newutm.txt

3	Properties of Aggregate      Since at least three-quarters of the volume of concrete is occupied by aggregate, it is  not surprising that its quality is of considerable importance. Not only may the  aggregate limit the strength of concrete, as weak aggregate cannot produce strong  concrete, but the properties of aggregate greatly affect the dura
bility and structural  performance of concrete.   Aggregate was originally viewed as an inert material dispersed throughout the cement  paste largely for economic reasons. It is possible, however, to take an opposite view  and to look on aggregate as a building material connected into a cohesive whole by means  of the cement paste, in a manner sim
ilar to masonry construction. In fact, aggregate is  not truly inert and its physical, thermal, and sometimes also chemical properties  influence the performance of concrete.   Aggregate is cheaper than cement and it is, therefore, economical to put into the mix  as much of the former and as little of the latter as possible. But economy is not the
	only reason for using aggregate: it confers considerable technical advantages on  concrete, which has a higher volume stability and better durability than the cement  paste alone.      General Classification of Aggregate    The size of aggregate used in concrete ranges from tens of millimetres down to  particles of the order of a tenth of a mill
imetre in cross-section. The maximum size  actually used varies but in any mix particles of different sizes are incorporated, the  particle size distribution being referred to as grading. In making low-grade concrete,  aggregate from deposits containing a whole range of sizes, from the largest to the  smallest, is sometimes used; this is referred 
to as all-in or pit-run aggregate. The  alternative, very much more common, and always used in the manufacture of good  concrete, is to obtain the aggregate in at least two size groups, the division being between  fine aggregate, often called sand, not larger 5 mm or 3/6 in., and coarse aggregate, which  comprises material at 5 mm or 3/16 in. in s
ize. In the United States, the division is made at  No. 4 ASTM sieve, which is 475 mm in size. More will be said about grading later, but this  basic division makes it possible to distinguish in the ensuing description between fine and  coarse aggregate. It should be noted that the use of the term aggregate (to mean coarse  aggregate) in contradis
tinction to sand is not correct, although comparatively common.   Sand is generally considered to have a lower size limit of about 0.07 mm or a little  less. Material between 0.06 mm and 0.002 mm is classified as silt, and particles smaller  still are termed clay. Loam is a soft deposit consisting of sand, silt, and clay in  about equal proportion
s.   All aggregate particles originally formed a part of a larger parent mass. This may have  been fragmented by natural processes of weathering and abrasion or artificially by  crushing. Thus many properties of the aggregate depend entirely on the properties of the  parent rock, e.g. chemical and mineral composition, petrographic description, spe
cific  gravity, hardness, strength, physical and chemical stability, pore structure, colour,  etc. On the other hand, there are some properties possessed by the aggregate but absent  in the parent rock: particle shape and size, surface texture, and absorption. All these  properties may have a considerable influence on the quality of the concrete e
ither fresh  or in the hardened state.   It is only reasonable to add, however, that, although these different properties of  aggregate per se can be examined, it is difficult to define a good aggregate other than  by saying that it is an aggregate from which good concrete (for the given conditions)  can be made. While aggregate whose properties a
ll appear satisfactory will always make  good concrete, the converse is not necessarily true and this is why the criterion of  performance in concrete has to be used. In particular, it has been found that aggregate  may appear to be unsatisfactory on some count but no trouble need be experienced when  it is used in concrete. For instance, a rock s
ample may disrupt on freezing but need not do  so when embedded in concrete, especially when the aggregate particles are well covered  by a paste of low permeability. However, aggregate considered poor in more than one  respect is unlikely to make a satisfactory concrete, so that tests on aggregate alone  are of help in assessing its suitability f
or use in concrete.      Classification of Natural Aggregates     So far, we have considered only aggregate formed from naturally occurring materials, and  the present chapter deals almost exclusively with this type of aggregate. Aggregate can,  however, also be manufactured from industrial  products: since these artificial aggregates  are general
ly either heavier or lighter than ordinary aggregate they are considered in  Chapter 9.  A further distinction can be made between aggregate reduced to its present size by natural  agents and crushed aggregate obtained by a deliberate fragmentation of rock.   From the petrological standpoint the aggregates, whether crushed or naturally reduced in 
 size, can be divided into several groups of rocks having common characteristics. The  classification of BS 812: Part 1: 1975 is most convenient and is given in Table 3.1. The  group classification does not imply suitability of any aggregate for concrete-making:  unsuitable material can be found in any group, although some groups tend to have a be
tter  record than others. It should also be remembered that many trade and customary names of  aggregates are in use, and these often do not correspond to the correct petrographic  classificatjjj.    Table 3.1: Classificatiii of Natural Aggregates According to Rock Type (BS 812: Part 1: 1975)     ASTM Standard C 294-69 (reapproved 1975) gives a de
scription of some of the more common  or important minerals found in aggregates. Mineralogical classification is of help in recognizing  properties of aggregate but cannot provide a basis for predicting its performance in concrete as there  are no minerals universally desirable and few invariably undesirable ones. The ASTM classification is  summa
rized below:   Silica minerals - (quartz, opal, chalcedony, tridymite, cristobalite)	Feldspars   Micaceous minerals	Carbonate minerals   Sulphate minerals   Iron sulphide minerals   Ferromagnesian minerals   Zeolites   Iron oxides   Clay minerals   The details of petrological and mineralogical methods are outside the scope of this  book, but i
t	is important to realize that geological examination of aggregate is a  useful aid in assessing its quality, and, in particular, in comparing a new aggregate  with one for which service records are available. Furthermore, adverse properties, such  as the presence of some unstable forms of silica, can be detected. In the case of  artificial aggreg
ates the influence of manufacturing methods and of processing can also  be studied.	Artificial Aggregates    Artificial aggregates are considered in Chapter 9 because they are lightweight or  high-density. One type will, however, be mentioned here because its use arises not from its  lightweight properties but because in the United Kingdom and 
in many other countries  there is a growing shortage of naturally occurring aggregates suitable for use in concrete.  The use of artificial aggregates is a natural step towards solving part of this problem,  and artificial aggregates manufactured from waste materials would appear to be an even  more sensible solution.   One such material, currentl
y being developed, is the ash produced from the  incinerators used to burn domestic refuse. The ash contains a proportion of both ferrous  and non-ferrous metals, both of which can be successfully removed and used again. The  remaining residue can be ground to a fine powder, blended with clay, pelletized and  fired in a kiln to produce an artifici
al aggregate.   Tests to date have shown the material capable of producing a concrete  with compressive strengths as high as 50 MPa (7000 psi) at 28 days. There will,  obviously, be problems with variations in the composition of the raw ash, and the long-term  durability characteristics of the material have yet to be determined, although results  
to date look promising. It is not envisaged that this material will be of use in high strength  structural concrete; however, it may be suitable for low strength concrete where currently  high-grade aggregates, of a quality far superior to that really required, are being used.  Previous use of this type of concrete is mentioned on Page 609.      S
ampling    Tests of various properties of aggregate are perforce performed on samples of the  material and, therefore, the results of the tests apply, strictly speaking, to the  aggregate in the sample only. Since, however, we are interested in the bulk of the  aggregate as supplied or as available for supply we should ensure that the sample is  t
ypical of the average properties of the aggregate. Such a sample is said to be representative,  and to obtain it certain precautions in procuring the sample have to be observed.   No detailed procedures can, however, be laid down because the conditions and situations  involved in taking samples in the field can vary widely from case to case. Never
theless, an  intelligent experimenter can obtain reliable results if he bears in mind at all times that the sample  taken is to be representative of the bulk of the material considered. An instance of such care  would be to use a scoop rather than a shovel so as to prevent rolling off of particles of some sizes  when the shovel is lifted. This bec
ame recognized in the 1967 revision of BS 812.   The main sample is made up of a number of portions drawn from different parts of the  whole. The minimum number of these portions, called increments, is ten, and they should  add up to a weight not less than that given in Table 3.2 for particles of different  sizes, as prescribed by BS 812 : Part 1:
1975. If, however, the source from which the  sample is being obtained is variable or segregated, a larger number of increments should  be taken and a larger sample ought to be dispatched for testing. This is particularly the case  in stockpiles when increments have to be taken from all parts of the pile, not only near its  surface but also from t
he centre.    Table 3.2: Minimum Weights of Samples for Testing (BS 812: Part 1: 1975)     It is clear from Table 3.2 that the main sample may be rather large, particularly when  large-size aggregate is used, and so the sample has to be reduced before testing. At all  stages of reduction it is necessary to ensure that the representative character 
of the  sample is retained so that the actual test sample has the same properties as the main  sample and ipso facto as the bulk of the aggregate.   There are two ways of reducing the size of a sample, each essentially dividing it into  two similar parts: quartering and riffling. For quartering, the main sample is  thoroughly mixed and in the case
 of fine aggregate dampened in order to avoid  segregation. The material is heaped into a cone and then turned over to form a new cone.  This is repeated twice, the material always being deposited at the apex of the cone so  that the fall of particles is evenly distributed round the circumference. The final cone  is flattened and divided into quar
ters. One pair of diagonally opposite quarters is  discarded, and the remainder forms the sample for testing or, if still too large, can be  reduced by further quartering. Care must be taken to include all fine material in the  appropriate quarter.   As an alternative, the sample can be split into halves using a riffler (Fig. 3.1). This is a box  
with a number of parallel vertical divisions, alternate ones discharging to the left and to the  right. The sample is discharged into the riffler over its full width, and the two halves are collected  into two boxes at the bottom of the chutes on each side. One half is discarded, and riffling of the  other half is repeated until the sample is redu
ced to the desired size. BS 812:1975 describes a  typical riffler.    Fig. 3.1. Riffler      Particle Shape and Texture    In addition to the petrological character of aggregate, its external characteristics are of importance,  in particular the particle shape and surface texture. The shape of three-dimensional bodies is rather  difficult to descr
ibe, and it is, therefore, convenient to define certain geometrical characteristics  of such bodies.   Roundness measures the relative sharpness or angularity of the edges and corners of a  particle. Roundness is controlled largely by the strength and abrasion resistance of the  parent rock and by the amount of wear to which the particle has been 
subjected. In the  case of crushed aggregate, the particle shape depends on the nature of the parent  material and on the type of crusher and its reduction ratio, i.e. the ratio of the size  of material fed into the crusher to the size of the finished product. A convenient broad  classification of roundness is that of BS 812 : Part 1 : 1975, given
 in Table 3.3.    Table 3.3: Particle Shape Classification of BS 812: Part 1: 1975 with Examples    A classification sometimes used in the United States is as follows -  Well-rounded - no original faces left  Rounded - faces almost gone  Subrounded - considerable wear, faces reduced in area  Subangular - some wear but faces untouched  Angular - li
ttle eviden                                                                                                                                                                                                                                                                                                                                                  
                                                                                                                                                                              ween any of the following -    (numbers).     The figure 67 in the expression for the angularity number represents the solid volume  of the most rounded gravel, so that the an
gularity number measures the percentage of  voids in excess of that in the rounded gravel (i.e. 33). The higher the number the more  angular the aggregate, the range for practical aggregate being between 0 and 11. A  development in measurement of angularity of aggregate, both coarse and fine but of  single size, is an angularity factor defined as 
the ratio of the solid volume of loose  aggregate to the solid volume of glass spheres of specified grading; thus no packing  is involved and the attendant error is avoided. The usefulness of the test is yet to be  determined.   The void content of aggregate can be calculated from the change in the volume of  air when a known decrease in pressure 
is applied; hence, the volume of air, i.e. the  volume of interstitial space, can be calculated.   An indirect proof of the dependence of the percentage of voids on the shape of  particles is obtained from Fig. 3.2, based on Shergold's data. The sample consisted  of a mixture of two aggregates, one angular, the other rounded, in varying proportion
s,  and it can be seen how increasing proportion of rounded particles decreases the  percentage of voids.    Fig. 3.2. Influence of angularity of aggregate on voids ratio ' (Crown copyright)     Another aspect of the shape of coarse aggregate is its sphericity, defined as a function  of the ratio of the surface area of the particle to its volume. 
Sphericity is related to the bedding  and cleavage of the parent rock, and is influenced by the type of crushing equipment when the  size of particles have been artificially reduced. Particles with a high ratio of surface area to  volume are of particular interest as they lower the workability of the mix. Elongated  and flaky particles are of this
 type. The latter can also affect adversely the  durability of concrete as they tend to be oriented in one plane, with water and air  voids forming underneath.   The presence of elongated or flaky particles in excess of 10 to 15 per cent of the  weight of coarse aggregate is generally considered undesirable, but  no recognized limits are laid down
.	  The weight of flaky particles expressed as a percentage of the weight of the sample is  called the flakiness index. Elongation index is similarly defined. Some particles are  both flaky and elongated, and are, therefore, countered in both categories.   The classification is made by means of simple gauges described in BS 812: Part 1:1975.  The 
division is based on the rather arbitrary assumption that a particle is flaky if its  thickness (least dimension) is less than 0.6 times the mean sieve size of the size  fraction to which the particle belongs. Similarly, a particle whose length (largest  dimension) is more than 1.8 times the mean sieve size of the size fraction is said to be  elon
gated. The mean size is defined as the arithmetic mean of the sieve size on which  the particle is just retained and the sieve size through which the particle just passes.  As closer size control is necessary, the sieves considered are not those of the standard  concrete aggregate series but: 75.0, 63.0, 50.0, 37.5, 28.0, 20.0, 14.0, 10.0, 6.30 an
d		5.00 mm (or about 3, 21/2, 2, 11/2, 1, 3/4, 1/2, 3/8, 1/4, and 3/16 in.) sieves. The flakiness and  elongation tests are useful for general assessment of aggregates but they do not  adequately describe the particle shape.   The classification of the surface texture is based on the degree to which the particle  surfaces are polished or dull, smo
oth or rough; the type of roughness has also to be  described. Surface texture depends on the hardness, grain size, and pore characteristics  of the parent material (hard, dense and fine-grained rocks generally having smooth  fracture surfaces) as well as on the degree to which forces acting on the particle  surface have smoothed or roughened it. 
Visual estimate of roughness is quite reliable,  but in order to reduce misunderstanding the classification of BS 812: Part 1 : 1975,  given in Table 3.4, should be followed. There is no recognized method of measuring the  surface roughness but Wright's approach is of interest : the interface between the particle  and a resin in which it is set is
 magnified, and the difference between the length of the profile  and the length of an unevenness line drawn as a series of chords is determined. This is  taken as a measure of roughness. Reproducible results are obtained, but method is  laborious and is not widely used.    Table 3.4: Surface Texture of Aggregates (BS 812: Part 1:1975) with Exampl
es     A recent attempt is the use of a shape coefficient and a surface texture coefficient  evaluated from a Fourier series method which a priori assumes ranges of the  harmonic system and also of a modified total roughness coefficient. It is doubtful whether  this type of approach is useful in evaluating and comparing the wide range of shapes an
d	 texture properties encountered in practice  Some other approaches are reviewed by Ozol.   It seems that the shape and surface texture of aggregate influence considerably the  strength of concrete. The flexural strength is more affected than the compressive  strength, and the effects of shape and texture are particularly significant in the case 
 of high strength concrete. Some data of Kaplan's are reproduced in Table 3.5 but this  gives no more than an indication of the type of influence, as some other factors may  not have been taken into account. The full role of shape and texture of aggregate in the  development of concrete strength is not known, but possibly a rougher texture results
	in a greater adhesive force between the particles and the cement matrix. Likewise, the  larger surface area of angular aggregate means that a larger adhesive force can be developed.    Table 3.5: Average Relative Importance of the Aggregate Properties Affecting the  Strength of Concrete     The shape and texture of fine aggregate have a signific
ant effect on the water  requirement of the mix made with the given aggregate. If these properties of fine  aggregate are expressed indirectly by its packing, i.e. by the percentage voids in a  loose condition (see p. 142), then the influence on the water requirement is quite  definite (see Fig. 3.3). The influence of the voids in coarse aggregate
 is less  definite.    Fig. 3.3. Relation between void content of sand in a loose condition and the water  requirement of concrete made with the given sand      Flakiness and the shape of coarse aggregate in general have an appreciable  effect on the workability of concrete. Fig. 3.4, reproduced from Kaplan's paper,  shows the pattern of the relat
ion between the angularity of coarse aggregate and the  compacting factor of concrete made with it. An increase in angularity from minimum to  maximum would reduce the compacting factor by about 0.09 but in practice there can  clearly be no unique relation between the two factors as other properties of aggregate  also affect the workability. Kapla
n's experimental results, however, do not confirm  that the surface texture is a factor.    Fig. 3.4. The relation between the angularity number of aggregate and the compacting  factor of concrete made with the given aggregate      Bond of Aggregate    Bond between aggregate and cement paste is an important factor in the strength of  concrete, esp
ecially the flexural strength, the full role of bond being only now  realized. Bond is due, in part, to the interlocking of the aggregate and the paste owing  to the roughness of the surface of the former. A rougher surface, such as that of  crushed particles, results in a better bond; better bond is also usually obtained with softer,  porous, and
 mineralogically heterogeneous particles. Generally, texture characteristics  which permit no penetration of the surface of the particles are not conducive to good bond.  In addition, bond is affected by other physical and chemical properties of aggregate, related  to its mineralogical and chemical composition, and to the electrostatic condition o
f	the particle  surface. For instance, some chemical bond may exist in the  limestone, dolomite, and  possibly siliceous aggregates, and at the surface of polished particles some capillary forces may  develop. However,  little is known about these phenomena, and relying on experience is still  necessary in predicting the bond between the aggregate
 and the surrounding cement paste   The determination of the quality of bond of aggregate is rather difficult and no accepted tests exist.  Generally, when bond is good, a crushed  concrete specimen should contain some aggregate particles  broken right  through, in addition to the more numerous ones pulled out from their  sockets. An excess  of fr
actured particles, however, might suggest that the aggregate is too weak. Because it depends on the  paste strength as well as on the properties of aggregate surface, bond strength increases with the  age of concrete; it seems that the ratio of bond strength to the strength of the paste increases with age  Thus, providing it is adequate, the bond 
strength per se may not be a controlling factor in the strength of  concrete. However, in high strength concrete there is probably a tendency for the bond strength  to be lower than the tensile strength of the cement paste so that preferential failure in bond takes place.  The problem of failure of concrete is discussed more fully in Chapter 5.   
	Strength of Aggregate    Clearly, the compressive strength of concrete cannot significantly exceed that of the  major part of the aggregate contained therein, although it is not easy to state what is  the strength of the individual particles. Indeed, it is difficult to test the crushing strength  of the aggregate by itself, and the required inf
ormation has to be obtained usually from  indirect tests: crushing strength of prepared rock samples, crushing value of bulk aggregate,  and performance of aggregate in concrete.   The latter simply means either previous experience with the given aggregate or a trial  use of the aggregate in a concrete mix known to have a certain strength with pre
viously  proven aggregates. If the aggregate under test leads to a lower compressive strength of  concrete, and in particular if numerous individual aggregate particles appear fractured  after the concrete specimen has been crushed, then the strength of the aggregate is  lower than the nominal compressive strength of the concrete mix in which the 
aggregate  was incorporated. Clearly, such aggregate can be used only in a concrete of lower  strength. This is, for instance, the case with laterite, a material widely spread in  Africa, South Asia and South America, which can rarely produce concrete stronger than  10 MPa (1500 psi).   Inadequate strength of aggregate represents a limiting case a
s the properties of  aggregate have some influence on the strength of concrete even when the aggregate by  itself is strong enough not to fracture prematurely. If we compare concretes made with  different aggregates we can observe that the influence of aggregate on the strength of  concrete is qualitatively the same whatever the mix proportions, a
nd is the same  regardless of whether the concrete is tested in compression or in tension. It is  possible that the influence of aggregate on the strength of concrete is due not only to  the mechanical strength of the aggregate but also, to a considerable degree, to its  absorption and bond characteristics.   In general, the strength and elasticit
y of aggregate depend on its composition, texture  and structure. Thus a low strength may be due to the weakness of constituent grains or  the grains may be strong but not well knit or cemented together.   The modulus of elasticity of aggregate is rarely determined; this is, however, not  unimportant as the modulus of elasticity of concrete is gen
erally higher the higher the  modulus of the constituent aggregate, but depends on other factors as well. The modulus  of elasticity of aggregate affects also the magnitude of creep and shrinkage that can be  realized by the concrete.   A good average value of the crushing strength of aggregate is about 200 MPa (30 000 psi)  but many excellent agg
regates range in strength down to 80 MPa (12 000 psi). One of the  highest values recorded is 530 MPa (77 000 psi) for a certain quartzite. Values for other rocks  are in Table 3.6. It should be noted that the required strength of concrete is considerably  higher than the normal range of concrete because the actual stresses at the points of contac
t of  individual particles within the concrete may be far in excess of the nominal compressive  stress applied.    Table 3.6: Compressive Strength of American Rocks Commonly Used as Concrete Aggregates     On the other hand, aggregate of moderate or low strength and modulus elasticity can be  valuable in preserving the durability of concrete. Volu
me changes of concrete, arising from  hygral or thermal reasons, lead to a lower stress in the cement paste when the aggregate is  compressible. Thus compressibility of aggregate would reduce distress in concrete while a  strong and rigid aggregate might lead to cracking of the surrounding cement paste.  It may be noted that no general relation ex
ists between the strength and modulus of  elasticity of different aggregates. Some granites, for instance, have been found to  have a modulus of elasticity of 45 GPa (6.5 x 10'6 psi), and gabbro and diabase a modulus  of 85.5 GPa (12.4 x 10'6 psi), the strength of all these rocks ranging between 145 and  170 MPa (21000 to 25 000 psi). Values of th
e modulus in excess of 160 GPa (23 x 10'6 psi)  have been encountered.   A test to measure the compressive strength of prepared rock cylinders was prescribed by  BS 812 : 1967. In this test, a 25.4 mm (1 in.) diameter cylinder 25.4 mm (1 in.) high, is  used, and the nominal crushing strength of an oven-dry specimen is determined to the  nearest 0.
5 MPa (or 100 psi). The preparation of the sample involves drilling, sawing and  grinding - all rather laborious operations The results of the crushing test are affected  by the presence of planes of weakness in the rock, and there is, therefore, some doubt  about the value of this test, particularly as the structural weakness in the rock may  not
 be significant once the rock has been comminuted to the size used in concrete. In  essence, the crushing strength test measures the quality of the parent rock rather than the  quality of the aggregate as used in concrete. For this reason, in 1975, the test was deleted from  BS 812, and tests on prepared specimens are nowadays less used than tests
 on bulk aggregate,  but are nevertheless useful when dealing with a potential new source of crushed aggregate.   Sometimes, the strength of a wet as well as of a dry specimen is determined. The ratio of wet  to dry strengths measures the softening effect, and when this is high, poor durability of the rock  may be suspected.   A test on the crushi
ng properties of bulk aggregate is the so-called crushing value  test of BS 812: Part 3: 1975. There is no explicit relation between this crushing value  and the compressive strength, but the results of the two tests are in agreement (Fig. 3.5).  The crushing value is a useful guide when dealing with aggregates of unknown  performance, particularl
y when lower strength may be suspected, as for instance with  limestone and some granites and basalts.    Fig. 3.5. Relation between the compressive strength of the parent rock and the  crushing value of aggregate obtained from the same rock     The material to be tested should pass a 14.0 mm (1/2 in.) test sieve and be retained on a  10.0 mm (3/8
 in.) sieve. When, however, this size is not available, particles of other sizes  may be used, but those larger than standard will in general give a higher crushing  value, and the smaller ones a lower value than would be obtained with the same rock of  standard size. The sample to be tested should be dried in an oven at 100 to 110 C (212  to 230 
F) for four hours, and then placed in a cylindrical mould and tamped in a  prescribed manner. A plunger is put on top of the aggregate and the whole assembly is  placed in a compression testing machine and subjected to a load of 400 kN (40 tons)  (pressure of 22.1 MPa (3200 psi)) over the gross area of the plunger, the load being  increased gradua
lly over a period of 10 minutes. After the load has been released, the  aggregate is removed and sieved on a 2.36 mm (No. 8 ASTM) test sieve in the case of a  sample of the 14.0 to 10.0 mm (1/2 to 3/8 in.) standard size; for samples of other sizes,  the sieve size is prescribed in BS 812 : Part 3 : 1975. The ratio of the weight of the material  pa
ssing this sieve to the total weight of the sample is called the aggregate crushing value.   In the United States, where large quantities of artificial lightweight aggregates are used,  attempts were made to develop a strength test for these aggregates, rather similar to the  crushing value test described above, but no test has been standardized. 
	The crushing value test is rather insensitive to the variation in strength of  weaker aggregates, i.e. those with a crushing value of over 25 to 30. This is so  because, having been crushed before the full load of 400 kN (40 tons) has been applied,  these weaker materials become compacted so that the amount of crushing during later  stages of th
e	test is reduced. For this reason, a ten per cent fines va                                                                                                                                                                                                                                                                                                  
	                                                                                                                                                                                                                            r honeycombed aggregate (such as expanded shale or foamed slag).     These penetrations should result in a percentage of fines 
passing a 2.36 mm (No. 8  ASTM) sieve of between 7.5 and 12.5 per cent. If y is the actual percentage of fines due  to a maximum load of x tons, then the load required to give 10 per cent fines is given  by    (formula).    It should be noted that in this test, unlike the standard crushing value test, a higher numerical  result denotes a higher st
rength of the aggregate. BS 882:1973 prescribes, minimum value of  100 kN (10 tons) for aggregate to be used in concrete wearing surfaces, and 50 kN (5 tons)  when used in other concretes.   The ten per cent fines value test shows a fairly good correlation with the standard crushing  value test for strong aggregates, while for weaker aggregates th
e	ten per cent fines value test  is more sensitive and gives a truer picture of differences between more or less weak samples.  For this reason, the test is of use in assessing light weight aggregates but there is no simple  relation between the test result and the upper limit of strength of concrete made with  the given aggregate.      Other Mech
anical Properties of Aggregate    Several mechanical properties of aggregate are of interest, especially when the  aggregate is to be used in road construction or is to be subjected to high  wear.   The first of these is toughness, which can be defined as the resistance of  aggregate to failure by impact. Toughness can be determined on prepared cy
lindrical  samples of rock: the minimum height from which a standard weight has to be dropped so  as to cause failure of the specimen represents the toughness of the material. This test was  devised in the days of horse-drawn and steel-tyred traffic, and was abandoned as an ASTM  test in 1965, although it would disclose adverse effects of weatheri
ng of the rock under test.   It is possible also to determine the impact value of bulk aggregate, and  toughness determined in this manner is related to the crushing value, and can, in fact,  be used as an alternative test. The size of the particles tested is the same as in the  crushing value test and the permissible values of the crushed fractio
n	smaller than a  2.36 mm (No. 8 ASTM) test sieve are also the same. The impact is applied by a standard  hammer falling 15 times under its own weight upon the aggregate in a cylindrical  container. This results in fragmentation in a manner similar to that produced by the  pressure of the plunger in the aggregate crushing value test. Full details 
of the test  are prescribed in BS 812 : Part 3: 1975, and BS 882 : 1973 prescribes the following  maximum values of the average of duplicate samples -  30 per cent when the aggregate is to be used in concrete for wearing  surfaces, and  45 per cent when to be used in other concretes.  These figures serve as useful guides, but it is clear that a di
rect correlation  between the crushing value and the performance of aggregate in concrete or the strength  of the concrete is not possible.   One advantage of the impact test is that it can be performed in the field with some  modifications, such as the measurement of quantities by volume rather than by weight,  but the test may not be adequate fo
r	compliance purposes.   In addition to strength and toughness, hardness or resistance to wear is an important  property of concrete used in roads and in floor surfaces subjected to heavy traffic.  Several tests are available, and it is possible to cause wear by abrasion, i.e. by rubbing  of a foreign material against the stone under test, or by a
ttrition of stone particles against  one another.   In the abrasion (Derry) test, a cylindrical specimen, similar to those used in the  crushing strength test, is subjected to wear by quartz (Leighton Buzzard) sand pressed  against the cylinder by a rotating metal disc. The abrasion value is expressed as 20  minus one-third of the loss of weight o
f	 the cylinder in grams. Good stone has an abrasion  value of not less than 17; stone with a value of less than 14 would be considered poor.   The details of the abrasion test were included in BS 812 up to the 1951  edition. Nowadays the test has gone out of favour both in Britain and in the United  States, and, in keeping with the tendency to te
st aggregate in bulk,  a new abrasion value test was introduced in BS 812 : 1967. According to BS 812 : Part 3  : 1975, aggregate particles between 14.0 and 10.2 mm are made up in a tray in a single  layer, using a setting compound. The sample is subjected to abrasion in a standard  machine, the grinding lap being turned 500 revolutions with Leigh
ton Buzzard sand being  fed continuously at a prescribed rate. The aggregate abrasion value is defined as the  percentage loss in weight on abrasion, so that a high value denotes low resistance to  abrasion.    The attrition (Deval) test also uses aggregate in bulk. Particles of known total  weight are subjected to wear in an iron cylinder rotated
 10 000 times  at 30 to 33  revolutions per minute. The proportion of broken material  expressed as a percentage  represents the attrition value. The test can be  performed on dry or wet aggregate, and  the difference in results indicates the influence of the condition of the aggregate on its  resistance to attrition. An attrition value of about 7
 to 8 is usually considered as the  maximum permissible, but a shortcoming of the test is that it gives only small numerical  differences between widely differing aggregates. The test was overed by ASTM  Standard D2-33 (reapproved in 1968) but was discontinued in 1971.    An American test combining attrition and abrasion is the Los Angeles  test; 
it is quite frequently used in other countries, too, because its results show  good correlation not only with the actual wear of aggregate when used in concrete but  also with the compressive and flexural strengths of  concrete made with the  given aggregate. In this test, aggregate of specified grading is placed in a  cylindrical drum, mounted ho
rizontally, with a shelf  inside. A charge of steel balls  is added, and the drum is rotated a specified number of revolutions. The tumbling and  dropping of the aggregate and the balls results in abrasion and attrition of the aggregate, and  this is measured in the same way as in the attrition test.   The Los Angeles test can be performed on aggr
egates of different sizes, the same wear  being obtained by an appropriate weight of the sample and of the charge of steel balls,  and by a suitable number of revolutions. The various quantities are prescribed by ASTM  Standard C 131-76. The Los Angeles test is, however, not very suitable for the  assessment of the behaviour of fine aggregate when
 subjected to attrition on prolonged  mixing. This problem has only recently been identified; limestone sand is probably one  of the more common materials to undergo this degradation. For this reason, unknown fine  aggregates should, in addition to standard tests, be subjected to a wet attrition test  to see how much material smaller than 75 um (N
o. 200 sieve) is produced. No standard  apparatus is available but some development has been made by Meininger.   Table 3.7 gives average values of crushing strength, aggregate crushing value,  abrasion, impact, and attrition for the different rock groups of BS 812 : Part 2 : 1975.  It should be noted that the values for hornfels and schists are b
ased on a few specimens  only; these groups would appear to be better than they really are, presumably because  only good quality hornfels and schists were tested. As a rule, they are not suitable for  use in concrete. Likewise, chalk is not included in the limestone group data as it is  not generally suitable as a concrete aggregate.    Table 3.7
: Average Test Values for British Rocks of Different Groups     As far as the crushing strength is concerned, basalt is extremely variable, fresh  basalts with little olivine reaching some 400 MPa (60 000 psi), while decomposed basalt  at the other end of the scale may have a strength of no more than 100 MPa (15 000 psi).  Limestone and porphyry s
how much less variation in strength, and in Britain porphyry has  a good general performance - rather better than that of granites, which tend to be  variable.   An indication of the accuracy of the results of the different tests is given by Table 3.8,  listing the number of samples to be tested in order to ensure a 0.9 probability  that the mean 
value for the samples is within +/-3 and also within +/-10 per cent of the  true mean. The aggregate crushing value shows up as particularly consistent. On the  other hand, the prepared specimens show a greater scatter of results than the bulk  samples, which is of course to be expected. While the various tests described in this  and succeeding se
ctions give an indication of the quality of the aggregate, it is not  possible to predict from the properties of aggregate the potential strength development  of concrete made with the given aggregate, and indeed it is not yet possible to  translate physical properties of aggregate into its concrete-making properties.    Table 3.8: Reproducibility
 of Test Results on Aggregate      Specific Gravity    Since aggregate generally contains pores both permeable and impermeable (see p 143),  the meaning of the term specific gravity has to be carefully defined, and there are indeed  several types of specific gravity.   The absolute specific gravity refers to the volume of the solid material exclud
ing all pores, and can,  therefore, be defined as the ratio of the weight of the solid, referred to vacuum, to the  weight of an  equal volume of gas-free distilled water, both taken at a stated temperature. Thus, in order to eliminate  the effect of totally enclosed impermeable pores the material has to be pulverized, and the test is both  labori
ous and sensitive. Fortunately, it is not normally required in concrete technology work.   If the volume of the solid is deemed to include the impermeable pores, but not the  capillary ones, the resulting specific gravity is prefixed by the word apparent. The apparent specific  gravity is then the ratio of the weight of the aggregate dried in an o
ven at 100 to 110 C (212 to 230 F)  for 24 hours to the weight of water occupying a volume equal to that of the solid including  the impermeable pores. The latter weight is determined using a vessel which can be accurately  filled with water to a specified volume. Thus, if the weight of the oven-dried sample is D, the  weight of the vessel full of
 water is B, and the weight of the vessel with the sample and topped up  with water is A, then the weight of the water occupying the same volume as the solid is B - (A - D).  The apparent specific gravity is then     (formula).    The vessel referred to earlier, and known as a pycnometer, is usually a one-litre jar with a  watertight metal conical
 screwtop having a small hole at the apex. The pycnometer can  thus be filled with water so as to contain precisely the same volume every time.   Calculations with reference to concrete are generally based on the  saturated surface-dry condition of the aggregate (see p. 146) because the water  contained in all the pores in the aggregate does not t
ake part in the chemical reactions  of cement and can, therefore, be considered as part of the aggregate. Thus, if a sample  of the saturated and surface-dry aggregate weighs C, the gross apparent specific gravity is    (formula).     This is the specific gravity most frequently and easily determined and necessary for  calculations of yield of con
crete or of the quantity of aggregate required for a given  volume of concrete.   The apparent specific gravity of aggregate depends on the specific gravity of the  minerals of which the aggregate is composed and also on the amount of voids. The  majority of natural aggregates have a specific gravity of between 2.6 and 2.7, and the  range of value
s is given in Table 3.9. The values for artificial aggregates extend from  considerably below to very much above this range (see Chapter 9).    Table 3.9: Apparent Specific Gravities of Different Rock Groups     As mentioned earlier, specific gravity of aggregate is used in the calculation of  quantities but the actual value of the specific gravit
y of aggregate is not a measure  of its quality. Thus the value of specific gravity should not be specified unless we  are dealing with a material of a given petrological character when a variation in  specific gravity would reflect the porosity of the particles. An exception to this is  the case of mass construction, such as a gravity dam, where 
a minimum density  of concrete is essential for the stability of the structure.       Bulk Density     It is well known that in the metric system the density of a material is  numerically  equal to its specific gravity although, of course, the latter is a  ratio while density  is expressed in kilogrammes per litre. However, in  concrete practice, 
expressing the  density in kilogrammes per cubic metre is  more common. In the Imperial system, specific  gravity has to be multiplied  by the unit weight of water (approximately 624 lb / ft3) in  order to be converted into absolute density (specific weight) expressed in pounds per  cubic foot.   This absolute density, it must be remembered, refer
s to the volume of the individual  particles only, and of course it is not physically possible to pack these particles so  that there are no voids between them. When aggregate is to be actually batched by volume  it is necessary to know the weight of aggregate that would fill a container of unit  volume. This is known as the bulk density of aggreg
ate, and this density is used to  convert quantities by weight to quantities by volume.   The bulk density clearly depends on how densely the aggregate is packed, and it follows  that for a material of a given specific gravity the bulk density depends on the size  distribution and shape of the particles: particles all of one size can be packed to 
a  limited extent but smaller particles can be added in the voids between the larger ones,  thus increasing the bulk density of the packed material. The shape of the particles  greatly affects the closeness of packing that can be achieved.   For a coarse aggregate of given specific gravity, a higher bulk density means that  there are fewer voids t
o be filled by sand and cement, and the bulk density test has at  one time been used as a basis of proportioning of mixes.   The actual bulk density of aggregate depends not only on the various characteristics of  the material which determine the potential degree of packing, but also on the actual  compaction achieved in a given case. For instance
, using spherical particles all of the  same size, the densest packing is achieved when their centres lie at the apexes of  imaginary tetrahedra. The bulk density is then 0.74 of the specific weight of the  material. For the loosest packing, the centres of spheres are at the corners of  imaginary cubes and the bulk density is only 0.52 of the spec
ific weight of the solid.   Thus, for test purposes, the degree of compaction has to be specified. BS 812 : Part 2  : 1975 recognizes two degrees: loose (or uncompacted) and compacted. The test is  performed in a metal cylinder of prescribed diameter and depth, depending on the maximum  size of the aggregate and also on whether compacted or uncomp
acted bulk density is  being determined.   For the determination of loose bulk density, the dried aggregate is gently  placed in the container to overflowing and then levelled by rolling a rod across the  top. In order to find the compacted or rodded bulk density, the  container is filled in three stages, each third of the volume being tamped a  p
rescribed number of times with a 16 mm (5/8 in.) diameter round-nosed rod. Again, the  overflow is removed. The net weight of the aggregate in the container divided by its  volume then represents the bulk density for either degree of compaction. The ratio of  the loose bulk density to the compacted bulk density lies usually between 0.87 and 0.96. 
 Knowing the apparent specific gravity for the saturated and surface-dry condition, Q,  the voids ratio can be calculated from the expression    (formula).    If the aggregate contains surface water it will pack less densely owing to the bulking  effect. This is discussed on p. 148. Moreover, the bulk density as determined in the  laboratory may n
ot be directly suitable for conversion of weight to volume of aggregate  for purposes of volume batching as the degree of compaction in the laboratory and on the  site may not be the same.   The bulk density of aggregate is of interest in connection with the use of  lightweight and heavy aggregates (see Chapter 9).      Porosity and Absorption of 
Aggregate    The presence of internal pores in the aggregate particles was mentioned in connection  with the specific gravity of aggregate, and indeed the characteristics of these pores  are very important in the study of its properties. The porosity of  aggregate, its permeability,  and absorption, influence such properties of aggregate as the bo
nd between it and the  cement paste, the resistance of concrete to freezing and thawing, as well as its chemical  stability and resistance to abrasion. As stated earlier, the apparent specific gravity  of aggregate also depends on its porosity and, as a consequence, the yield of concrete  for a given weight of aggregate is affected.   The pores in
 aggregate vary in size over a wide range, the largest being large enough to be  seen under a microscope or even with the naked eye, but even the smallest aggregate pores  are generally larger than the gel pores in the cement paste. Pores smaller than 4 um are of  special interest as they are generally believed to affect the durability of aggregat
es subjected  to alternating freezing and thawing (see p. 464).   Some of the aggregate pores are wholly within the solid; others open onto the surface of the  particle. The cement paste, because of its viscosity, cannot penetrate to a great depth any but  the largest of the aggregate pores, so that it is the gross volume of the particle that is c
onsidered  solid for the purpose of calculating the aggregate content in concrete. However, water can  enter the pores, the amount and rate of penetration depending on their size, continuity  and total volume. The order of porosity of some common rocks is given in Table 3.10, and  since aggregate represents some three-quarters of the volume of con
crete it is clear that the  porosity of aggregate materially contributes to the overall porosity of concrete.    Table 3.10: Porosity of some Common Rocks     When all the pores in the aggregate are full it is said to be saturated and surface-dry.  If aggregate in this condition is allowed to stand free in dry air, e.g. in the  laboratory, some of
 the water contained in the pores will evaporate and the aggregate  will be less than saturated, i.e. air-dry. Prolonged drying in an oven would reduce the  moisture content of the aggregate still further until, when no moisture whatever is  left, the aggregate is said to be bone-dry. These various stages are shown  diagrammatically in Fig. 3.6, a
nd some typical values of absorption are given in Table  3.11. At the extreme right of Fig. 3.6, the aggregate contains surface moisture and is  darker in colour.    Table 3.11: Typical Values of Absorption of Different Aggregates    Fig. 3.6. Diagrammatic representation of moisture in aggregate     The water absorption of aggregate is determined 
by measuring the increase in weight of  an oven-dried sample when immersed in water for 24 hours (the surface water being  removed). The ratio of the increase in weight to the weight of the dry sample, expressed  as a percentage, is termed absorption. Standard procedures are prescribed in BS 812 :  Part 2 : 1975.   Some typical values of absorptio
n of different aggregates are given in Table 3.11,  based on Newman's data. The moisture content in the air-dry condition is also  tabulated. It may be noted that gravel has generally a higher absorption than crushed  rock of the same petrological character since weathering results in the outer layer of  the gravel particles being more porous and 
absorbent.   Although there is no clear-cut relation between the strength of concrete and the water  absorption of aggregate used, the pores at the surface of the particle affect the bond  between the aggregate and the cement paste, and may thus exert some influence on the  strength of concrete.   Normally, it is assumed that at the time of settin
g of concrete the aggregate is in a  saturated and surface-dry condition. If the aggregate is batched in a dry condition it  is assumed that sufficient water will be absorbed from the mix to bring the aggregate to  a saturated condition, and this absorbed water is not included in the net or effective  mixing water. It is possible, however, that wh
en dry aggregate is used the particles  become quickly coated with cement paste which prevents further ingress of water  necessary for saturation. This is particularly so with coarse aggregate, where water has  further to travel from the surface of the particle. As a result, the effective  water / cement ratio is higher than would be the case had 
full absorption of water by the  aggregate been possible. This effect is significant mainly in rich mixes where rapid  coating of aggregate can take place; in lean, wet mixes the saturation of aggregate  proceeds undisturbed. In practical cases the actual behaviour of the mix is affected  also by the order of feeding the ingredients into the mixer
.   The absorption of water by aggregate results also in some loss of workability with  time, but beyond about 15 minutes the loss becomes small .   Since absorption of water by dry aggregate slows down or is stopped owing to the  coating of particles with cement paste, it is often useful to determine the quantity of  water absorbed in 10 to 30 mi
nutes instead of the total water absorption, which may  never be achieved in practice.      Moisture Content of Aggregate    It was mentioned in connection with the specific gravity that in fresh concrete the  volume occupied by the aggregate is the volume of the particles including all the pores.  If no water movement into the aggregate is to tak
e place the pores must be full of  water, i.e. the aggregate must be in a saturated condition. On the other hand, any water  on the surface of the aggregate will contribute to the water in the mix and will occupy  a volume in excess of that of the aggregate particles. The basic state of the aggregate  is then saturated and surface-dry.   Aggregate
 exposed to rain collects a considerable amount of moisture on the surface of  the particles, and, except at the surface of the stockpile, keeps this moisture over  long periods. This is particularly true of fine aggregate, and the surface or free  moisture (in excess of that held by aggregate in a saturated and surface-dry condition)  must be all
owed for in the calculation of batch quantities. The surface moisture is  expressed as a percentage of the weight of the saturated and surface-dry aggregate, and  is termed the moisture content.   Since absorption represents the water contained in aggregate in a saturated and  surface-dry condition, and the moisture content is the water in excess 
of that state,  the total water content of a moist aggregate is equal to the sum of absorption and  moisture content.   As the moisture content of aggregate changes with weather and changes also from one  part of a stockpile to another, the value of the moisture content has to be determined  frequently and a number of methods have been developed. 
The oldest one consists simply  of finding the loss in weight of an aggregate sample when dried on a tray over a source  of heat. Care is necessary to avoid over-drying: the sand should be brought to a just  free-flowing condition, and must not be heated further. This stage can be determined by  feel or by forming the sand into a pile by means of 
a conical mould; when the mould has  been removed the material should slump freely. If the sand has acquired a brownish  tinge, this is a sure sign that excessive drying has taken place. This method of  determining the moisture content of aggregate, colloquially referred to as the "frying-pan  method" is simple, can be used in the field and is qui
te reliable.   In the laboratory the moisture content of aggregate can be determined by means of a  pycnometer. The apparent specific gravity of the aggregate on a saturated and surface-dry  basis, Q, must be known. Then, if B is the weight of the pycnometer full of water, C  the weight of the moist sample and A the weight of the pycnometer with t
he sample and  topped up with water, the moisture content of the aggregate is    (formula).     The test is slow and requires great care in execution (e.g. all air must be expelled  from the sample) but can yield accurate results.   In the siphon can test the volume of water displaced by a known weight of moist  aggregate is measured, the siphon m
aking this determination more accurate. Preliminary  calibration for each aggregate is required as the results depend on its specific gravity  but, once this has been done, the test is rapid and accurate.   The moisture content of aggregate can also be found using a steelyard moisture meter:  the moist aggregate is added to a vessel containing a f
ixed amount of water and  suspended at one end of a steelyard until it balances. We measure thus the quantity of  water that has to be replaced by the moist aggregate for a constant weight and total  volume. For this condition it can be shown that the amount of displaced water is  proportional to the moisture content of the aggregate. A calibratio
n curve for any  aggregate used has to be obtained. The moisture content can be determined with an  accuracy of 1/2 per cent.   In the buoyancy meter test the moisture content of the aggregate of known specific  gravity is determined from the apparent loss in weight on immersion in water. The  balance can read the moisture content directly if the 
size of the sample is adjusted,  according to the specific gravity of the aggregate, to such a value that a saturated and  surface-dry sample has a standard weight when immersed. The test is rapid and gives the  moisture content to the nearest 1/2 per cent. A simple version of the test is prescribed  by ASTM Standard C 70-79.   Numerous other meth
ods have been developed. For instance, moisture can be removed by  burning the aggregate with methyl alcohol, the resulting loss in weight of the sample being  measured. There are also proprietary meters based on the measurement of pressure of gas  formed in a closed vessel by the reaction of calcium carbide with the moisture in the  sample.   Ele
ctrical devices which give instantaneous or continuous reading of the  moisture content of aggregate in a storage bin, on the basis of the variation of  resistance or capacitance with a change in the moisture content of the aggregate, have  been developed. In some batching plants, meters of this type are used in automatic  devices which regulate t
he quantity of water to be added to the mixer but an accuracy  greater than 1 per cent of moisture cannot in practice be achieved. Microwave meters  have recently been developed but these are very expensive.   It can be seen that a great variety of tests is available but, however accurate the  test, its result is significant only if a representati
ve sample has been used.  Furthermore, if the moisture content of aggregate varies between adjacent parts of a  stockpile, the adjustment of mix proportions becomes laborious. Since the variation in  moisture content occurs mainly in the vertical direction from a water-logged bottom of a  pile to its drying or dry surface, care in laying out of st
ockpiles is necessary: storing in  horizontal layers, having at least two stockpiles and allowing each pile to  drain before  use, and not using the bottom 300 mm (12 in.) or so, all help to keep the variation in  moisture content to a minimum. Coarse aggregate holds very much less water than sand,  has a less variable moisture content, and genera
lly causes fewer difficulties.      Bulking of Sand    The presence of moisture in aggregate necessitates correction of the actual mix  proportions: the weight of water added to the mix has to be decreased by the weight of  the free moisture in the aggregate, and the weight of the aggregate must be increased  by a like amount. In the case of sand,
 there is a second effect of the presence of  moisture: bulking. This is the increase in the volume of a given weight of sand caused  by the films of water pushing the sand particles apart. While bulking per se does not  affect the proportioning of materials by weight, in the case of volume batching, bulking  results in a smaller weight of sand oc
cupying the fixed volume of the measuring box. For  this reason, the mix becomes deficient in sand and appears "stony", and the concrete may  be prone to segregation and honeycombing. Also, the yield of concrete is reduced.  The remedy, of course, lies in increasing the apparent volume of sand to allow for bulking.   The extent of bulking depends 
on the percentage of moisture present in the sand and on  its fineness. The increase in volume relative to that occupied by a saturated and  surface-dry sand increases with an increase in the moisture content of the sand up to a  value of some 5 to 8 per cent, when bulking of 20 to 30 per cent occurs. Upon further  addition of water, the films mer
ge and the water moves into the voids between the  particles so that the total volume of sand decreases until, when fully saturated  (flooded), its volume is approximately the same as the volume of dry sand for the same  method of filling the container. This is apparent from Fig. 3.7, which also shows that finer  sand bulks considerably more and r
eaches maximum bulking at a higher water content  than does coarse sand. Extremely fine sand has been known to bulk as much as 40 per  cent at a moisture content of 10 per cent, but such a sand is in any case unsuitable for the  manufacture of good quality concrete.    Fig. 3.7. Decrease in true volume of sand due to bulking (for a constant volume
 of moist sand)     Coarse aggregate shows only a negligible increase in volume due to the presence of free  water, as the thickness of moisture films is very small compared with the particle size.   Since the volume of saturated sand is the same as that of dry sand, the most convenient  way of determining bulking is by measuring the decrease in v
olume of the given sand when  inundated. A container of known volume is filled with loosely packed moist sand. The  sand is then tipped out, the container is partially filled with water and the sand is  gradually fed back, with stirring and rodding to expel all air bubbles. The volume of  sand in the saturated state, Vs is now measured. If Vm, is 
the initial volume of the  sand (i.e. the volume of the container), then bulking is given by -    (formula).    With volume batching, bulking has to be allowed for by increasing the total volume of  (moist) sand used. Thus volume Vs is multiplied by a factor -    (formula),    sometimes known as the bulking factor, and a graph of bulking factor ag
ainst moisture of  three typical sands is shown in Fig. 3.8.    Fig. 3.8. Bulking factor for sands with different moisture contents     The bulking factor can also be found from the bulk density of dry and moist sand, Dd and  Dm respectively, and the moisture content per unit volume of sand, m / Vm. The bulking  factor is then -    (formula).     
Since Dd represents a ratio of the weight of dry sand, w , to its bulk volume Vs (the  volume of dry and inundated sand being the same) -    (formula),    i.e. the two factors are identical.      Deleterious Substances in Aggregate    There are three broad categories of deleterious substances that may be found in  aggregates: impurities which inte
rfere with the process of hydration of cement; coatings  preventing the development of good bond between aggregate and the cement paste; and  certain individual particles which are weak or unsound in themselves. All or part of an  aggregate can also be harmful through the development of chemical reactions between the  aggregate and the cement past
e: these chemical reactions are discussed on page 158.    Organic Impurities    Natural aggregates may be sufficiently strong and resistant to wear and yet they may not  be satisfactory for concrete-making if they contain organic impurities which interfere with  the chemical reactions of hydration.  The organic matter found in aggregate consists  
usually of products of decay of vegetable matter (mainly tannic acid and its derivative) and  appears in the form of humus or organic loam. Such materials are more likely to be  present in sand than in coarse aggregate, which is easily washed.   Not all organic matter is harmful and it is best to check its effects by making actual test  cubes. Gen
erally, however, it saves time to ascertain first whether the amount of organic  matter is sufficient to warrant further tests. This is done by the so-called colorimetric test  of ASTM Standard C 40-73. The acids in the sample are neutralised by a 3 % solution of  NaOH, prescribed quantities of aggregate and of solution being placed in a bottle. T
he  mixture is vigorously shaken to allow the intimate contact necessary for chemical action,  and then left to stand for 24 hours, when the organic content can be judged by the colour  of the solution: the greater the organic content the darker the colour. If the colour of the  liquid above the test sample is not darker then the standard yellow c
olour defined by the  standard, the sample can be assumed to contain only a harmless amount of organic  impurities.   If the observed colour is darker than the standard, i.e. if the solution appears brownish or  brown, the aggregate has a rather high organic content, but this does not necessarily mean  that the aggregate is not fit for use in conc
rete. The organic matter present may not be  harmful to concrete or the colour may be due to some iron-bearing minerals. For this  reason, further tests are necessary: concrete cubes are made using the suspected  aggregates and their strength is compared with concrete of the same mix proportion but  made with aggregates of known quality.   In earl
ier edition, BS 812 contained the colorimetric tests, but in 1967 there was  introduced a measurement of the pH values of cement mortars under standard conditions.  The test is rather laborious and is suitable for laboratory work only, and was deleted from  BS 812 in 1975. In consequence, at present the calorimetric test is the best means for a  p
reliminary assessment of suitability of aggregates: If no colour change in excess of that  specified is observed, no further tests are necessary unless contamination with industrial  affluence has occurred.   In some countries the quantity of organic matter in aggregate is determined from the loss  of weight of a sample on treating with hydrogen p
eroxide.   It is interesting to note that in some cases the effects of organic impurities may be only  temporary. In one investigation concrete made with a sand containing organic matter had  a 24-hour strength equal to 53 per cent of the strength of similar concrete made with clean sand.  At 3 days this ratio rose to 82 per cent, then to 92 per c
ent at 7 days, and at 28 days equal  strengths were recorded.    Clay and Other Fine Material    Clay may be present in aggregate in the form of surface coatings which interfere with the  bond between aggregate and the cement paste. Since good bond is essential to ensure a  satisfactory strength and durability of concrete the problem of clay coati
ng is an important  one.   There are two more types of fine material which can be present in aggregate: silt and  crusher dust. Silt is a material between 2 um and 60 um, reduced to this size by natural  processes of weathering; silt may thus be found in aggregate won from natural deposits.  On the other hand crusher dust is a fine material formed
 during the process of  comminution of rock into crushed stone or, less frequently of gravel into crushed  sand. In  a properly laid out processing plant this dust should be removed by washing. Other soft  or loosely adhering coatings can be removed during the processing of  the aggregate.  Well-bonded coatings cannot be so removed, but if  they a
re chemically stable and have no  deleterious effect there is no objection to the use of aggregate with such a coating  although shrinkage may be increased. However, aggregates with chemically reactive  coatings, even if physically stable, can lead to serious trouble.   Silt and fine dust may form coatings similar to those of  clay,  or may be pre
sent in the  form of loose particles not bonded to the course aggregate. Even when they are in the  latter form, silt and fine dust should not be present in excessive quantities because, owing  to their fineness and therefore large surface area, silt and fine dust increase the amount of  water necessary to wet all the particles in the mix.   In vi
ew of the above, it is necessary to control the clay, silt and fine dust contents of  aggregate. BS 882 : 1973 limits the content of  all three materials together to not more  than -    15 per cent by weight in crushed stone sand,  3 per cent by weight in natural or crushed gravel sand, and  1 per cent by weight in coarse aggregate.     ASTM stand
ard C 33-78 lays down similar requirements, but distinguishes between  concrete subject to abrasion and other concretes. In the former case, the amount material  passing a 75 um (No. 200) test sieve is limited to 3 per cent of the weight of sand, instead  of  the 5 per cent value permitted for other concretes. The corresponding for coarse  aggrega
te is laid down as 1 per cent.   In the same standard, the content of clay and friable particles is specified separately as 3  per cent in fine and 2 to 10 per cent in coarse aggregate, depending on the use of the  concrete.   It may be noted that different test methods are prescribed in different specifications so that  the results are not direct
ly comparable.   The clay, silt and fine dust content of fine aggregate can be determined by the  sedimentation method described in BS 812 : Part 1 : 1975. The sand sample is placed in a  sodium oxalate solution in a stoppered jar and rotated with the axis of the jar horizontal  for 15 minutes at approximately 80 revolutions per minute. The fine s
olids become  dispersed and the amount of suspended material is then measured by means of an  Andreason pipette. A simple calculation gives the percentage of clay, fine silt and fine dust  in the sand, the separation size being 20 um.   A similar method, with suitable modifications can be used for coarse aggregate, but it is  simpler to wet-sieve 
the aggregate on a 75 um (No. 200) test sieve, as prescribed in BS  812 : Part 1 : 1975 and ASTM Standard C 117 - 76. This type of sieving is resorted to  because fine dust and clay adhering to larger particles would not be separated in ordinary  dry sieving. In wet sieving, on the other hand, the aggregate is placed in water and  agitated suffici
ently vigorously for the finer material to be brought into suspension. By  decantation and sieving, all material smaller than a 75 um (No. 200) test sieve can be  removed. To protect this sieve from damage by large particles during decantation a  1.18 mm (No. 16 ASTM) sieve is placed above the 75 um (No. 200) sieve.   For natural sands and crushed
 gravel sands, there is also a field test available which can be  performed quite easily and rapidly, with very little laboratory equipment. 50 ml of  an  approximately 1 per cent solution of common salt in water is place in a 250 ml BS  measuring cylinder. Sand, as received , is added until the total volume of the mixture in the  cylinder is 150 
ml. The cylinder is now covered with the palm of the hand, shaken  vigorously, repeatedly turned upside down and then allowed to stand for 3 hours. The silt  which became dispersed on shaking will now settle in a layer above the sand, and the  height of this layer can be expressed as a percentage of the height of the sand below.   It should be rem
embered that this a volumetric ratio, which cannot easily be converted to  a ratio by weight since the conversion factor depends on the fineness of the material. It has  been suggested that for natural sand the weight ratio is obtained by multiplying the  volumetric ratio by a factor of 1/4, the corresponding figure for crushed gravel sand being  
1/2, but with some aggregate an even wider variation is obtained. These conversions are  not reliable, and BS 882 : 1973 recommends that the volumetric content exceeds 8 per  cent tests by the more accurate methods, described earlier, should be made.    Salt Contamination    Sand won from the seashore or from a river estuary contains salt, and has
 to be processed;  more than 10 per cent aggregate used in Britain is of marine origin. The simplest course is  to wash the sand in fresh water, but special care is required with deposits just above the  high-water mark in which large quantities of salt, sometimes over 6 per cent of weight of  sand, may be found. Generally, sand from the sea bed, 
washed even in sea water, does not  contain harmful quantities of salts.	The British Code of Practice for the Structural Use of Concrete CP 110 : 1972 specifies  the maximum total chloride and all other sources (see p. 106). It is fair to add that it is not  the total chloride content in aggregate that matters in practice, but rather the soluble
	chloride, or even the degree to which the soluble chloride reacts during the hydration of  cement. In consequent, if the situation is such that the use of an aggregate with an  excessive chloride content is economically desirable, tests should be made to determine its  suitability; of course, compliance with CP 110 : 1972 cannot be claimed in su
ch a case.   If salt is not removed it will absorb moisture from the air and cause efflorescence - unsightly  white deposits on the surface of the concrete (see also p. 455).  A slight corrosion of  reinforcement may also result, but this is not believed to progress to a dangerous degree,  especially when the concrete is of good quality and adequa
te cover to reinforcement is  provided. No trouble need be expected in mass concrete structures. Special problems  arising from the presence of various salts in aggregates found in arid region, such as the  Middle East are treated by Fookes and Collis.   Sea sands are often extremely fine and the  grading of any new sand should be carefully  check
ed.  Sea-dredged coarse aggregate may have a large shell content.  This has no  adverse effect on strength but workability in concrete made with aggregate having a large  shell content is slightly reduced. The shell content of particles larger than 5 mm can be  determined by hand picking, using the method of an amendment of the British Standard  B
S	812 : Part 2: 1975.    Unsound Particles    Test on aggregate sometimes reveal that the majority of the component particles are  satisfactory but that a few are unsound: the quantity of such particles must clearly be  limited.   There are two broad types of unsound particles: those that fail to maintain their integrity,  and those that lead to d
isruptive expansion on freezing or even exposure to water.  The  disruptive properties are characteristics of certain rock groups, and will therefore be  discussed in relation to the durability of aggregate in general (mainly in the next section).  In this section, non-durable impurities only will be considered.   Shale and other particles of low 
density are regarded as unsound and so are soft inclusions  such as clay lumps, wood, and coal, as they lead to pitting and scaling.  If present in large  quantity (over 2 to 5 per cent of the weight of the aggregate) these particles may adversely  affect the strength of concrete and should certainly not be permitted in concrete which is  exposed 
to abrasion.   Coal, in addition to being a soft inclusion, is undesirable for other reasons: it can swell,  causing disruption of concrete and, if present in large quantities in a finely divided form, it  can disturb the process of hardening of the cement paste. However, discrete particles of  hard coal amounting to no more a 1/4 per cent of the 
weight of the aggregate have no  adverse effect on the strength of concrete.   The presence of coal and other materials of low density can be determined by flotation in a  liquid of suitable specific gravity, as, for instance, by the method of ASTM Standard C  123 - 69 (reapproved 1975). If the danger of pitting and scaling is not thought importan
t,  and strength of concrete is the main consideration, a trial mix should be made.   Mica should be avoided because in the presence of active chemical agents produced  during the hydration of cement, alteration of mica to other forms may result. Also, free  mica in fine aggregate, even in quantities of a few per cent of the weight of the aggregat
e  affects adversely the water requirement and hence the strength of concrete. It appears that  mica in the form of muscovite is much more harmful than biotite. These facts should be  borne in mind when materials such as china clay sand are considered for use in concrete.   Gypsum and other sulphates must not be present; their existence in many Mi
ddle East  aggregates leads to difficulties, but up to 5 per cent of  SO3 by weight of cement  (including that in the cement) is often tolerated  there.   Iron pyrites and marcasite represent the most common expansive  inclusions in  aggregate. These sulphides react with water and oxygen in the air to form a ferrous  sulphate which subsequently de
composes to form the  hydroxide, while the sulphate ions  react with calcium aluminates in the cement. Surface staining of the concrete and  disruption of the cement paste  (pop-outs) may result, particularly under warm and humid  conditions.   Not all pyrites are reactive but, since the decomposition of pyrites takes place only in lime  water, it
 is possible to test a suspect aggregate for reactivity by placing the material in a  saturated solution of lime. If the aggregate is reactive a blue-green gelatinous  precipitate of ferrous sulphate appears within a few minutes, and on exposure to air this  changes to brown ferric hydroxide. The absence of this reaction means that no staining  ne
ed be feared. Lack of reactivity was found by Midgley to be associated with the  presence of a number of metal cations, while their absence makes the pyrites active.  Generally, particles of pyrites likely to cause trouble are those between 5 and 10 mm (or  3/16 and 3/8 in.) in size.   The permissible quantities of unsound particles laid down by A
STM Standard C 33-78  are summarized in Table 3.12.    Table 3.12 Permissible Quantities of Unsound Particles Prescribed by ASTM Standard C 33-78     The majority of impurities discussed in the present section are found in  natural aggregate deposits and are much less frequently encountered in crushed aggregate.  However, some processed aggregates
, such as mine tailings, can contain harmful  substances. For instance, small quantities of lead soluble in limewater (e.g. 0.1 per cent of  PbO by weight of aggregate) greatly delay the set and reduce the early strength of  concrete; the long-term strength is unaffected.      Soundness of Aggregate    This is the name given to the ability of aggr
egate to resist excessive changes in volume as  a result of changes in physical conditions. Lack of soundness is thus distinct from  expansion caused by the chemical reactions between the aggregate and the alkalis in  cement.   The physical causes of large or permanent volume changes of aggregate are freezing and  thawing, thermal changes at tempe
ratures above freezing, and alternating wetting and  drying.   Aggregate is said to be unsound when volume changes, induced by the above causes,  result in deterioration of the concrete. This may range from local scaling and so-called  pop-outs to extensive surface cracking and to disintegration over a considerable depth,  and can thus vary from n
o more than impaired appearance to a structurally dangerous  situation.   Unsoundness is exhibited by porous flints and cherts, especially the lightweight ones with  a fine-textured pore structure, by some shales, by limestones with laminae of expansive  clay, and by other particles containing clay minerals, particularly of the montmorillonite or 
 illite group. For instance, an altered dolerite has been found to move as much as 0.0006  with wetting and drying, and concrete containing this aggregate might fail under  conditions of alternating wetting and drying, and will certainly do so on freezing and  thawing.   A test for soundness of aggregate is prescribed by the ASTM Standard C 88-76.
 A sample  of graded aggregate is subjected alternately to immersion in a saturated solution of sodium  or magnesium sulphate (generally the more severe of the two) and drying in an oven. The  formation of salt crystals in the pores of the aggregate tends to disrupt the particles,  probably in a manner similar to the action of ice. The reduction i
n size of the particles, as  shown by a sieve analysis, after a number of cycles of exposure denotes the degree of  unsoundness. The test is no more than qualitative in predicting the behaviour of the  aggregate under actual site conditions, and cannot be used as a basis of acceptance or  rejection of unknown aggregates. Specifically, there is no 
clear reason why soundness as  tested by ASTM Standard C 88-76 should be related to performance in concrete subjected  to freezing and thawing.   Other tests consist of subjecting the aggregate to cycles of alternating freezing and  thawing, and sometimes this treatment is applied to mortar or concrete made with the  suspect aggregate. Unfortunate
ly, none of the tests gives an accurate indication of the  behaviour of aggregate under actual conditions of moisture and temperature changes  above the freezing point.   Likewise, there are no tests which could satisfactorily predict the durability of aggregate  in the concrete under conditions of freezing and thawing. The main reason for this is
 that  the behaviour of aggregate is related to the presence of the surrounding cement paste, so  that only a service record can satisfactorily prove the durability of aggregate.   Nevertheless, certain aggregates are known to be susceptible to frost damage and it is on  these that our attention is centred. These are: porous cherts. shales. limest
ones,  particularly laminated limestones, and some sandstones. A common characteristic of these  rocks with a poor record is their high absorption, but it should be emphasized that many  durable rocks also exhibit high absorption (see Fig. 3.9).    Fig. 3.9. Distribution of sound and unsound aggregate samples as a function of absorption     For fr
ost damage to occur there must exist critical conditions of water content and lack of  drainage. These are governed, inter alia, by the size, shape and continuity of pores in the  aggregate, because these characteristics of the pores control the rate and amount of absorption  and the rate at which water can escape from the aggregate particle. Inde
ed, these features of the  pores are more important than merely their total volume as reflected by the magnitude of  absorption.   It has been found that pores smaller than 4 to 5 um are critical, for they are large  enough to permit water to enter but not large enough to allow easy drainage under the  pressure of ice. This pressure, in fully conf
ined space at -20 C (-4 F), may be as high as  200 MPa (29 000 psi). Thus, if splitting of aggregate particles and disruption of the  surrounding cement paste are to be avoided, flow of water towards unfilled pores within  the aggregate particle or into the surrounding paste must be possible before the hydraulic  pressure becomes high enough to ca
use disruption.   This argument illustrates the statement made earlier that the durability of aggregate  cannot be fully determined other than when it is embedded in cement paste: the particle  may be strong enough to resist the pressure of ice but expansion may cause disruption of  the surrounding mortar.   It has been said that the pore size is 
an important factor in the durability of aggregate. In  most aggregates, pores of different sizes are present so that we are really confronted with  a pore size distribution. A means of expressing this quantitatively has been developed by  Brunauer, Emmett and Teller. The specific surface of the aggregate is determined from  the amount of a gas so
rbate required to form a layer one molecule thick over the entire  internal surface of the aggregate pores. The total volume of the pores is measured by  absorption, and the ratio of the volume of pores to their surface represents the hydraulic  radius of the pores. This value, familiar from flow problems in hydraulics, gives an  indication of the
 pressure required to produce flow.      Alkali - aggregate Reaction    During the last forty years, some deleterious chemical reactions between the aggregate  and the surrounding cement paste have been observed. The most common reaction is that  between the active silica constituents of the aggregate and the alkalis in cement. The  reactive forms
 of silica are opal (amorphous), chalcedony (cryptocrystalline fibrous), and  tridymite (crystalline). These reactive materials occur in: opaline or chalcedonic cherts,  siliceous limestones, rhyolites and rhyolitic tuffs, dacite and dacite tuffs, andesite and  andesite tuffs, and phyllites.   The reaction starts with the attack on the siliceous m
inerals in the aggregate by the  alkaline hydroxides derived from the alkalis (Na2O and K2O) in the cement. As a result, an  alkali-silicate gel is formed, and alteration of the borders of the aggregate takes place. The  gel is of the "unlimited swelling" type: it imbibes water with a consequent tendency to increase  in volume. Since the gel is co
nfined by the surrounding cement paste, internal pressures result and  eventually lead to expansion, cracking and disruption of the cement paste (pop-outs).  Thus expansion appears to be due to hydraulic pressure generated through osmosis, but  expansion can also be caused by the swelling pressure of the still solid products of the  alkali - silic
a reaction. For this reason, it is believed that it is the swelling of the hard  aggregate particles that is most harmful to concrete. Some of the relatively soft gel is later  leached out by water and deposited in the cracks already formed by the swelling of the  aggregate. The size of the siliceous particles controls the speed with which reactio
n  occurs, fine particles (20 to 30 um) leading to expansion within a month or two, larger  ones only after some years.   Detailed studies of the alkali - aggregate reaction have been reported by Diamond. He  views the reaction as being primarily due to the high concentration of hydroxide ions in  pore solutions but with the alkali cations being o
f critical importance because their  concentration influences the reaction rates and the physical characteristics of the products  of reaction. The necessity of the presence of Ca(OH)2 has been suggested.   While we can predict that with given materials an alkali - aggregate reaction will take  place, it is not generally possible to estimate the d
eleterious effects from the knowledge of  the quantities of the reactive materials alone. For instance, the actual reactivity of  aggregate is affected by its particle size and porosity as these influence the area over which  the reaction can take place. Since the quantity of alkalis depends on the cement only, their  concentration at the reactive
 surface of aggregate will be governed by the magnitude of  this surface. The minimum alkali content of cement at which expansive reaction may take  place is 0.6 per cent of the soda equivalent. This is calculated from stoichiometry as the  actual Na2O content plus 0.658 times the K2O content of the clinker. In exceptional cases,  however, cements
 with an even lower alkali content have been known to cause  expansion. Within limits, the expansion of concrete made with a given reactive  aggregate is greater the higher the alkali content of the cement and, for a given  composition of cement, the greater its fineness.   Other factors influencing the progress of the alkali - aggregate reaction 
include the  availability of non-evaporable water in the paste and the permeability of the paste.  Moisture is necessary and the reaction is accelerated under conditions of alternating  wetting and drying. Higher temperature accelerates the reaction, at least in the range 10 to  38 C (50 to 100 F). It can thus be seen that various physical and che
mical factors make  the problem of alkali - aggregate reaction highly complex. In particular, the gel can change  its constitution by absorption and thus exert a considerable pressure, while at other times  diffusion of the gel out of the confined area takes place. It may be noted that, as the  hydration of cement progresses, much of the alkali is
 concentrated in the aqueous phase.  As a consequence, pH rises and all silica minerals become soluble.   It is not surprising, therefore, that, although we know that certain types of aggregate tend  to be reactive, there is no simple way of determining whether a given aggregate will cause  excessive expansion due to reaction with alkalis in the c
ement. Service record has  generally to be relied upon but as little as 0.5 per cent of defective aggregate can cause  damage. If no record is available it is possible only to determine the potential  reactivity of the aggregate but not to prove that reaction will take place. A quick chemical  test is prescribed by ASTM Standard C 289-71 (reapprov
ed 1976): the reduction in the  alkalinity of a normal solution of NaOH when placed in contact with pulverized aggregate  at 80 C (176 F) is determined, and the amount of dissolved silica is measured. The  interpretation of the result is in many cases not clear, but generally a potentially  deleterious reaction is indicated if the plotted test res
ult falls to the right of the boundary  line of Fig. 3.10, reproduced from the ASTM Standard but based on Mielenz and Witte's  paper. However, potentially deleterious aggregates represented by points lying above  the dashed line in Fig. 3.10 may be extremely reactive with alkalis so that a relatively low  expansion may result. These aggregates sho
uld therefore be tested further to determine  whether their reactivity is deleterious by the mortar bar test described below. The test is of  little value with lightweight aggregates.    Fig. 3.10. Results of chemical test of ASTM Standard C 289-71 (reapproved 1976)     In the mortar bar test for the physical reactivity of aggregate, the suspected
 aggregate,  crushed if need be and made up to a prescribed grading, is used in making special sand-cement  mortar bars, using a cement with an equivalent alkali content of not less than 0.6  per cent. The bars are stored over water at 38 C (100 F), at which temperature the  expansion is more rapid and usually higher than at higher or lower temper
atures. The  reaction is also accelerated by the use of a fairly high water / cement ratio. The details of  procedure are prescribed by ASTM Standard C 227-71 (reapproved 1976); a modification  has been suggested by Brotschi and Mehta. The aggregate under test is considered  harmful if it expands more than 0.05 per cent after 3 months or more than
 0.1 per cent  after 6 months.   This test has shown a very good correlation with field experience, but a considerable time  is required before judgement on the aggregate can be pronounced. On the other hand, as  mentioned earlier, the results of the chemical test, which is rapid, are often not conclusive.  Likewise, petrographic examination, alth
ough a useful tool in identifying the mineral  constituents, cannot establish that a given mineral will result in abnormal expansion. A  rapid and conclusive test for aggregate reactivity is thus still to be developed, and to use  more than one of the existing tests is the best that can be done at the moment.   It has been found that expansion due
 to the alkali - aggregate reaction can be reduced or  eliminated by the addition to the mix of reactive silica in a finely powdered form. This  apparent paradox can be explained by reference to Fig. 3.11, showing the relation between  the expansion of a mortar bar and the content of reactive silica of size between 850 and  300 um (No. 20 and No. 
50 ASTM) sieves, i.e. not in a powdered form. In the range of  low silica contents, the greater quantity of silica for a given amount of alkalis increases  expansion, but with higher values of silica content the situation is reversed: the greater the  surface area of the reactive aggregate the lower the quantity of alkalis available per unit of  t
his area, and the less alkali-silica gel can be formed. On the other hand, owing to the  extremely low mobility of calcium hydroxide, only that adjacent to  the surface of the aggregate is available for reaction, so that the quantity of calcium  hydroxide per unit area of aggregate is independent of the magnitude of the total surface  area of the 
aggregate. Thus, increasing the surface area increases the calcium hydroxide / alkali  ratio of the solution at the boundary of the aggregate. Under such circumstances an innocuous  (non-expanding) calcium alkali silicate product is formed.    Fig. 3.11. Relation between expansion after 224 days and reactive silica content in the aggregate     By 
a similar argument, finely divided siliceous material added to the coarse reactive  particles already present would reduce expansion, although the reaction with alkalis still  takes place. These pozzolanic additions, such as crushed pyrex glass or fly ash, have  indeed been found effective in reducing the penetration of the coarser aggregate parti
cles.  There are indications that fly ash is not quite as good as other pozzolanas. The  performance of any pozzolana in this respect should be tested following the ASTM  Standard C 441-69 (reapproved 1975). With a sufficient amount of added silica, the initial  reaction lowers the alkali concentration to that obtained with a low alkali cement whe
n no  additive is present. It is essential, however, that a sufficient amount of pulverized silica be  added; it is generally recommended that 20 g of reactive silica be added for each gram of  alkali in excess of 0.5 per cent of the weight of the cement. Thus the amount of  pozzolana is quite large. Inadequate amounts can actually aggravate the s
ituation and  increase expansion if a particularly bad silica - alkali ratio is reached (cf. Fig. 3.11). A side  effect of the addition of pozzolanas which has to be borne in mind is the increase in the  water requirement of the mix.   The alkali - aggregate reaction of the type described is widespread in many countries,  notably North America, Sc
andinavia, India, Australia and New Zealand. The reaction was  first encountered in Great Britain in 1978 in several structures ten to thirty years old.    Alkali - carbonate Reaction    Another type of deleterious aggregate reaction is that between some dolomitic limestone  aggregates and alkalis in the cement. Expansion of concrete, similar to t
hat occurring as a  result of the alkali - silica reaction, takes place under humid conditions. Typically, reaction  zones up to 2 mm (or 0.1 in.) are formed around the active aggregate particles. Cracking  develops within these rims and leads to a network of cracks and a loss of bond between  the aggregate and the cement paste.   Tests have shown
 that de-dolomitization occurs, but the reactions involved are still  imperfectly understood; in particular, the role of clay in the aggregate is not clear but  expansive reaction seems to be nearly always associated with the presence of clay. Also, in  expansive aggregates the dolomite and calcite crystals are very fine. One suggestion is  that t
he expansion is due to moisture uptake by the previously unwetted clay, the  de-dolomitization being necessary only to provide access of moisture to the locked-in  clay; another is that clay increases the reactivity of the aggregate so that dolomite and  calcium silicate hydrate produce Mg(OH)2, silica gel, and calcium carbonate with a  volume inc
rease of about 4 per cent. A good review of current thinking is given by  Walker.	It should be stressed that only some dolomitic limestones cause expansive reaction in  concrete. No simple test to identify them has been developed; in case of doubt, help can be  obtained from an investigation of the rock texture, rock expansion in sodium hydroxid
e  (ASTM Standard C 586-69 (reapproved 1975)) or when used in test beams made with  cements having a high alkali content. ASTM is preparing a new standard for a test for  length change of concrete due to alkali - carbonate rock reaction.   One distinction between the silica - and carbonate - alkali reactions which should be borne  in mind is that 
in the latter the alkali is regenerated. It is probably for this reason that  pozzolanas are not effective in controlling the alkali - carbonate expansion. Fortunately,  reactive carbonate rocks are not very widespread and can usually be avoided.      Thermal Properties of Aggregate    There are three thermal properties of aggregate that may be si
gnificant in the performance  of concrete: coefficient of thermal expansion, specific heat, and conductivity. The last two  are of importance in mass concrete or where insulation is required, but not in ordinary  structural work, and are discussed in the section dealing with the thermal properties of  concrete (see p. 487).   The coefficient of th
ermal expansion of aggregate influences the value of this coefficient  for concrete containing the given aggregate: the higher the coefficient of the aggregate the  higher the coefficient of the concrete, but the latter depends also on the aggregate content  in the mix and on the mix proportions in general.   There is, however, another aspect of t
he problem. It has been suggested that if the  coefficients of thermal expansion of the coarse aggregate and of the cement paste differ  too much, a large change in temperature may introduce differential movement and a break  in the bond between the aggregate  particles and the surrounding paste. However, possibly  because the differential movemen
t is affected also by other forces, such as those due to  shrinkage, a large difference between the coefficients is not necessarily detrimental when  the temperature does not vary outside the range of, say, 4 to 60 C (40 to 140 F).  Nevertheless, when the two coefficients differ by more than 5.5 x 10'-6 per  C (3 x 10'-6  per  F) the durability of
 concrete subjected to freezing and thawing may be affected.  The coefficient of thermal expansion can be determined by means of a dilatometer devised  by Verbeck and Hass for use with both fine and  coarse aggregate. The linear coefficient  of thermal expansion varies with the type of parent rock, the range for the more common  rocks being about 
0.9 x 10'-6 to 16 x 10'-6 per  C (0.5 x 10'-6 to 8.9 x 10'-6 per  F), but the  majority of aggregates lie between approximately 5 x 10'-6 to 13 x 10'-6 per  C (3 x 10'-6 and  7 x 10'-6 per  F) (see Table 3.13). For hydrated Portland cement paste, the coefficient varies  between 11x10'-6 and 16 x 10'-6 per  C (6 x 10'-6 and 9 x 10'-6 per  F), but v
alues up to  20.7 x 10'-6 per  C (11.5 x 10'-6 per  F) have also been reported, the coefficient varying with  the degree of saturation. Thus, a serious difference in coefficients occurs only with aggregates  of a very low expansion; these are certain granites, limestones and marbles.    Table 3.13: Linear Coefficient of Thermal Expansion of Differ
ent Rock Types	  If extreme temperatures are expected, the detailed properties of any given aggregate have  to be known. For instance, quartz undergoes inversion at 574 C and expands suddenly by  0.85 per cent. This would disrupt the concrete, and for this reason fire-resistant concrete is  never made with quartz aggregate.      Sieve Analysis  
	This somewhat grandiose name is given to the simple operation of dividing a sample of  aggregate into fractions, each consisting of particles of the same size. In practice, each  fraction contains particles between specific limits, these being the openings of standard  test sieves.   The test sieves used for concrete aggregate have square openin
gs and their properties are  prescribed by BS 410 : 1976. Sieves used to be described by the size of the opening (in  inches) for larger sizes, and by the number of openings per lineal inch for sieves smaller  than about 1/8 in. Thus a No. 100 test sieve has 100 x 100 openings in each square inch,  the size of the opening and the width of the wire
 of which the mesh is made being laid  down in previous editions of BS 410; a summary is given in an appendix to the 1976  edition. Nowadays, sieve sizes are designated by the nominal aperature size in millimetres  or micrometres.   Sieves smaller than 4 mm (0.16 in.) are normally made of wire cloth although, if required,  this can be used up to 1
6	mm (0.62 in.). The wire cloth is made of phosphor bronze but  for some coarser sieves brass and mild steel can also be used. The screening area, i.e. the  area of the openings as a percentage of the gross area of the sieve, varies between 34 and  53 per cent.   Coarse test sieves (4 mm (0.16 in.) and larger) are made of perforated mild steel pla
te,  with a screening area of 44 to 65 per cent.   All sieves are mounted in frames which can nest. It is thus possible to place the sieves one  above the other in order of size with the largest sieve at the top, and the material retained  on each sieve after shaking represents the fraction of aggregate coarser than the sieve in  question but fine
r	than the sieve above. 200 mm (8 in.) diameter frames are used for 5 mm  (3/16 in.) or smaller sizes, and 300 or 400 mm (12 or 18 in.) diameter frames for 5 mm  (3/16 in.) and larger sizes. It may be remembered that 5 mm (or 3/16 in., No. 4 ASTM) is  the dividing line between the fine and coarse aggregate.   The sieves used for concrete aggregate
 consist of a series in which the clear opening of  any sieve is approximately one-half of the opening of the next larger sieve size. The BS  test sieve sizes in Imperial units for this series were as follows: 3 in., 1 1/2 in., 3/4 in., 3/8  in., 3/16 in., Nos. 7, 14, 25, 52, 100, and 200.   For determination of oversize and undersize aggregate, a
nd especially for research work  on aggregate grading, additional sieve sizes are required. The full sequence of test sieves is  based theoretically on the ratio of 4/2 for the openings of two consecutive sieves.  However, recently both the British (BS 410:1976) and American (ASTM Ell-70  (reapproved 1977)) sieves have been standardized generally 
in accordance with the R40/3  sieve series of the International Standards Organization. Not all of these sizes form a true  geometric series but follow "preferred numbers".   Table 3.14 gives the standard sieve sizes according to their fundamental description by  aperture in millimetres or micrometres and also the previous British and ASTM  design
ations and approximate apertures in inches.    Table 3.14: Standard American and British Sieve Sizes     The international sieve sizes for aggregate are given by the standard ISO / DIS 6274: 1980;  there is a lack of unanimity evident here as three series are recommended (Table 3.15) of  which one, series C, does not overlap with the sieve sizes o
f	Table 3.14.	Table 3.15: International Sieve Sizes of ISO Standard 6274     In the United Kingdom, the sieve sizes used, according to British Standard 812: Part 1: 1975,  are as follows: 75.0, 63.0, 50.0, 37.5, 28.0, 20.0, 14.0, 10.0, 6.30, 5.00, 3.35, 2.36, 1.70, 1.18 mm  and 850, 600, 425, 300, 212, 150 and 75 um. Of these, four sizes (50.0,
 28.0, 14.0 and	6.30 mm)  do not fit into any of the sieve sizes of Table 3.14 or Table 3.15; they are part of a yet different  ISO series of sieve sizes. For grading purposes, the sieve sizes normally used are: 75.0, 50.0,  37.5, 20.0, 10.0, 5.00, 2.36, 1.18 mm, and 600, 300, and 150 um.   We can thus see that in discussing aggregate grading we h
ave to contend with two sets of  sieve sizes. In this book, results of measurements made with Imperial size sieves will be  reported by the exact metric equivalent, but grading curves for design purposes (see  Chapter 10) will, wherever available, be based on the new BS metric sieve sizes.   Before the sieve analysis is performed, the aggregate sa
mple has to be air-dried in order  to avoid lumps of fine particles being classified as large particles and also to prevent  clogging of the finer sieves. The weights of the reduced samples for sieving, as  recommended by BS 812: Part 1: 1975, are given in Table 3.16 and Table 3.17 shows the  maximum weight of material with which each sieve can co
pe. If this weight is exceeded on  a sieve, material which is really finer than this sieve may be included in the portion  retained. The material on the sieve in question should, therefore, be split into two parts  and each should be sieved separately.    Table 3.16: Minimum Weight of Sample for Sieve Analysis According to BS 812: Part 1: 1975    
Table 3.17: Maximum Weight to be Retained at the Completion of Sieving According to  BA 812: Part 1: 1975	The actual sieving operation can be performed by hand, each sieve in turn being shaken  until not more than a trace continues to pass. The movement should be backwards and  forwards, sideways left and right, circular clockwise and anticloc
kwise, all these motions  following one another so that every particle "has a chance" of passing through the sieve.  In most modern laboratories a sieve shaker is available, usually fitted with a time switch so  that uniformity of the sieving operation can be ensured. None the less, care is necessary in  order to make sure that no sieve is overloa
ded. (See Table	3.17.)   The results of a sieve analysis are best reported in tabular form, as shown in Table 3.18.  Column (2) shows the weight retained on each sieve. This is expressed as a percentage of  the total weight of the sample and is shown in column (3). Now, working from the finest  size upwards the cumulative percentage (to the neares
t one per cent) passing each sieve  can be calculated [column (4)1, and it is this percentage that is used in the plotting of  grading curves.    Table 3.18: Example of Sieve Analysis    Grading Curves    The results of a sieve analysis can be grasped much more easily if represented graphically,  and for this reason grading charts are very extensi
vely used. By using a chart it is possible  to see at a glance whether the grading of a given sample conforms to that specified, or is  too coarse or too fine, or deficient in a particular size.   In the grading chart commonly used, the ordinates represent the cumulative percentage  passing and the abscissae the sieve opening plotted to a logarith
mic scale. Since the  openings of sieves in a standard series are in the ratio of 1/2, a logarithmic plot shows these  openings at a constant spacing. This is illustrated in Fig. 3.12 which represents the data of  Table 3.18.    Fig. 3.12. Example of a grading curve (see table 3.18)     It is convenient to choose a scale such that the scale spacin
g between two adjacent sieve  sizes is approximately equal to 20 per cent on the ordinate scale; a visual comparison of  different grading curves can then be made from memory.    Fineness Modulus    A single factor computed from the sieve analysis is sometimes used, particularly in the  United States. This is the fineness modulus, defined as the s
um of the cumulative  percentages retained on the sieves of the standard series: 150, 300, 600 um, 1.18, 2.36,  5.00 mm (ASTM Nos. 100, 50, 30, 16, 8, 4) and up to the largest sieve size present. It  should be remembered that, when all the particles in a sample are coarser than, say, 600  um (No. 30 ASTM), the cumulative percentage retained on 300
 um (No. 50 ASTM)  should be entered as 100; the same value, of course, would be entered for 150 um (No.  100). The value of the fineness modulus is higher the coarser the aggregate (see column  (5), Table 3.18).   The fineness modulus can be looked upon as a weighted average size of a sieve on which  the material is retained, the sieves being cou
nted from the finest. For instance, a fineness  modulus of 4.00 can be interpreted to mean that the fourth sieve, 1.18 mm (No. 16  ASTM) is the average size. However, it is clear that one parameter, the average, cannot be  representative of a distribution: thus the same fineness modulus can represent an infinite  number of totally different size d
istributions or grading curves. The fineness modulus  cannot, therefore, be used as a description of the grading of an aggregate, but it is valuable  for measuring slight variations in the aggregate from the same source, i.e. as a day-to-day  check. Nevertheless, within certain limitations, the fineness modulus gives an indication of  the probable
 behaviour of a concrete mix made with aggregate having a certain grading,  and the use of the fineness modulus in assessment of aggregates and in mix design has  many supporters.      Grading Requirements    We have seen how to find the grading of a sample of aggregate but it still remains to  determine whether or not a particular grading is suit
able. A related problem is that of  combining fine and coarse aggregates so as to produce a desired grading. What, then, are  the properties of a "good" grading curve?   Since the strength of fully compacted concrete with a given water / cement ratio is  independent of the grading of the aggregate, grading is, in the first instance, of importance 
 only in so far as it affects workability. As, however, the development of strength  corresponding to a given water / cement ratio requires full compaction, and this can be  achieved only with a sufficiently workable mix, it is necessary to produce a mix that can be  compacted to a maximum density with a reasonable amount of work.   It should be s
tated at the outset that there is no one ideal grading curve but a compromise  is aimed at. Apart from the physical requirements, the economic problem must not be  forgotten: concrete has to be made of materials which can be produced cheaply so that no  narrow limits can be imposed on aggregate.   It has been suggested that the main factors govern
ing the desired aggregate grading are:  the surface area of the aggregate, which determines the amount of water necessary to wet  all the solids; the relative volume occupied by the aggregate; the workability of the mix;  and the tendency to segregation.   Segregation is discussed on p. 223, but it should be observed here that the requirements of 
 workability and absence of segregation tend to be partially opposed to one another: the  easier it is for the particles of different sizes to pack, smaller particles passing into the  voids between the larger ones, the easier it is also for the small particles to be shaken out  of the voids, i.e. to segregate in the dry state. In actual fact, it 
is the mortar (i.e. a mixture  of sand, cement and water) that should be prevented from passing freely out of the voids  in the coarse aggregate. It is also essential  for the voids in the combined aggregate to be  sufficiently small to prevent  the cement paste from passing through and separating out.   The problem of segregation is thus rather s
imilar to that of filters, although the  requirements in the two cases are of course diametrically opposite: for the concrete to be  satisfactory it is essential that segregation be avoided.   There is a further requirement for a mix to be satisfactorily workable: it must contain a  sufficient amount of material smaller than a 300 um (No. 50 ASTM)
 sieve. Since the  cement particles are included in this material, a  richer mix requires a lower content of fine  sand than a lean mix. If the grading of sand is such that it is deficient in finer particles,  increasing the  fine / coarse aggregate ratio may not prove a satisfactory remedy, as it may  lead to an excess of middle sizes and possibl
y to harshness. (A mix is said to be harsh  when one size fraction is present in excess, as shown by a steep step in the middle of a  grading curve, so that particle interference results.) This need for an adequate amount of  fines (provided they are structurally sound) explains why minimum contents of particles  passing 300 um (No. 50 ASTM) and s
ometimes also 150 um (No. 100) sieves are laid  down, as for instance in Tables 3.23 and 3.24 (pp. 186 and 187). However, it is now  thought that the U.S. Bureau of Reclamation requirements of Table 3.24 for the minimum  percentage of particles passing the 300 and 150 um (Nos. 50 and 100 ASTM) sieves are  too high.   It may be further noted that b
lended cements (see page 84) would automatically provide  an adequate amount of fines. The presence of fines of all provenance (i.e. aggregate, filler,  and cement) can be assured by using the following total content of particles smaller than  125 um.  The volume of entrained air can be taken as equivalent to one-half the volume of fines and  shou
ld be included in the above figures.   German and Dutch mix specifications are essentially based on this approach.   The requirement that the aggregate occupies as large a relative volume as possible is in the  first instance an economic one, the aggregate being cheaper than the cement paste, but  there are also technical reasons why too rich a mi
x is undesirable. It has also been assumed  that the greater the amount of solid particles that can be packed into a given volume of  concrete the higher its strength. This maximum density theory has led to the advocacy of  grading curves parabolic in shape, or in part parabolic and then straight (when plotted to a  natural scale), as shown in Fig
. 3.13. It was found, however, that the aggregate graded to  give maximum density makes a harsh and somewhat unworkable mix. The workability is  improved when there is an excess of paste above that required to fill the voids in the sand,  and also an excess of mortar (sand plus cement) above that required to fill the voids in the  coarse aggregate
.    Fig. 3.13. Fuller's grading curve     Let us now consider the surface area of the aggregate particles. The water / cement ratio of  the mix is generally fixed from strength considerations. At the same time, the amount of  cement paste has to be sufficient to cover the surface of all the particles so that the lower  the surface area of the agg
regate the less paste, and therefore the less water, is required.   Taking for simplicity a sphere of diameter D as representative of the shape of the  aggregate we have the ratio of the surface area to volume of 6/D. This ratio of the surface  of the particles to their volume (or, when the particles have a constant specific gravity, to  their wei
ght) is called specific surface. For particles of a different shape, a coefficient other  than 6/D would be obtained but the surface area is still inversely proportional to the particle  size, as shown in Fig. 3.14 reproduced from Shacklock and Walker's report. It should be  noted that a logarithmic scale is used for both the ordinates and the abs
cissae since the sieve  sizes are in geometrical progression.    Fig. 3.14. Relation between specific surface and particle size     In the case of graded aggregate, the grading and the overall specific surface are related to  one another, although of course there are many grading curves corresponding to the same  specific surface. If the grading e
xtends to a larger maximum aggregate size, the overall  specific surface is reduced and the water requirement decreases, but the relation is not  linear. For instance increasing the maximum aggregate size from 10 mm to 63 mm  (3/8 in. to 2 1/2 in.) can, under certain conditions, reduce the water requirement for a  constant workability by as much a
s 50 kg per cubic metre (85 lb / yd3) of concrete. The  corresponding decrease in the water / cement ratio may be as much as 0.15. Some typical  values are shown in Fig. 3.15.    Fig. 3.15. Influence of maximum size of aggregate on mixing water requirement for a  constant slump     The practical limitations of the maximum size of aggregate that ca
n be used under given  circumstances and the problem of influence of the  maximum size on strength in general  are discussed on Page 195.   It can be seen that, having chosen the maximum size of aggregate and its grading, we can  express the total surface area of the particles using the specific surface as a parameter,  and it is the total surface
 of the aggregate that determines the water requirement or the  workability of the mix. Mix  design on the basis of the specific surface of the aggregate  was first suggested by Edwards as far back as 1918, and interest in this method was  renewed forty years later. Specific surface can be determined using the water permeability  method, but no si
mple field test is available, and a mathematical approach is made  difficult by the variability in the shape of different aggregate particles.   This, however, is not the only reason why the design of mixes on the basis of the specific  surface of aggregate is not universally recommended  The application of surface area  calculations was found to 
break down for aggregate particles smaller than about 150 um  (No. 100 ASTM) sieve, and for cement. These particles, and also some larger sand  particles, appear to act as a lubricant in the mix and do not seem to require wetting in  quite the same way as coarse particles. An indication of this can be obtained from some  results of tests by Glanvi
lle, Collins and Matthews, reproduced in part in Table	3.19.    Table 3.19: Water / Cement Ratio Required to Produce a Given Workability for Various  Amounts of Crusher Dust (Smaller than 150 um (No. 100) Sieve) in the Aggregate     Because specific surface gives a somewhat misleading picture of the workability to be  expected (largely owing to an
 overestimate of the effect of fine particles), an empirical  surface index was suggested by Murdock and its values as well as those of the  specific surface are given in Table 3.20.    Table 3.20: Relative Values of Surface Area and Surface Index     The overall effect of the surface area of an aggregate of given grading is obtained by  multiplyi
ng the percentage weight of any size fraction by the coefficient corresponding to  that fraction, and summing all the products. According to Murdock the surface index  (modified by an angularity index) should be used, and in fact the values of this index are  based on empirical results. On the other hand, Davey found that for the same total  speci
fic surface of the aggregate the water requirement and the compressive strength of  the concrete are the same for very wide limits of aggregate grading. This applies both to  continuously and gap-graded aggregate, and in fact three of the four gradings listed in  Table 3.21, reproduced from Davey's paper, are of the gap type.    Table 3.21: Proper
ties of Concretes Made with Aggregates of the Same Specific Surface     An increase in the specific surface of the aggregate for a constant water / cement ratio has  been found to lead to a lower strength of concrete, as shown for instance in Table 3.22,  reproducing Newman and Teychenne's results. The reasons for this are not quite clear,  but it
 is possible that a reduction in density of the concrete consequent upon an increase in  fineness of the aggregate is instrumental in lowering the strength.    Table 3.22: Specific Surface of Aggregate and Strength of Concrete for a 1:6 Mix with a  Water / Cement Ratio of 0.60     It seems then that the surface area of the aggregate is an importan
t factor in determining  the workability of the mix, but the exact role played by the finer particles has by no means  been ascertained.   The type gradings of Road Note No. 4 represent different values of overall specific  surface. For instance, when river sand and gravel are used, the four grading curves, Nos. 1  to 4, of Fig. 3.16 correspond to
 the specific surface of 1.6, 2.0, 2.5 and 3.3 m2 / kg,  respectively. In practice, when trying to approximate type gradings, the properties of  the mix will remain largely unaltered when compensation of a small deficiency of fines by a  somewhat larger excess of coarser particles is applied but the departure must not be too  great. The deficiency
 and excess are, of course, mutually interchangeable in the above  statement.    Fig. 3.16. Road Note No. 4 type grading curves for 19.05 mm (3/4 in.) aggregate     There is no doubt then that the grading of aggregate is a major factor in the workability of  a concrete mix. Workability, in turn, affects the water and cement requirements, controls 
 segregation, has some effect on bleeding, and influences the placing and finishing of the  concrete. These factors represent the important characteristics of fresh concrete and affect  also its properties in the hardened state: strength, shrinkage, and durability.   Grading is thus of vital importance in the proportioning of concrete mixes, but i
ts exact  role in mathematical terms is not fully known, and the behaviour of this type of semi-liquid  mixture of granular materials is still imperfectly understood.   Finally, it must be remembered that far more important than devising a "good" grading is  ensuring that the grading is kept constant; otherwise, variable workability results and, a
s	 this is usually corrected at the mixer by a variation in the water content, concrete of  variable strength is obtained.      Practical Gradings    From the brief review in the previous section it can be seen how important it is to use  aggregate with a grading such that a reasonable workability and a minimum segregation  are obtained. The impor
tance of the latter requirement cannot be over-emphasized: a  workable mixture which could produce a strong and economical concrete will result in a  honeycombed, weak, not durable and variable end product if segregation takes place.   The process of calculation of the proportions of aggregates of different size to achieve the  desired grading com
es within the scope of mix design, and is described in Chapter 10  Here, the properties of   some "good" grading curves will be discussed. It should be  remembered, however, that in practice the aggregate available locally or within an economic  distance has to be used, and this can generally produce satisfactory concrete, given an  intelligent ap
proach and sufficient care. The curves most commonly referred to as a basis of  comparison are those of the Road Research Note No. 4 on the Design of Concrete Mixes.  They have been prepared for aggregates of 19.05 and 38.1 mm (3/4 in. and 1 1/2 in.) maximum  size, and are reproduced in Figs. 3.16 and 3.17 respectively. Similar curves for aggregat
e	with a  9.52 mm (3/8 in.) maximum size have been prepared by McIntosh and Erntroy, and are shown  in Fig. 3.18.    Fig. 3.17. Road Note No. 4 type grading curves for 38.1 mm (1 1/2 in.) aggregate    Fig. 3.18. McIntosh and Erntroy's type grading curves for 9.52 mm (3/8 in.) aggregate     Four curves are shown for each maximum size of aggregate, 
but due to the presence of  over- and under-size aggregate and also because of  variation within any fraction size,  practical gradings are more likely to lie in  the vicinity of these curves than to follow them  exactly. It is therefore convenient to talk about grading zones, and these are marked on all  the  diagrams. In some specifications defi
nite limits of grading, rather than a  single curve,  are laid down.   Curve No. 1 represents the coarsest grading in each of the Figs. 3.16 to  3.18. Such a  grading is comparatively workable and can, therefore, be used  for mixes with a low  water / cement ratio or for rich mixes; it is, however, necessary to make sure that  segregation does not
 take place. At the other extreme, curve No. 4 represents a fine  grading: it will be cohesive but not very workable. In particular, an excess of material  between 1.20 and 4.76 mm (No. 16 and 3/16 in.) test sieves will produce a harsh concrete,  which, although it may be suitable for compaction by vibration, is difficult to place by hand.  If the
 same workability is to be obtained using aggregates with grading curves Nos. 1 and 4,  the latter would require a considerably higher water content: this would mean a lower strength  if both concretes are to have the same aggregate / cement ratio or, if the same strength is  required, the concrete made with the fine aggregate would have to be con
siderably richer,  i.e. each cubic metre would contain more cement than when the coarser grading is used.   The change between the extreme gradings is progressive. In the case of gradings lying  partly in one zone, partly in another, there is, however, a danger of segregation when too  many intermediate sizes are missing (cf. gap grading). If, on 
the other hand, there is an  excess of middle-sized aggregate the mix will be harsh and difficult to compact by hand  and possibly even by vibration. For this reason, it is preferable to use aggregate with  gradings similar to type rather than totally dissimilar ones.   Figs. 3.19 and 3.20 show the range of gradings used with 152.4 mm (6 in.) and 
76.2 mm  (3 in.) maximum aggregate size respectively, as given by McIntosh. The actual  gradings, as usual, run parallel with the limits rather than crossing over from one to the  other.    Fig. 3.19. Range of gradings used with 152.4 mm (6 in.) aggregate    Fig. 3.20. Range of gradings used with 76.2 mm (3 in.) aggregate     In practice, the use 
of separate fine and coarse aggregate means that a grading can be  made up to conform exactly with a type grading at one intermediate point, generally the 5  mm (3/16 in.) size. Good agreement can usually also be obtained at the ends of the curve  (150 um (No. 100) sieve and the maximum size used). If coarse aggregate is delivered in  single-size 
fractions, as is usually the case, agreement at additional points above 5 mm (3/16  in.) can be obtained, but for sizes below 5 mm (3/16 in.) blending of two or more sands is  necessary.      Grading of Fine and Coarse Aggregates    Since for any but unimportant work fine and coarse aggregates are batched separately, the  grading of each type of a
ggregate should be known and controlled.   Formerly, two classes of fine aggregate were recognized, but it has been shown that by  adjusting the ratio of the fine to coarse aggregate a good concrete could be obtained with  either class of aggregate. For this reason, in the 1954 revision of BS 882 the classification  of fine aggregate was altered t
o	four grading zones. The grading requirements for these  are reproduced in Table 3.23 and Fig. 3.21, and any fine aggregate whose grading falls  wholly within the limits of any one zone is considered suitable. A tolerance of a total  amount of 5 per cent on certain sieves is permitted, but the aggregate must not be finer  than the exact limits of
 the finest grading (No. 4) or coarser than the coarsest grading (No.  1). The only exception is in the case of crushed stone where 20 per cent is allowed to pass  the 150 um (No. 100) test sieve in all zones.    Fig. 3.21. Grading limits for sand in zones 1 to 4 of BS 882: 1973    Table 3.23: BS and ASTM Grading Requirements for Fine Aggregate   
	In BS 882 : 1973, the division into zones is based primarily on the percentage passing the  600 um (No. 30 ASTM) sieve, as shown by the values in Table 3.23. The main reason  for this is that a large number of sands divide themselves naturally at just that size, the  gradings above and below being approximately uniform. Furthermore, the content 
of  particles finer than the 600 um (No. 30 ASTM) sieve has a considerable influence on the  workability of the mix, and provides a fairly reliable index of the overall specific surface of  the sand.   For comparison, the requirements of ASTM Standard C 33-78 are in part included in  Table 3.23 (see Fig. 3.22). The limits of the latter specificati
on are much narrower than  the overall limits of BS 882 : 1973. The requirements of the U.S. Bureau of Reclamation  are given in Table 3.24. It may be noted that in the case of air-entrained concrete lower  quantities of the finest particles are acceptable, the entrained air acting effectively as very  fine aggregate. ASTM Standard C 33-78 also al
lows reduced percentages passing sieves  300 um and 150 um (No. 50 and No. 100 ASTM) when the cement content is above 297  kg / m3 (500 lb / yd3) or if air entrainment is used with at least 237 kg of cement per cubic  metre of concrete (400 1b / yd3).    Fig. 3.22. ASTM Standard C 33-78 grading limits for fine aggregate    Table 3.24: U.S. Bureau 
of Reclamation Grading Requirements for Fine Aggregate     Sand falling into any zone can generally be used in concrete, although under some  circumstances the suitability of a given sand may depend on the grading and shape of the  coarse aggregate.   The suitability of the fine sand of zone 4 for use in reinforced concrete has to be tested.  Sinc
e the greater part of this sand is smaller than the 600 um (No. 30 ASTM) sieve, a  gap-graded or a nearly gap-graded aggregate is obtained, and special care in choosing the  mix proportions must be exercised. The sand content of the mix should generally be low,  and suggested values of coarse / fine aggregate ratio are given in Table 3.25.    Tabl
e 3.25: Suggested Proportions by Weight of Coarse to Fine Aggregate for Sand of  Different Zones     Nevertheless, quite good concrete can be obtained with sand of zone 4, particularly using  vibration. Work at the Building Research Establishment has shown that increasing the  content of particles smaller than 150 um (No. 100) in crushed rock fine
 aggregate from  10 to 25 per cent results in only a small decrease in the compressive strength of concrete,  typically by 10 per cent.   At the other extreme, the coarse sand of zone 1 produces a harsh mix, and a high sand  content may be necessary for higher workability. This sand is more suitable for rich mixes  or for use in concrete of low wo
rkability.   Zone 2 represents a medium sand generally suitable for the "standard" 1:2 fine to coarse  mix (when the maximum size of aggregate is 20 mm (3/4 in.)).   In general terms, the ratio of coarse to fine aggregate should be higher the finer the  grading of the fine aggregate: typical values are given in Table 3.25. When crushed rock  coars
e aggregate is used, a slightly higher proportion of sand is required than with gravel  aggregate in order to compensate for the lowering of workability by the sharp, angular  shape of the crushed particles.   The choice of correct proportions is particularly important as the grading of the sand  approaches the fine outer limit of zone 4 or the co
arse outer  limit of zone 1. It is worth  noting, however, that, proportioned correctly,  fine sand can be utilized with success, and  this is of considerable economic importance in the United Kingdom where there is a  predominance of fine sand; in the past there has been a great deal of prejudice against this  type of material.   An example of us
ing sand from any of the four zones to produce equally good concrete is  given in Table 3.26, based on results obtained at the Building Research Station. The  actual sand gradings are shown in Fig. 3.21. An aggregate / cement ratio of 6.04 and a  water / cement ratio of 0.60, both by weight, were used throughout. To keep the  workability approxima
tely constant the coarse / fine aggregate ratio was varied so that the  overall specific surface of aggregate remained at 2.55 m2 / kg. Table 3.26 shows that  concrete of similar quality was obtained in all cases.    Table 3.26: Properties of Concretes Made with Aggregates of Constant Overall Specific Surface     By way of contrast, Table 3.27 sho
ws test results for the case when the same materials  were used but the coarse / fine aggregate ratio was kept constant. The use of finer sand led  to a higher water requirement and consequently a lower strength of concrete.    Table 3.26: Properties of Concretes Made with Aggregates of Fixed Proportions     The requirements of BS 882 : 1973 for t
he grading of coarse aggregate are reproduced in  Table	3.28: values are given both for graded aggregate and for nominal one-size fractions.  For comparison, some of the limits of ASTM Standard C 33-78 are given in Table 3.29.    Table 3.28: Grading Requirements for Coarse Aggregate According to BS 882 : 1973    Table 3.29: Grading Requirements fo
r Coarse Aggregate According to ASTM Standard C 33-78	 The actual grading requirements depend to some extent on the shape and surface  characteristics of the particles. For instance, sharp, angular particles with rough surfaces  should have a slightly finer grading in order to reduce the possibility of interlocking and to  compensate for the hi
gh friction between the particles. The actual grading of crushed  aggregate is affected primarily by the type of crushing plant employed. A roll granulator  usually produces fewer fines than other types of crushers, but the grading depends also on  the amount of material fed into the crusher.   The grading limits for all-in aggregate prescribed by
 BS 882 : 1973 are reproduced in  Table 3.30. It should be remembered that this type of aggregate is not used except for  small and unimportant jobs, mainly because it is difficult to avoid segregation in  stockpiling.    Table 3.30: Grading Requirements for All-in Aggregate According to BS 882 : 1973    Oversize and Undersize    Strict adherence 
to size limits of aggregate is not possible: breakage during handling will  produce some undersize material, and wear of screens in the quarry or at the crusher will  result in oversize particles being present.   In the United States it is usual to specify over- and undersize screen sizes as 7/6 and 5/6,  respectively, of the nominal sieve size; a
ctual values are given in Table 3.31. The quantity  of aggregate smaller than the undersize or larger than the oversize is generally severely  limited.    Table 3.31: Sizes of Over- and Undersize Screens of U.S. Bureau of Reclamation     The grading requirements of BS 882 : 1973 allow some under- and oversize both for  coarse and fine aggregate. T
he figures for the former are given in Table 3.28, and it can be  seen that between 5 and 15 per cent oversize is permitted. However, no aggregate must be  retained on a sieve one size larger (in the standard series) than the nominal maximum size.  5 to 10 per cent undersize material is allowed. In the case of single-size aggregate, some  undersiz
e is also allowed, and the amount passing the sieve next smaller than the nominal  size is also prescribed. It is important that this fine fraction of coarse aggregate be not  neglected in the calculation of the actual grading.   For fine aggregate, a total departure of 5 per cent from zone limits is allowed but not  beyond the coarser limit of zo
ne 1 or the finer limit of zone 4 (see Table 3.23).	Gap-Graded Aggregate    As mentioned earlier, aggregate particles of a given size pack so as to form voids that can  be penetrated only if the next smaller size of particles is sufficiently small. This means that  there must be a minimum difference between the sizes of any two adjacent parti
cle  fractions. In other words, sizes differing but little cannot be used side by side, and this has  led to advocacy of gap-graded aggregate.   Gap grading can then be defined as a grading in which one or more intermediate size  fractions are omitted. The term continuously graded is used to describe conventional  grading when it is necessary to d
istinguish it from gap grading. On a grading curve, gap  grading is represented by a horizontal line over the range of sizes omitted. For instance,  the top grading curve of Fig. 3.23 shows that no particles of size between 10.0 and  2.36 mm (3/8 in. and No. 8 ASTM) sieve are present. In some cases, a gap between  10.0 and 1.18 mm (3/8 in. and No.
 16 ASTM) sieve is considered suitable. Omission of  these sizes would reduce the number of stockpiles of aggregate required and lead to  economy. In the case of aggregate of 20.0 mm (3/4 in.) maximum size, there would be two  piles only: 20.0 to 10.0 mm and sand screened through a 1.18 mm (No. 16 ASTM) screen.  The particles smaller than 1.18 mm 
(No. 16 ASTM) sieve size could easily enter the voids  in the coarse aggregate so that the workability of the mix would be higher than that of a  continuously graded mix of the same sand content.    Fig. 3.23. Typical gap gradings     Tests by Shacklock have shown that, for a given aggregate / cement ratio and water / cement  ratio, a higher worka
bility is obtained with a lower sand content in the case of gap-graded  aggregate than when continuously graded aggregate is used. However, in the more  workable range of mixes, gap-graded aggregate showed a greater proneness to  segregation. For this reason, gap grading is recommended mainly for mixes of relatively  low workability that are to be
 compacted by vibration. Good control and, above all, care  in handling so as to avoid segregation, are essential.	It may be observed that even when some "ordinary" aggregates are used gap grading  exists; for instance, sand belonging to zone 4 of BS 882: 1973 is almost completely  deficient in particles between the 5.00 and 2.36 or 1.18 mm (3/1
6 in. and No. 8 or  No. 16 ASTM) sieve sizes. Thus, whenever we use a zone 4 sand without blending with  a coarser sand, we are in fact using a gap-graded aggregate.   Gap-graded aggregate can be used in any concrete, but there are two cases of interest:  preplaced aggregate concrete (see p. 254) and exposed aggregate concrete; in the latter, a  p
leasing finish is obtained since a large quantity of only one size of coarse aggregate  becomes exposed after treatment.   From time to time, various claims of superior properties have been made for concrete  made with gap-graded aggregate, but these do not seem to have been substantiated.  Strength, both compressive and tensile, does not appear t
o be affected. Likewise, Fig.  3.24, showing McIntosh's results, confirms that, using given materials with a fixed  aggregate / cement ratio (but adjusting the sand content), approximately the same  workability and strength are obtained with gap and continuous gradings; Brodda and  Weber reported a slight negative influence of gap grading on stren
gth.    Fig. 3.24. Workability and strength of 1:6 concretes made with gap- and continuously  graded aggregates     Similarly, there is no difference in shrinkage of the concretes made with aggregate of  either type of grading, although it might be expected that a framework of coarse  particles almost touching one another would result in a lower t
otal change in dimensions  on drying. The resistance of concrete to freezing and thawing is lower when gap-graded  aggregate is used.   It seems, therefore, that the rather extravagant claims made by advocates of gap grading  are not borne out. The explanation lies probably in the fact that, while gap grading makes  it possible for maximum packing
 of particles to occur, there is no means of ensuring that it  will occur. Both gap-graded and continuously graded aggregate can be used to make good  concrete, but in each case the right percentage of sand has to be chosen. Thus, once again,  it can be seen that we should not aim at any ideal grading but find the best combination of  the availabl
e aggregates.      Maximum Aggregate Size    It has been mentioned before that the larger the aggregate particle the smaller the surface  area to be wetted per unit weight. Thus, extending the grading of aggregate to a larger  maximum size lowers the water requirement of the mix, so that, for a specified workability  and richness, the water / ceme
nt ratio can be lowered with a consequent increase in  strength.   This behaviour has been verified by tests with aggregates up to 38.1 mm (1 1/2 in.)  maximum size, and is usually assumed to extend to larger sizes as well. Experimental  results show, however, that above the 38.l mm (1 1/2 in.) maximum size the gain in  strength due to the reduced
 water requirement is offset by the detrimental effects of lower  bond area (so that volume changes in the paste cause larger stresses at interface) and of  discontinuities introduced by the very large particles, particularly in rich mixes. Concrete  becomes grossly heterogeneous and the resultant lowering of strength may possibly be  similar to t
hat caused by a rise in the crystal size and coarseness of texture of rocks.   This adverse effect of increase in the size of the largest aggregate particles in the mix  exists, in fact, throughout the range of sizes, but below 38.1 mm (1 1/2 in.) the effect of  size  on the decrease in the water requirement is dominant. For larger sizes, the bala
nce of the  two  effects depends on richness of the mix as shown in Fig. 3.25. Thus the best maximum size  of aggregate from the standpoint of strength is a function of the richness of the mix.  Specifically, in lean concrete (165 kg of cement per cubic metre (280 lb / yd3)), the use of  150 mm (or 6 in.) aggregate is advantageous. However, in str
uctural concrete of usual  proportions, from the point of view of strength there is no advantage in using aggregate  with a maximum size greater than about 25 or 40 mm (1 or 1 1/2 in.). Moreover, the use of larger  aggregate would require the handling of a separate stockpile and might increase the risk of  segregation. However, a practical decisio
n would also be influenced by the availability and  cost of different size fractions. There are, of course, structural limitations too: the  maximum size of aggregate should be no more than one fifth to one quarter of the  thickness of the concrete section and is related also to the spacing of reinforcement. The  governing values are prescribed in
 codes of practice.    Fig. 3.25. Influence of maximum size of aggregate on the 28-day compressive strength of  concretes of different richness      Use of "Plums"    The original idea of the use of aggregate as an inert filler can be extended to the inclusion  of large stones in a normal concrete: thus the apparent yield for a given amount of cem
ent  is increased.   These stones are called "plums" and used in a large concrete mass they can be as big as a  300 mm (1 ft) cube but should not be greater than one-third of the least dimension to be  concreted. The volume of plums should not exceed 20 to 30 per cent of the total volume  of the finished concrete, and they have to be well disperse
d throughout the mass. This is  achieved by placing a layer of normal concrete, then spreading the plums, followed by  another layer of concrete, and so on. Each layer should be of such thickness as to ensure  at least 100 mm (4 in.) of concrete around each plum. Care must be taken to ensure that  no air is trapped underneath the stones and that t
he concrete does not work away from  their underside. The plums must have no adhering coating.   The placing of plums requires a large amount of labour and also breaks the continuity of  concreting. It is, therefore, not surprising that with the current high ratio of the cost of  labour to the cost of cement the use of plums is not economical exce
pt under special  circumstances, but their use is covered by standards in some countries, e.g. South Africa.      Handling of Aggregate    Handling and stockpiling of coarse aggregate can easily lead to segregation. This is  particularly so when discharging and tipping permits the aggregate to roll down a slope. A  natural case of such segregation
 is a scree (talus): the size of particles is uniformly graded  from largest at the bottom to smallest at the top.	A description of the precautions necessary in handling operations is outside the scope of  this book, but one vital recommendation should be mentioned: coarse aggregate should be  split into size fractions 5 to 10, 10 to 20, 20 to 4
0 mm (or 3/16 to 3/8, 3/8 to 3/4, 3/4 to 1  1/2 in.), etc. These fractions should be handled and stockpiled separately and remixed only  when being fed into the concrete mixer in the desired proportions. Thus segregation can  occur only within the narrow size range of each fraction, and even this can be reduced by  careful handling procedures.   C
are is necessary to avoid breakage of the aggregate: particles greater than 40 mm (or  1 1/2 in.) should be lowered into bins by means of rock ladders and not dropped from  a height.   On big and important jobs the results of segregation and breakage in handling (i.e. excess  of undersize particles) are eliminated by "finish rescreening" immediate
ly prior to feeding  into the batching bins over the mixer. The proportions of different sizes are thus controlled  much more effectively but the complexity and cost of the operations are correspondingly  increased. This is, however, repaid by easier placing of uniformly workable concrete and  by a possible saving in cement due to the uniformity o
f the concrete.  14    Properties of Hardened Concrete    The properties of fresh concrete are important only in the first few hours of its history  whereas the properties of hardened concrete assume an importance which is retained for  the remainder of the life of the concrete. The important properties of hardened concrete  are strength, deformat
ion under load, durability, permeability and shrinkage. In general,  strength is considered to be the most important property and the quality of concrete is  often judged by its strength. There are, however, many occasions when other properties  are more important, for example, low permeability and low shrinkage are required for  water-retaining s
tructures. Although in most cases an improvement in strength  results in an improvement of the other properties of concrete there are exceptions. For  example, increasing the cement content of a mix improves strength but results in higher  shrinkage which in extreme cases can adversely affect durability and permeability. One of  the primary object
ives of this chapter is to help the reader to understand the factors which  affect each of the important properties of hardened concrete.   Since the properties of concrete change with age and environment it is not  possible to attribute absolute values to any of them. Laboratory tests give only an  indication of the properties which concrete may 
have in the actual structure as the quality  of the concrete in the structure depends on the workmanship on site. For these reasons it is  important to be able to judge the quality of concrete in situ. The direct method of testing  drilled cores is expensive and limited because the removal of too many cores weakens a  structure. Nondestructive tes
ts have therefore been developed for the assessment of  concrete quality. The limitations and applications of nondestructive testing together with a  brief description of the techniques is given at the end of this chapter.      14.1 Strength    The strength of concrete is defined as the maximum load (stress) it can carry. As the  strength of concr
ete increases its other properties usually improve and since the tests for  strength, particularly in compression, are relatively simple to perform concrete  compressive strength is commonly used in the construction industry for the purpose of  specification and quality control. Concrete is a comparatively brittle material which is  relatively wea
k in tension.    Compressive Strength    The compressive strength of concrete is taken as the maximum compressive load it can  carry per unit area. Concrete strengths of up to 80 N mm-2 can be achieved by selective  use of the type of cement, mix proportions, method of compaction and curing conditions.  Concrete structures, except for road pavemen
ts, are normally designed on the  basis that concrete is capable of resisting only compression, the tension being carried by  steel reinforcement. In the United Kingdom a 150 mm cube is commonly used for  determining the compressive strength. The standard method described in BS 1881: Part  116 requires that the test specimen should be cured in wat
er at 20 +/- 2 C and crushed, by  loading it at a constant rate of stress increase of between 12 and 24 N mm-2 min-1,  immediately after it has been removed from the curing tank.    Tensile Strength    The tensile strength of concrete is of importance in the design of concrete roads and  runways. For example, its flexural strength or modulus of ru
pture (tensile strength in  bending) is utilised for distributing the concentrated loads over a wider area of road  pavement. Concrete members are also required to withstand tensile stresses resulting from  any restraint to contraction due to drying or temperature variation.   Unlike metals, it is difficult to measure concrete strength in direct t
ension and  indirect methods have been developed for assessing this property. Of these the split  cylinder test is the simplest and most widely used. This test is fully described in BS 1881:  Part 117 and entails diametrically loading a cylinder in compression along its entire length.  This form of loading induces tensile stresses over the loaded 
diametrical plane and the  cylinder splits along the loaded diameter. The magnitude of the induced tensile stress fct at  failure is given by    (formula)    where F is the maximum applied load and l and d are the cylinder length and diameter  respectively.   The flexural strength of concrete is another indirect tensile value which is also  common
ly used (BS 1881: Part 118). In this test a simply supported plain concrete beam is  loaded at its third points, the resulting bending moments inducing compressive and tensile  stresses in the top and bottom of the beam respectively. The beam fails in tension and the  flexural strength (modulus of rupture) fcr is defined by    (formula)    where F
 is the maximum applied load, L the distance between the supports, and b and d  are the beam breadth and depth respectively at the section at which failure occurs.   The tensile strength of concrete is usually taken to be about one-tenth of its  compressive strength. This may vary, however, depending on the method used for  measuring tensile stren
gth and the type of concrete. In general the direct tensile strength  and the split cylinder tensile strength vary from 5 to 13 per cent and the flexural strength  from 11 to 23 per cent of the concrete cube compressive strength. In each case, as the  strength increases the percentage decreases. As a guide, the modulus of rupture may be  taken as 
0.7 / (cube strength) N mm-2 and the direct tensile strength as 0.45 / (cube  strength) N mm- 2 although, where possible, values based on tests using the  actual concrete in question should be obtained.      14.2 Factors influencing Strength    Several factors which affect the strength of concrete are shown in figure 14.1. In this section their  i
nfluence is discussed with particular reference to compressive strength. In general, tensile strength is  affected in a similar manner.    Figure 14.1 Factors affecting strength of concrete    Influence of the Constituent Materials    Cement    The influence of cement on concrete strength, for given mix proportions, is determined by its fineness  
and chemical composition through the processes of hydration (see chapter 12). The gain in concrete  strength as the fineness of its cement particles increases is shown in figure 14.2. The gain in strength  is most marked at early ages and after 28 days the relative gain in strength is much reduced. At some  later age the strength of concrete made 
with fine cements may not be very different from that made  with normal cement (300 m2 kg- 1).   The role of the chemical composition of cement in the development of concrete  strength can best be appreciated by studying table 14.1 and figures 14.3 and 14.4. It is apparent that  cements containing a relatively high percentage of tricalcium silicat
e	(C3S) gain strength much more  rapidly than those rich in dicalcium silicate (C2S), as shown in figure 14.3; however, at later ages the  difference in the corresponding strength values is small. In fact there is a tendency for concretes made  with low-heat cements eventually to develop slightly higher strengths (figure 14.4). This is possibly  d
ue to the formation of a better quality gel structure in the course of hydration.    Figure	14.2 Effect of cement fineness on the development of concrete strength, based on  Bennett and Collings (1969)    TABLE 14.1 Chemical composition of various Portland cements with similar fineness    Figure 14.3 Development of strength of typical concrete mad
e	with different Portland cements  (see table 14.1)    Figure 14.4 Development of strength of typical concrete made with different Portland cements  (see table 14.1)    Water    A concrete mix containing the minimum amount of water required for complete hydration of its  cement, if it could be fully compacted, would develop the maximum attainable 
strength at any given  age. A water - cement ratio of approximately 0.25 (by weight) is required for full hydration of the  cement but with this water content a normal concrete mix would be extremely dry and virtually  impossible to compact.   A partially compacted mix will contain a large percentage of voids and the  concrete strength will drop. 
On the other hand, while facilitating placing and compaction, water in  excess of that required for full hydration produces a somewhat porous structure resulting from loss of  excess water, even for a fully compacted concrete. In practice a concrete mix is designed with a view  to obtaining maximum compaction under the given conditions. In such a 
concrete, as the ratio of  water to cement increases the strength decreases in a manner similar to that illustrated in figure 15.3  (see chapter 15).    Aggregate    When concrete is stressed, failure may originate within the aggregate, the matrix or at the  aggregate - matrix interface; or any combination of these may occur. In general the aggreg
ates are  stronger than the concrete itself and in such cases the aggregate strength has little effect on the strength  of concrete.   The bond (aggregate - matrix interface) is an important factor determining  concrete strength. Bond strength is influenced by the shape of the aggregate, its surface texture and  cleanliness. A smooth rounded aggre
gate will result in a weaker bond between the aggregate and  matrix than an angular or irregular aggregate or an aggregate with a rough surface texture. The  associated loss in strength however may be offset by the smaller water - cement ratio required for the  same workability. Aggregate shape and surface texture affect the tensile strength more 
than the  compressive strength. A fine coating of impurities, such as silt and clay, on the aggregate surface  hinders the development of a good bond. A weathered and decomposed layer on the aggregate can  also result in a poor bond as this layer can readily become detached from the sound aggregate beneath.   The aggregate size also affects the st
rength. For given mix proportions, the  concrete strength decreases as the maximum size of aggregate is increased. On the other hand, for a  given cement content and workability this effect is opposed by a reduction in the water requirement  for the larger aggregate. However, it is probable that beyond a certain size of aggregate there is no  obvi
ous advantage in further increasing the aggregate size except perhaps in some instances when  larger aggregate may be more readily available. The optimum maximum aggregate size varies with  the richness of the mix, being smaller for the richer mixes, and generally lies between 10 and 50 mm.  Concrete of a given strength can be produced with aggreg
ates having a variety of different gradings  provided due care is exercised to ensure that segregation does not occur. The suitability of a grading  to some extent depends on the shape and texture of the aggregate. Aggregates which react with the  alkali content of a cement adversely affect concrete strength. This is rarely a problem in the United
	Kingdom, even though some cases have been reported since 1976 (BRE Digest 258).    Admixtures    As a general rule, admixtures can only affect concrete strength by changing the hydration processes  and the air content of the mix and / or by enabling changes to be made to the mix proportions, most  importantly to the water - cement ratio. Acceler
ating admixtures increase the rate of hydration thereby  providing an increased early strength with little or no change at later ages unless the increased rate of  heat evolution causes internal cracking in which case a lower strength will result. In contrast, with  retarding admixtures the early strength of concrete is reduced owing to the delay 
in setting time.  Provided no air is entrained, the concrete strength will be approximately the same as that of the  control mix within a few days. Air entrainment in concrete will cause a reduction in strength at all  ages and to achieve a required strength the mix cement content has to be increased. However, in  practice the increased yield and 
improvement in workability is also taken into account in the mix  design when using air-entraining admixtures and there is generally no significant change in concrete  strength for the usual range of air content. Most water-reducing admixtures, including  superplasticizers, do not have any significant effect on the hydration of the cement. Thus, w
hen these  admixtures are used to improve workability no significant change in strength should be expected,  providing of course that the air content remains unchanged. On the other hand if the water content of  the mix is reduced while maintaining workability an increase in strength corresponding to the new  water - cement ratio will result.   It
 should be noted that the effect of particular admixtures in concrete depends on the precise  nature of the admixtures themselves, the constituent materials and proportions of the mix and the  ambient conditions (particularly temperature). In consequence, although for practical purposes the  above mentioned guidelines apply, to obtain an accurate 
estimate of their effect on concrete strength  in individual circumstances it is necessary to carry out trial mixes.    Influence of the Methods of Preparation    When concrete materials are not adequately mixed into a consistent and homogeneous mass, some  poor quality concrete is inevitably the result. Even when a concrete is adequately mixed ca
re must be  taken during placing and compaction to minimise the probability of the occurrence of bleeding,  segregation and honeycombing all of which can result in patches of poor quality concrete. A properly  designed concrete mix is one that does not demand the impossible from site operatives before it can be  fully compacted in its final locati
on. If full compaction is not achieved the resulting voids produce a  marked reduction in concrete strength.    Influence of Curing    Curing of concrete is a prerequisite for the hydration of the cement content. For a given concrete, the  amount and rate of hydration and furthermore the physical make-up of the hydration products are  dependent on
 the time - moisture - temperature history.   Generally speaking, the longer the period during which concrete is kept in  water, the greater its final strength. It is normally accepted that a concrete made with ordinary  Portland cement and kept in normal curing conditions will develop about 75 per cent of its final  strength in the first 28 days.
 This value varies with the nominal strength of concrete however,  increasing as concrete strength increases. The gain in strength with age up to the time of loading can  now be used only for estimating the static modulus of elasticity of concrete in structures (BS 8110:  Part 2). The age factors for strength prescribed for this purpose are given 
in table 14.2.    TABLE 14.2 Age factors for strength of concrete     The development of concrete strength under various curing conditions is  shown in figure 14.5. It is apparent that concrete left in air achieves the lowest strength values at all  ages owing to the evaporation of the free mixing water from the concrete. The gain in strength  dep
ends on a number of factors such as relative humidity, wind velocity and the size of structural  member of test specimen. Figure 14.5 shows that both the increased hydration due to improvements in  initial curing (moist or water curing at normal temperature) and the condition of the concrete at the  time of testing have a significant effect on the
 final apparent strength of concrete. It should also be  noted that moist (or water) curing after an initial period in air results in a resumption of the hydration  process and that concrete strength is further improved with time, although the optimum strength may  not be realised.    Figure 14.5 Effect of curing and condition of concrete when tes
ted on concrete strength,  based on Gilkey (1937)     The temperature at which concrete is cured has an important bearing on the  development of its strength with time. The rate of gain in strength of concrete made with ordinary  Portland cement increases with increase in concrete temperature at early ages (figure 14.6), although  at later ages th
e concrete made and cured at lower temperatures shows a somewhat higher strength.  Figure 14.7 shows how a high temperature during the placing and setting of concrete can adversely  affect the development of its strength from early ages. On the other hand, when the initial temperature  is lower than the subsequent curing temperature, then higher t
emperatures during final curing result  in significantly higher strengths (figure 14.8). A possible explanation for this behaviour is that a rapid  initial hydration appears to form a gel structure (hydration product) of an inferior quality and this  adversely affects concrete strength at later ages. Concretes made with other Portland cements woul
d  respond to temperature in a somewhat similar manner.    Figure 14.6 Comparative compressive strength of concrete cast, sealed and maintained at different  temperatures, based on Price (1951)    Figure 14.7 Comparative compressive strength of sealed concrete specimens maintained at  different temperatures for 2 hours after casting and subsequent
ly cured at 21 C, based on Price (1951)    Figure 14.8 Comparative compressive strength of concrete cast, sealed and maintained at 10 C for  the first 24 hours and subsequently cured at different temperatures, based on Price (1951)     It has been suggested that the strength of concrete can be related to the product  of age and curing temperature,
 commonly known as maturity. However, such relationships are  dependent on a number of factors such as curing temperature history, particularly the temperature at  early ages (see figures 14.7 and 14.8), and are therefore limited in their general applicability for  predicting concrete strength.    Influence of Test Conditions    The conditions und
er which tests to determine concrete strength are carried out can have a  considerable influence on the strength obtained and it is important that these effects are understood if  test results are to be correctly interpreted.    Specimen shape and size    Three basic shapes used for the determination of compressive strength are the cube, cylinder 
and  square prism. Each shape gives different strength results and furthermore for a given shape the  strength also varies with size. Figure 14.9 shows the influence of specimen diameter and it can be  seen that as the size decreases, the apparent strength increases. The measured strength of concrete is  also affected by the height - diameter rati
o (figure 14.10). For height - diameter ratios less than 2  strength begins to increase rapidly owing to the restraint provided by the machine platens. Strength  remains sensibly constant for height - diameter ratios between 2 and 3 and thereafter shows a slight  reduction. The relative influence of slenderness may be modified by several inherent 
characteristics of  concrete such as strength, air-entrainment, strength of aggregate, degree of moist curing and moisture  content at the time of testing.    Figure 14.9 Effect of specimen size on the apparent 28-day concrete compressive strength for  specimens with a height - diameter ratio of 2 and aggregate whose maximum diameter is  one-quart
er of the diameter of the specimen, based on Price (1951)    Figure 14.10 Effect on height - diameter ratio on concrete compressive strength for specimens  moist-cured at room temperature and test wet    BS 1881: Part 116 specifies the use of concrete cubes for determining compressive strength; 150 mm  cubes are widely used in the United Kingdom f
or quality-control purposes. However, cored cylindrical  specimens are used for measuring the compressive strength of concrete in situ and precast members  (BS 1881: Part 120). This standard gives a set of equations for converting the measured strength of a  core into an equivalent in situ cube strength. It should be noted that the estimation of e
quivalent  standard cube strengths from core strengths is no longer considered to be valid, because of large  variations in the relationship between the two strengths arising from variations in site conditions  including percentage of reinforcement, dimensions of members and the methods of compaction and  curing. The effect on the measured strengt
h of the variations in the height - diameter ratio of drilled  cores is taken into account by using the correction factors given in table 14.3. The effect of the shape  of the test specimen is taken into account by multiplying the cylinder (core) strength, for a cylinder  (core) having a height - diameter ratio of 2, by 1.25 to obtain the equivale
nt cube strength. Since the  relationship between height - diameter ratio and strength depends on the type of concrete, the use of  one set of correction factors of the type given in table 14.3 can only be suitable for a limited range of  concrete materials. It should be noted that the correction factors given in table 14.3 are most likely to  fav
our low-strength concretes. The estimation of the equivalent in situ cube strength from  compression tests performed on cored cylindrical specimen can be further influenced by a multitude of  other factors and the reader is advised to consult the Concrete Society Technical Report on concrete  core testing for strength (1976), the British Standard 
Guide on Assessment of Concrete Strength in  Existing Structures (BS 6089), Murphy (1984) and Munday and Dhir (1984).    TABLE 14.3 Correction factor for compression tests on cylinders (cores)    Specimen moisture content and temperature    This should not be confused with the effect of moisture and temperature during curing. The strength  of conc
rete can be influenced by the absence or presence of moisture and by temperature only when  these conditions generate internal stresses which change the magnitude of the external load required  to bring about failure. Since the mode of failure in different strength tests is different, it follows that  the influence of moisture content and temperat
ure on the apparent strength varies.   In the case of compression tests, air-dry concrete has a significantly higher  strength than concrete tested in a saturated condition (figure 14.5). The lower strengths of wet  concrete can be attributed mainly to the development of internal pore pressure as the external load is  applied.   The flexural stren
gth of saturated concrete is greater however than that of  concrete which is only partially dry owing to tensile stresses developed near the surface of the dry  concrete by differential shrinkage. The initial drying is therefore critical with the apparent strength  reaching a minimum value within the first few days; thereafter it begins to increas
e gradually as the  concrete approaches a completely dry state. A thoroughly dry concrete specimen in which tensile  stresses are not being induced has a higher apparent flexural strength than saturated concrete. The  indirect tensile or split cylinder strength is lower for saturated concrete than for thoroughly dried  concrete.   Since the influe
nce of the moisture content on concrete strength varies with  the type of test, standard strength tests (BS 1881: Part 116) should be performed on specimens in a  saturated condition.    Method of loading    The compressive strength of concrete increases as the lateral confining pressure increases (figure  14.11).    Figure 14.11 Effect of lateral
 compression on concrete compressive strength     The apparent strength of concrete is affected by the rate at which it is loaded.  In general, for static loading, the faster the loading rate the higher the indicated strength. However,  the relative effects of the rate of loading vary with the nominal strength, age and extent of moist  curing. Hig
h-strength mature concretes cured in water are most sensitive to loading rate and  particularly so for loading rates greater than 600 N mm-2 min-1. BS 1881: Part 116 requires concrete  in compression tests to be loaded at 12 to 24 N mm-2 min-1 and within this small range of loading  rates variations in the measured strength of concrete will be ins
ignificant. The standard rates of  loading for flexural and split cylinder tests correspond to rates of increase of tensile stress of 1.2 to 6.0  N mm-2 min-1 and 1.2 to 2.4 N mm-2 min-1 respectively (BS 1881: Parts 117 and 118). For the  flexural test the standard requires the use of the lower loading rates for low strength concretes and the  hig
her loading rates for high strength concretes.   When loads on a structure are predominantly cyclic (repeated loading and  unloading) in character, the effects of fatigue should also be considered. This kind of loading produces  a reduction in strength. A reduction in strength of as much as 30 per cent of the normal static strength  value can take
 place, although this depends on the stress - strength ratio, the frequency of loading and  the type of concrete.   Structural concrete is commonly subjected to sustained loads. It is probable that  concrete can withstand higher loads if a constant load is maintained before  loading to failure. Improvement in compressive strength can occur for sus
tained loads up to 85 per  cent of the normal static strength although the actual gain in strength depends on the duration and  magnitude of load, type of concrete and age. The increase in strength is probably due to consolidation  of the concrete under sustained load and the redistribution of stresses within the concrete.      14.3 Deformation   
 Concrete deforms under load, the deformation increasing with the applied load and being commonly  known as elastic deformation (figure 14.12). Concrete continues to deform with time, under constant  load; this is known as time-dependent deformation or creep (figure 14.13). Deformation due to  concrete shrinkage is discussed in section 14.5 and is
 specifically excluded here. Another important  consideration in design is the deformation resulting from changes in temperature. The coefficient of  thermal expansion for concrete is about 12 x 10'-6 per C.    Figure 14.12 A typical stress - strain curve showing different moduli of elasticity    Figure 14.13 A typical illustration of deformation 
of concrete subjected to constant load    Elastic Deformation    Unlike that for metals, the load - deformation relationship for concrete subjected to a continuously  increasing load is nonlinear in character. The nonlinearity is most marked at higher loads. When the  applied load is released the concrete does not fully recover its original shape 
(see figure 14.12). Under  repeated loading and unloading the deformation at a given load level increases, although at a  decreasing rate, with each successive cycle. All these characteristics of concrete indicate that  it should be considered as a quasi-elastic material and when computing the elastic constants, namely,  the modulus of elasticity 
and Poisson's ratio, the method employed should be clearly stated. It is only  for simplicity and convenience that the elastic modulus is assumed to be constant in both concrete  technology and the design of concrete structures.    Modulus  of elasticity    This is defined as the ratio of load per unit area (stress) to the elastic deformation per 
unit length  (strain)    modulus of elasticity E = stress / strain = (formula).    The modulus is used when estimating the deformation, deflection or stresses under normal working  loads. Since concrete is not a perfectly elastic material, the modulus of elasticity depends on the  particular definition adopted (see figure 14.12).   The initial tan
gent modulus is of little value for structural applications since it  has significance only for low stresses during the first load cycle. The tangent modulus is difficult to  determine and is applicable only within a narrow band of stress levels. The secant modulus is easily  determined and takes into account the total deformation at any one point
. The method prescribed in  BS 1881: Part 121 requires repeated loading and unloading before the specimen is loaded for  determination of its secant modulus of elasticity from a stress - strain curve, which then approaches a  straight line for stresses up to one-third of the concrete cylinder strength. The modulus of elasticity for  most concretes
, at 28 days, ranges from 15 to 40 kN mm-2. As a guide, the modulus may be assumed  to be 3.8 / (concrete strength, N mm-2) kN mm-2 for normal weight concrete. For structural design  purposes, the short term elastic modulus values for normal weight concrete given in BS 8110: Part 2  may be used.    Poisson's ratio    When concrete is subjected to 
axial compression, it contracts in the axial direction and expands  laterally. Poisson's ratio is defined as the ratio of the lateral strain to the associated axial strain and  varies from 0.1 to 0.3 for normal working stresses. A value of 0.2 is commonly used.    Factors influencing the Elastic Behaviour of Concrete    Concrete is a multiphase ma
terial and its resistance to deformation under load is dependent on the  stiffness of its various phases such as aggregate, cement paste and voids, and the interaction between  individual phases. In general the factors which influence the strength of concrete also affect  deformation although the extent of their influence may well vary. The modulu
s of elasticity increases  with strength, although the two properties are not directly related because different factors exert varying  degrees of influence on strength and modulus of elasticity. Although relationships between the two  properties may be derived they are only applicable within the range of variables considered. In short  there is n
o unique relationship between strength and modulus of elasticity.   The influence of stress level and rate of loading on both the axial and lateral  strains is shown in figure 14.14 for a concrete which has been cured in water (20 C +/- 1 C) and  air-dried before loading in uniaxial compression. It should be noted that the slower the rate of loadi
ng  the greater the strains at a given stress level. This is due to the fact that during loading both elastic and  creep strains occur, the creep strains increasing with the duration of time for which the load acts. The  variation of tangent modulus and Poisson's ratio, for the same concrete, with stress level and rate of  loading are shown in fig
ure 14.15. The general trends depicted in the figure apply to all concretes  although the exact relationship may vary with the type of concrete and the methods employed for  measuring the properties.    Figure 14.14 Stress - strain relationships of concrete for different rates of loading, based on  Dhir and Sangha (1972)    Figure 14.15 Variation 
of tangent modulus and Poisson's ratio with stress level and rate of  loading, based on Dhir and Sangha (1972)    Creep Deformation    When concrete is subjected to a sustained load it first undergoes an instantaneous deformation  (elastic) and thereafter continues to deform with time (figure 14.13). The increase in strain with time  is termed cre
ep. It should be noted that after an initially high rate of creep the creep continues but at a  continuously decreasing rate, except when the sustained load is large enough to cause failure, in  which case the rate of creep increases before failure occurs. The removal of the sustained load results in  an immediate reduction in strain and this is f
ollowed by a gradual decrease in strain over a period of  time. The gradual decrease in strain is called creep recovery. Creep and creep recovery are related  phenomena with somewhat similar characteristics. Creep is not wholly reversible and some permanent  strain remains after creep recovery is complete.   Since concrete in service is subjected 
to sustained loads for long periods of time the effect of  creep strains, which normally exceed the elastic strains, must be considered in structural design.  Failure to include the effects of creep strains may lead to serious underestimates of beam deflections  and over-all structural deformation and, in those cases where structural stability is 
involved. the  provision of members with inadequate strength. In prestressed concrete, allowance must be made for  the loss of tension in the prestressing tendons resulting from the shortening of a member under the  action of creep. Creep strain can also be beneficial in that it can relieve local stress concentrations  which might otherwise lead t
o structural damage. A classic example of this is the reduction of  shrinkage stresses in restrained members.    Factors influencing creep    Both the type of concrete (as described by its ingredients, curing history, strength and age) and the  relative magnitude of the applied stress with respect to concrete strength (stress - strength ratio) aff
ect  the creep strain. For a given concrete the creep strain is almost directly proportional to the  stress - strength ratio for ratios up to about one-third. For the same stress - strength ratio, the creep  strain increases as both cement content and water - cement ratio increase (figure 14.16) and decreases as  the relative humidity and age at l
oading increase (figure 14.17).    Figure 14.16 Effect of cement content and water - cement ratio on ultimate creep strain of concrete  at 50 per cent relative humidity loaded at an age of 28 days, based on CEB-FIP (1970)    Figure 14.17 Effect of relative humidity and age at loading on creep, based on Neville (1970)     The influence of the const
ituent materials of concrete on creep is somewhat  complex. The different types of cement influence creep because of the associated different rates of  gain in concrete strength. For example, concrete made with rapid-hardening Portland cement shows  less creep than concrete made with ordinary Portland cement and loaded at the same age. Normal  roc
k aggregates have a restraining effect on concrete creep and the use of large, high modulus  aggregates can be beneficial in this respect.   The environmental conditions of concrete subjected to a sustained load can  also have a marked effect on the magnitude of creep. For example, concrete cured under humid  conditions and then loaded in a relati
vely dry atmosphere undergoes 3 greater creep strain than if the  original conditions had been maintained. A concrete allowed to reach moisture equilibrium before  loading, however, is not adversely affected. Creep decreases as the mass of concrete increases owing  to the slower rate at which loss of water can take place within a large mass of con
crete.   Several mathematical equations have been proposed for estimating creep  strains. Both the rate of creep and its ultimate magnitude are dependent on a multitude of factors.  These expressions can only be truly applicable to concretes similar to those for which they were  designed. Nevertheless, creep strains are usually estimated for desig
n purposes. It has been noted that,  for a given concrete, creep strain depends on the stress - strength ratio. For practical purposes,  concrete creep strain may be assumed to be directly proportional to the elastic deformation up to a  stress - strength ratio of about two-thirds. On this assumption the ultimate creep strain may be  estimated for
 a given water - cement ratio and cement content using figure 14.16. This creep value  may be modified for different environmental moisture conditions and ages on loading using figure  14.17. It may be assumed that 50 per cent of the total estimated creep will take place within one  month and 75 per cent within the first six months after loading. 
     14.4 Durability    Besides its ability to sustain loads, concrete is also required to be durable. The durability of concrete  can be defined as its resistance to deterioration resulting from external and internal causes (figure  14.18). The external causes include the effects of environmental and service conditions to which  concrete is subje
cted, such as weathering, chemical actions and wear. The internal causes are the  effects of salts, particularly chlorides and sulphates, in the constituent materials, interaction between  the constituent materials, such as alkali - aggregate reaction, volume changes, absorption and  permeability.    Figure 14.18 Factors affecting durability of co
ncrete     In order to produce a durable concrete care should be taken to select suitable  constituent materials. It is also important that the mix contains adequate quantities of materials in  proportions suitable for producing a homogeneous and fully compacted concrete mass.    Weathering    Deterioration of concrete by weathering is usually bro
ught about by the disruptive action of alternate  freezing and thawing of free water within the concrete and expansion and contraction of the concrete,  under restraint, resulting from variations in temperature and alternate wetting and drying.   Damage to concrete from freezing and thawing arises from the expansion of  porewater during freezing; 
in a condition of restraint, if repeated a sufficient number of times, this  results in the development of hydraulic pressure capable of disrupting concrete. Road kerbs and slabs,  dams and reservoirs are very susceptible to frost action.   The resistance of concrete to freezing and thawing can be improved by increasing  its impermeability. This c
an be achieved by using a mix with the lowest possible water - cement ratio  compatible with sufficient workability for placing and compacting into a homogeneous mass.  Durability can be further improved by using air-entrainment, an air content of 3 to 6 per cent of the  volume of concrete normally being adequate for most applications. The use of 
air-entrained concrete  is particularly useful for roads where salts are used for de-icing. BS 8110: Part 1 recommends the use  of entrained air for a concrete of characteristic strength below 50 N mm-2 where it is likely to be  exposed to freezing and thawing actions while wet and its surfaces are subject to the effects of  de-icing salts. The le
vel of air content required varies with the maximum size of aggregate used, ranging  from an average of 7 per cent for 10 mm to 4 per cent for 40 mm nominal maximum aggregate sizes.  It is also important that, wherever possible, provision be made for adequate drainage of exposed  concrete surfaces. A dry concrete is not affected by freezing. The t
ype of cement used has no effect  although during the very early stages of hydration the use of a cement with a high rate of hydration  can be beneficial. Damage to a structure resulting from the expansion and contraction of concrete  should be minimised by providing joints which permit such movement without restraint.    Chemical Attack    In gen
eral, concrete has a low resistance to chemical attack. There are several chemical agents which  react with concrete but the most common forms of attack are those associated with leaching,  carbonation, chlorides and sulphates. Chemical agents essentially react with certain compounds of the  hardened cement paste and the resistance of concrete to 
chemical attack therefore can be affected by  the type of cement used. The resistance to chemical attack improves with increased impermeability.    Leaching    Calcium hydroxide, Ca(OH)2, in hardened cement paste dissolves readily in water . Thus, if concrete in  service absorbs or permits the passage of water through it the calcium hydroxide in t
he hardened  cement is removed, or leached out. Leaching can seriously impair the durability of concrete. Hydraulic  structures in which water may pass through cracks, areas of segregated or porous concrete or along  poor construction joints suffer from this kind of attack. Concrete may also absorb rain or groundwater  and the presence of carbon d
ioxide in such waters enhances the process of leaching. Free carbon  dioxide and organic acids are found in acidic peaty groundwater.   A homogeneous and dense concrete with low permeability significantly reduces  the effectiveness of the leaching action. The use of pozzolanas (e g. pulverized-fuel ash) can also be  helpful, as it reduces the perm
eability of concrete and fixes Ca(OH)2, thus lowering the soluble lime  in concrete. Care should be taken in selecting and proportioning the constituent materials and in  curing the concrete to ensure that shrinkage cracking is minimised.    Carbonation    Carbon dioxide (CO2), in the presence of moisture, reacts with Ca(OH)2 present in hydrated c
oncrete  to form calcium carbonate (CaCO3). This process is known as the carbonation of concrete, which can  take place even with the level of CO2 present in the normal atmosphere. Carbonation continues from  the concrete surface, including the surfaces of any cracks, throughout the life of the concrete but the  products of carbonation reduce the 
rate of carbonation as the depth of the carbonated layer increases.  The data available show that, for a given concrete, the depth of carbonation is roughly proportional to  the square-root of time. The key factor in the advancement of the carbonation front is the permeability  of concrete, which in turn depends on the mix proportions, particularl
y the water - cement ratio, and  how well the concrete has been compacted and cured. The environmental conditions that are also  important are the relative humidity and temperature. The rate of carbonation increases with  temperature and this effect is more pronounced at relative humidities within the optimum range of 50  to 75 per cent. Thus, in 
certain Middle East regions depths of carbonation much greater than those  generally encountered in the UK have been reported (Smith and Evans, 1986).   The effect of carbonation on concrete durability is significant only for concrete  in which steel is embedded (see chapter 5). Concrete is a highly alkaline material (pH about 13). This  is largel
y due to the presence of soluble Ca(OH)2, although sodium and potassium oxides also  contribute a little. This alkalinity results in the formation of a passive oxide film on the surface of  embedded steel in concrete which protects it against corrosion. The carbonation of concrete depletes  Ca(OH)2 present and thereby lowers the alkalinity of carb
onated concrete to below 9, at which point  the protective oxide film is destroyed. Thus, corrosion can occur when carbonation in concrete reaches  embedded steel. The products of corrosion occupy a volume greater than that of the steel which has  been affected and the associated bursting forces can result in significant cracking and subsequent  s
palling of the concrete in these locations. The provision of adequate concrete cover, that is a depth of  cover greater than the anticipated depth of carbonation, to all otherwise unprotected embedded steel,  is essential if corrosion resulting from the effect of carbonation is to be avoided.    Chloride attack    Soluble chlorides in de-icing sal
ts or occurring naturally in soils, seawater and ground water can enter  concrete by absorption through its surface, by capillary attraction along interconnected voids or by  direct access through cracks in the concrete. For plain concrete the presence of chlorides, in this  context, is not generally a cause for concern in the UK, although salt we
athering (see Sulphate attack)  does cause problems in some other regions. For concrete in which steel is embedded the presence of  chlorides is a potentially more serious condition, particularly in prestressed concrete.   Chloride ions in concrete adjacent to the steel can readily destroy, even at  high alkalinities, the passive oxide film which 
normally protects the steel against corrosion (see  Carbonation). The presence of chloride ions not only effectively initiates the electro-chemical process  of corrosion, subject to the availability of moisture and oxygen, but it also increases the electrical  conductivity of the concrete and thus results in an increased corrosion current leading 
to an  accelerated rate of corrosion. The disruptive effects of the products of corrosion have been discussed  previously (see Carbonation). For a more detailed discussion of reinforcement corrosion the reader is  referred to the Report by A.C.I. Committee 222 (1985).   Adequate protection against external sources of soluble chlorides can generall
y be achieved  by the use of properly designed mixes producing dense impermeable concrete with adequate concrete  cover to the embedded steel, together with careful attention to all aspects of detail design particularly  those associated with crack control. Where appropriate a protective coating or lining can also be  provided.   When significant 
quantities of chlorides are introduced into concrete at the mixing stage,  through the choice of constituent materials (such as water, aggregates or admixtures containing  calcium chloride), little can generally be done afterwards to rectify the situation. The presence of  internal chlorides not only promotes more rapid corrosion of the steel, but
 also has a number of other  undesirable effects (Fookes et al. ,1976) including the reduced resistance of concrete to sulphate  attack. Limits on the total chloride content of concrete mixes are prescribed in BS 8110: Part 1 to  ensure adequate durability in this context.    Sulphate attack    Most sulphate solutions react with the calcium hydrox
ide, Ca(OH)2, and calcium aluminate, C3A, of  hydrated cement to form calcium sulphate and calcium sulphoaluminate compounds. Calcium and  magnesium sulphates are most active and occur widely in soils (particularly clays), groundwater and  seawater. Although these compounds, unlike calcium hydroxide, do not readily dissolve in water  their volume 
is greater than the volume of the compounds of cement paste from which they are  formed. This increase in volume within the hardened concrete contributes towards the breakdown of  its structure.   The intensity and rate of sulphate attack depend on a number of factors such  as type of sulphate (magnesium sulphate is the most vigorous), its concent
ration and the continuity of  its supply to concrete. The concentration of sulphates in solution is expressed in parts of SO3 per  million (p.p.m.), by weight. Permeability and the presence of cracks also affect the severity of the  attack. The type of cement is a very important factor and the resistance of various cements to sulphate  attack incr
eases in the following order: ordinary and rapid-hardening Portland cement, Portland  blastfurnace cement and low-heat Portland cement, sulphate-resisting Portland cement and  supersulphated cement and finally high-alumina cement. As previously stated, owing to their inherent  ability to fix the Ca(OH)2 liberated by hydration of Portland cements, 
pozzolanas such as  pulverized-fuel ash can be employed to effect improvement in the resistance of concrete to sulphate  attack. Chlorides reduce the resistance of concrete to sulphate attack. The use of ground granulated  blastfurnace slag (ggbfs) also has a beneficial effect against sulphate attack.   To protect concrete from sulphate attack eve
ry effort should be made to produce an  impermeable concrete. Requirements given in BS 8110: Part 1 regarding the type of cement,  water - cement ratio and minimum cement content for concrete exposed to various degrees of sulphate  attack are reproduced in table 14.4. The number of construction joints should be minimised since these  can be partic
ularly prone to attack. For concrete structures housed in sulphate-bearing soils, protective  linings such as the various proprietary self-adhesive membranes, or protective coatings, such as  bitumens, tars and epoxy resins, may also be applied on exterior surfaces although the long-term  effectiveness of some coatings cannot always be assured.   
 TABLE 14.4 Requirements for concrete exposed to sulphate attack, based on BS 8110: Part 1     Another form of external attack, not usually found in the UK but widely reported in regions  with climate and geological environments normally associated with desert terrain, is that known as  salt weathering (Fookes et al., 1976). This is typified by a 
physical disintegration of the concrete  surface resulting from entry of soluble salts (sulphates and / or chlorides) into the pore spaces within  the concrete matrix and the pores in the aggregates, and their subsequent crystallisation within these  pores. Internal stresses induced in the concrete by the volume changes of the salt crystals, assoc
iated  with moisture and temperature changes can lead to cracking and subsequent disintegration of the  surface concrete. Progressive deterioration of the surface, which commonly occurs within a short  distance of the ground surface, might normally be expected. This form of attack is unaffected by the  type of cement and can only be completely eli
minated by the provision of an adequate protective  lining or surface coating.   External forms of protection together with the use, where appropriate, of sulphate resistant  cements, will not be fully effective in the prevention of sulphate attack if significant quantities of  sulphates have already been introduced into the concrete, via its cons
tituent materials, during mixing.  Limits on the soluble sulphate content of concrete mixes are prescribed in BS 8110: Part 1 to minimise,  in a practical manner, the possible effects of internal sulphates (Fookes et al., 1976).    Wear    The main causes of wear of concrete are the cavitation effects of fast-moving water, abrasive material  in wa
ter, wind blasting and attrition and impact of traffic. Certain conditions of hydraulic flow result  in the formation of cavities between the flowing water and the concrete surface. These cavities are  usually filled with water vapour charged with extraordinarily high energy and repeated contact with  the concrete surface results in the formation 
of pits and holes, known as cavitation erosion. Since even  a good-quality concrete will not be able to resist this kind of deterioration the best remedy is therefore  the elimination of cavitation by producing smooth hydraulic flow. Where necessary, the critical areas  may be lined with materials having greater resistance to cavitation erosion.  
 In general, the resistance of concrete to erosion and abrasion increases with  increase in strength. The use of a hard and tough aggregate tends to improve concrete resistance to  wear.    Alkali - Aggregate Reactions    Certain natural aggregates react chemically with the alkalis present in Portland cement. When this  happens these aggregates ex
pand or swell resulting in cracking and disintegration of concrete.   The most common form of alkali - aggregate reaction (AAR) is alkali - silica reaction (ASR)  which is associated with siliceous rocks containing silica in its active (amorphous) form such as cherts,  siliceous limestones and certain volcanic rocks. The new alkali silicate gel fo
rmed as a result can imbibe  a considerable amount of water, thus producing a volume expansion and, where this expansion is  sufficient, an eventual disruption of concrete. The minimum alkali content of cement required to  produce enough alkali silicate gel (expansive reaction) to damage concrete is 0.6 per cent of the soda  equivalent. Moisture i
s necessary for the alkali - silica reaction and increased temperatures accelerate  this reaction. The simplest preventive measure for this type of deterioration is not to use alkali reactive  aggregate with cements having a high alkali content. Adverse effects of the alkali - silica reaction can be  minimised by using suitable proportions of eith
er pulverized-fuel ash (pfa) or ground granulated  blastfurnace slag (ggbfs) in the concrete mix. For further guidance on minimising the risk of ASR in  concrete, the reader is referred to BRE Digest 258 and the Concrete Society Report (1985).   The second type of expansive alkali - aggregate reaction is known as alkali - carbonate reaction  (ACR)
, which occurs with certain argillaceous dolomitic limestones containing some clay, although  not all such reactive rocks necessarily produce deleterious expansions. The environmental factors  affecting the alkali - carbonate reaction and the preventative measures are essentially the same as for the  alkali - silica reaction, but there is some dou
bt about the effectiveness of pozzolanic materials in  controlling the alkali - carbonate reaction. The reactive carbonate rocks are relatively more rare in  occurrence than the reactive silica rocks.    Volume Changes    Principal factors responsible for volume changes are the chemical combination of water and cement  and the subsequent drying of
 concrete, variations in temperature and alternate wetting and drying.  When a change in volume is resisted by internal or external forces this can produce cracking. The  greater the imposed restraint, the more severe the cracking. The presence of cracks in concrete  reduces its resistance to the action of leaching, corrosion of reinforcement, att
ack by sulphates and  other chemicals, alkali - aggregate reaction and freezing and thawing, all of which may lead to  disruption of concrete. Severe cracking can lead to complete disintegration of the concrete surface  particularly when this is accompanied by alternate expansion and contraction.   Volume changes can be minimised by using suitable
 constituent materials  and mix proportions having due regard to the size of structure. Adequate moist curing is also essential  to minimise the effects of any volume changes.    Permeability and Absorption    Permeability refers to the ease with which water can pass through the concrete. This should not be  confused with the absorption property o
f concrete and the two are not necessarily related. Absorption  may be defined as the ability of concrete to draw water into its voids. Low permeability is an  important requirement for hydraulic structures and in some cases watertightness of concrete may be  considered to be more significant than strength although, other conditions being equal, c
oncrete of  low permeability will also be strong and durable. A concrete which readily absorbs water is  susceptible to deterioration.   Concrete is inherently a porous material. This arises from the use of water in excess of that  required for the purpose of hydration in order to make the mix sufficiently workable and the difficulty  of completel
y removing all the air from the concrete during compaction. If the voids are  interconnected concrete becomes pervious although with normal care concrete is sufficiently  impermeable for most purposes. Concrete of low permeability can be obtained by suitable selection of  its constituent materials and their proportions followed by careful placing,
 compaction and curing. In  general for a fully compacted concrete, the permeability decreases with decreasing water - cement  ratio. Permeability is affected by both the fineness and the chemical composition of cement. Coarse  cements tend to produce pastes with relatively high porosity. Aggregates of low porosity are preferable  when concrete wi
th a low permeability is required. Segregation of the constituent materials during  placing can adversely affect the impermeability of concrete.      14.5 Shrinkage    Shrinkage of concrete is caused by the settlement of solids and the loss of free water from the plastic  concrete (plastic shrinkage), by the chemical combination of cement with wat
er (autogenous  shrinkage) and by the drying of concrete (drying shrinkage). Where movement of the concrete is  restrained, shrinkage will produce tensile stresses within the concrete which may cause cracking.  Most concrete structures experience a gradual drying out and the effects of drying shrinkage should  be minimised by the provision of move
ment joints and careful attention to detail at the design stage.    Plastic Shrinkage    Shrinkage which takes place before concrete has set is known as plastic shrinkage. This occurs as a  result of the loss of free water with, or without, significant settlements of solids in the mix. Since  evaporation usually accounts for a large proportion of 
the water losses plastic shrinkage is most  common in slab construction and is characterised by the appearance of surface cracks which can  extend quite deeply into the concrete. Crack patterns associated with significant settlement (plastic  settlement cracks) generally coincide with the line of the reinforcement. Preventive measures are  usually
 based on methods of reducing water loss. This can be achieved in practice by making the mix  more cohesive and by covering concrete with wet hessian or polythene sheets or by spraying it with a  membrane curing compound.    Autogenous Shrinkage    In a set concrete, as hydration proceeds, a net decrease in volume occurs since the hydrated cement 
 gel has a smaller volume than the sum of the cement and water constituents. As hydration continues  in an environment where the water content is constant, such as inside a large mass of concrete, this  decrease in volume of the cement paste results in shrinkage of the concrete. This is known as  autogenous shrinkage because, as the name implies, 
it is self-produced by the hydration of cement.  However, when concrete is cured under water, the water taken up by cement during hydration is  replaced from outside and furthermore the gel particles absorb more water, thus producing a net  increase in volume of the cement paste and an expansion of the concrete. On the other hand if  concrete is k
ept in a dry atmosphere water is drawn out of the hydrated gel and additional shrinkage,  known as drying shrinkage, occurs.   Several factors influence the rate and magnitude of autogenous shrinkage. These include the  chemical composition of cement, the initial water content, temperature and time. The autogenous  shrinkage can be up to 100 x 10'
-6 of which 75 per cent occurs within the first three months.    Drying Shrinkage    When a hardened concrete, cured in water, is allowed to dry it first loses water from its voids and  capillary pores and only starts to shrink during further drying when water is drawn out of its cement  gel. This is known as drying shrinkage and in some concretes
 it can be greater than 1500 x 10'-6, but  a value in excess of 800 x 10'-6 is usually considered to be undesirable for most structural applications.  After an initial high rate of drying shrinkage concrete continues to shrink for a long period of time but  at a continuously decreasing rate (see figure 14.19). For practical purposes, it may be ass
umed that for  small sections 50 per cent of the total shrinkage occurs in the first year.    Figure 14.19 Drying shrinkage and expansion characteristics of concrete     When concrete which has been allowed to dry out is subjected to a moist  environment, it swells. However, the magnitude of this expansion is not sufficient to recover all the  ini
tial shrinkage even after prolonged immersion in water. Concrete subjected to cyclic drying and  wetting approaches the same shrinkage level as that caused by complete drying (figure 14.19). A test  procedure for determining shrinkage is described in BS 1881: Part 5.    Factors affecting Drying Shrinkage    Several factors influence the over-all d
rying shrinkage of concrete. These include the type, content  and proportion of the constituent materials of concrete, the size and shape of the concrete structure,  the amount and distribution of reinforcement and the relative humidity of the environment.   In general, drying shrinkage is directly proportional to the water - cement ratio  and inv
ersely proportional to the aggregate - cement ratio (see figure 14.20). Because of the interaction  of the effects of aggregate - cement and water - cement ratios, it is possible to have a rich mix with a  low water - cement ratio giving higher shrinkage than a leaner mix with a higher water - cement ratio  (Dhir et al., 1978). For a given water -
 cement ratio shrinkage increases with increasing cement  content.    Figure 14.20 Effect of water - cement and aggregate - cement ratios on drying shrinkage of  concrete at 20 C and 50 per cent relative humidity, based on Lea (1970)      Since the aggregate exerts a restraining influence on shrinkage the maximum aggregate  content compatible with
 other required properties is desirable. When the aggregate itself is susceptible  to large moisture movement, this can aggravate shrinkage (or swelling) of the concrete (BRE Digest  35, 1971, and Dhir et al., 1978) and may result in excessive cracking and large deflections of beams  and slabs. The composition and fineness of cement can also affec
t	its shrinkage characteristics. In  general, shrinkage increases as the specific surface area of cement increases (table 14.5) although this  effect is slight and is usually overshadowed by the effects of water - cement ratio and  aggregate - cement ratio. Increases in dicalcium silicate (C2S) content and ignition loss usually result in  increase
d	shrinkage. Tricalcium aluminate (C3A) appears to influence the expansion of concrete  under moist conditions. Nevertheless, the shrinkage characteristics of concrete cannot reliably be  predicted from an analysis of the chemical composition of its cement. In general, admixtures which  reduce the water requirement of concrete without affecting it
s	other properties will reduce its  shrinkage. Air-entrainment itself has no significant influence on shrinkage. Calcium chloride may  considerably increase shrinkage.    TABLE 14.5 Influence of the fineness of cement on drying shrinkage of concrete  (aggregate - cement ratio 3) after 500 days, Bennett and Loat (1970)     The size and shape of a s
pecimen affects the rate of moisture movement in concrete and this  in turn influences the rate of volume change. Since drying begins from the surface, it follows that the  greater the surface area per unit mass, the greater the rate of shrinkage. For a given shape, the initial  rate of shrinkage is greater for small specimens although there will 
be little difference in the ultimate  drying shrinkage, if this stage is ever reached for very large masses of concrete.   The shrinkage of reinforced concrete is less than that of plain concrete owing  to the restraint developed by the reinforcement. This restraint induces tensile stresses in the concrete  which may be large enough to cause crack
ing.   The relative humidity and temperature of the environment have a significant effect on both  the rate and magnitude of shrinkage in as much as they affect the movement of water in concrete. The  duration of initial moist curing has little effect on ultimate shrinkage although it affects the initial rate  of shrinkage.      14.6 Evaluation of
 the Quality of Concrete From Nondestructive Testing    The quality of concrete is usually taken to mean its strength and durability although other properties  such as resistance to deformation and shrinkage can be significant in determining structural  behaviour. In general most of the properties of concrete improve with increasing strength and f
or this  reason the quality of concrete is often judged by its strength. Nondestructive testing, as the name  implies, requires that the material under test is not damaged during testing.   Direct measurement of the strength of concrete involves destructive stresses and thus cannot  be used for determining the quality of concrete in structures. Fu
rthermore, the compressive strength  test, as described in BS 1881: Part 116 can only indicate the potential strength of the mix. The actual  concrete strength within a concrete unit or structure depends on the conditions of placing, compaction  and curing. Although test samples can be cored from the structure and tested for evaluating the  qualit
y	of concrete, this operation is too expensive for general use. Moreover, only a limited number  of cores can be taken without damaging the structure.   It is for these reasons that attempts have been made, in the last four decades, to determine  some suitable nondestructive test for determining the quality of concrete. Several tests have now been
	developed (BS 1881: Part 201) but those which have been most widely accepted include vibrational  methods for estimating strength, durability and uniformity and for detecting flaws, and hardness  methods for estimating strength. Although the tests are simple to perform they have certain limitations.  Nevertheless, when applied rationally the tec
hniques often provide information which cannot otherwise  be obtained by direct methods. The vibrational methods can also be beneficially employed in laboratory  investigations where progressive changes in the quality of concrete due to environmental effects are to  be evaluated.    Resonance Method    This method is used to determine the dynamic 
modulus of elasticity. A concrete specimen of a  well-defined shape, such as a beam similar to the one used in the flexural test, is subjected to vibration  in either the longitudinal, flexural or torsional mode. Basically the experimental technique is similar in  each case and the only difference is the way in which the beam is supported and exci
ted. For the  longitudinal mode of vibration, as described in BS 1881: Part 5, the beam is clamped at its centre and  the exciter and pick-up units are brought into contact with the ends of the beam, without exerting  restraint, so that the ends are free to vibrate in the longitudinal direction (see figure 14.21). The  exciter unit is driven by a 
variable-frequency oscillator and forces the specimen into longitudinal  vibration. These vibrations are received by the pick-up unit and after amplification their amplitude is  indicated on a meter.    Figure 14.21 A typical arrangement for measuring the longitudinal resonance of a concrete beam     During the test the frequency of the oscillator
 is varied so that resonance is  obtained at the fundamental frequency, indicated by maximum deflection on the meter. The dynamic  modulus of elasticity Ed is given by    (formula)   where n is the natural frequency of the fundamental mode of longitudinal vibration of the specimen  (Hz), l the length of the specimen (mm) and p its density (kg m-3)
.	The value of the constant F  depends on the shape of the specimen and the mode of vibration.   In general, changes in the quality of concrete are related to changes in the  dynamic modulus of elasticity but these relationships are not unique and are affected by several factors  such as the constituent materials, particularly the type of aggregat
e, mix proportions and curing  conditions. However, for a given concrete the variation in dynamic modulus of elasticity can give a  good indication of variations in strength and static modulus of elasticity, and this test is particularly  useful for assessing the progressive change in strength and durability as affected by various factors  such as
 the action of alternate freezing and thawing and sulphate solutions. The main disadvantage of  the resonance method is that it cannot be useful for assessing the properties of concrete in an actual  structure since it requires specimens of shapes for which relationships between frequency and the  dynamic modulus of elasticity are known.    Ultras
onic Pulse Method    In this method the velocity of an ultrasonic pulse passing through the concrete is  determined (BS 1881: Part 203). This technique is now widely used for assessing the  quality of concrete in structures and several models of the apparatus are commercially  available. The pulse, produced by an electro-acoustical transducer plac
ed in contact with  the concrete under test, passes through the concrete and is picked up and converted into  an electrical signal by a second electro-acoustical transducer (figure 14.22). The time taken  by the pulse to travel through the concrete is measured by an electrical timing  unit, to the nearest 0.1 us, and the pulse velocity v is calcul
ated using the relationship    (formula)   where L is the length of the path (m) and t is the time taken (s). The accuracy of the  velocity thus  obtained depends on the length of the path although after a certain length the sharpness of  the signal decreases and there is no further gain in the accuracy of the measurement.  Concrete thicknesses be
tween 0.1 and 15 m can be tested quite satisfactorily.    Figure 14.22 Ultrasonic pulse apparatus     The main advantages of the ultrasonic pulse method are that it can be employed  for assessing the quality of concrete in structures and there are no restrictions concerning  the shape of the concrete mass although access to the structure from both
 sides is  desirable. As for the dynamic modulus of elasticity, there is no unique relationship between  the pulse velocity and strength as it is influenced by the concrete constituents and curing  conditions. The effect of the coarse aggregate is of particular  significance since its influence on the pulse velocity is more marked than its influen
ce on  strength. Thus the evaluation of the quality of concrete in structures is usually made on a  comparative basis and the technique is frequently employed to detect inferior parts within a  structure (Tomsett, 1980). However, when the mix proportions remain constant and only  one type of coarse aggregate is used then it is possible to determin
e a specific relationship  between strength and pulse velocity for in situ concrete. Since the pulse velocity is affected  by moisture it is important to have moisture conditions in  the test specimens similar to those in the in situ concrete when establishing  strength - pulse-velocity relationships. The pulse is transmitted most effectively by s
olid  media and cracks or cavities are indicated by a reduction in the pulse velocity. The method  is useful for a continuous assessment of the effects of deteriorating agencies such as frost  and chemicals for both in situ and laboratory concrete. It  can also be used to obtain values for the dynamic modulus of elasticity of concrete using  the r
elationship    (formula).   Hardness Method    The Schmidt rebound hammer is the most widely accepted instrument for measuring the  surface hardness of concrete, as described in BS 1881: Part 202. A fixed mass of steel is  charged with kinetic energy through a spring system by gradually pressing the plunger  against the surface to be tested (figur
e 14.23). The steel mass is released and impinges on  the plunger, which remains in contact with concrete. After this impact the mass rebounds  and the magnitude of the rebound is a measure of the hardness of the surface, indicated by  a rider on a linear scale graduated in empirical rebound  numbers. When the hammer is used to test a nonvertical 
surface the rebound reading is  corrected because of the change in the impact energy.    Figure 14.23 Schmidt hammer     Hardness (rebound number) is a relative property and there can be no physical  relationship between it and the other properties of concrete. Empirical relationships  between rebound number and strength have been established and 
in general the higher the  rebound number, the greater the strength. As for the dynamic modulus of elasticity and  ultrasonic pulse velocity, there is no unique relationship between rebound number and  strength. For this reason it is advisable to determine the  strength - hardness relationship for each concrete instead of relying on secondhand val
ues.  When the hammer is used for assessing the strength of in situ concrete the test procedure  and environmental conditions should be similar to those employed during calibration. Since  for a given concrete the rebound number can vary because of differences in hardness  between aggregate and matrix and the possible variation in aggregate minera
logy, it is  necessary that several readings are taken and the average value used. Control over the  preparation of the test surface is important for proper use of the hammer. Provided the  limitations of the method are borne in mind and the hammer is used intelligently, it can be a  useful tool for assessing the strength of concrete in structures
. It is an inexpensive  mechanical device and is easy to use. The technique can also be applied to assess the  uniformity of concrete, for example, to locate the possible existence of an area of  unsatisfactory concrete in a wall, or to detect debonded areas in floor screeds (Chaplin,  1980).    Other Methods    These include the use of gamma rays
 for detecting voids in concrete and for locating  reinforcement, electrical methods for measuring moisture content and electromagnetic  methods for measuring the depth below the concrete surface of steel reinforcing bars.  Some of these are described in BS 4408: Part 1 and Part 3. Recently developed tests which  are of interest for the in situ as
sessment of concrete strength are pull-out, break-off and  penetration tests. The application of these tests, along with  ultrasonic pulse and hardness tests, is dealt with in the British Standard Guide to  Assessment of Concrete Strength in Existing Structures (BS 6089). The pull-out tests  essentially involve determining the force required to pu
ll out a steel rod fixed into the  surface of the concrete (Malhotra, 1975;  Chabowski and Brydon-Smith, 1980). The break-off test measures the flexural strength of  concrete (Johansen, 1979). The penetration test measures the resistance of concrete to  penetration of a probe fixed into the surface (ASTM C 803-75). Introduction    Concrete is a ma
n-made composite the major constituent of which is natural aggregate,  such as gravel and sand or crushed rock. Alternatively artificial aggregates, for example,  blastfurnace slag, expanded clay, broken brick and steel shot may be used where  appropriate. The other principal constituent of  concrete is the binding medium used to bind the aggregat
e particles together to form a  hard composite material. The most commonly used binding medium is the product formed  by a chemical reaction between cement and water. Other binding mediums are used on a  much smaller scale for special concretes in which the cement and water of normal  concretes are replaced either wholly or in part by epoxide or p
olyester resins. These  polymer concretes known as resin-based or resin-additive concretes  respectively are costly and generally not suitable for use where fire-resistant properties are  required but they are useful for repair work and other special applications. Resin-based  concretes have been used, for example, for precast chemical resistant p
ipes and lightweight  drainage channels. This section deals only with normal concretes in which cement and  water form the binding medium.   In its hardened state concrete is a rock-like material with a high compressive  strength. By virtue of the ease with which fresh concrete in its plastic state may be  moulded into virtually any shape it may b
e used to advantage architecturally or solely for  decorative purposes.  Special surface finishes, for example, exposed aggregate, can also be used to great effect.   Normal concrete has a comparatively low tensile strength and for structural  applications it is normal practice either to incorporate steel bars to resist any tensile forces  (reinfo
rced concrete) or to apply compressive forces to the concrete to counteract these  tensile forces (prestressed concrete). Concrete is also used in conjunction with other  materials, for example, it may form the compression flange of a box section the remainder  of which is steel (composite construction).  Concrete is used structurally in buildings
 for foundations, columns, beams and slabs, in  shell structures, bridges, sewage-treatment works, railway sleepers, roads, cooling towers,  dams, chimneys, harbours, off-shore structures, coastal protection works and so on. It is  used also for a wide range of precast concrete products which includes concrete blocks,  cladding panels, pipes and l
amp standards.   The impact strength, as well as the tensile strength, of normal concretes is low  and this can be improved by the introduction of randomly orientated fibres into the  concrete mix. Steel, polypropylene, asbestos and glass fibres have all been used with some  success in precast products, for example, pipes, building panels and pile
s. Steel fibres also  increase the flexural strength, or modulus of rupture, of concrete and this particular type of  fibre-reinforced concrete has been used in ground paving slabs for roads where flexural  and impact strength are both important. Fibre-reinforced  concretes are however essentially special-purpose concretes and for most purposes th
e  normal concretes described in this book are used.   In addition to its potential from aesthetic considerations, concrete requires little  maintenance and has good fire resistance. Concrete has other properties which may on  occasions be considered less  desirable, for example, the time-dependent deformations associated with drying shrinkage  an
d other related phenomena. However, if the effects of environmental conditions, creep,  shrinkage and loading on the dimensional changes of concrete structures and structural  elements are fully appreciated, and  catered for at the design stage, no subsequent difficulties in this respect should arise.   A true appreciation of the relevant properti
es of any material is necessary if a  satisfactory end product is to be obtained and concrete, in this respect, is no different from  other materials. 2 Cements of Different Types      The previous chapter dealt with the properties of Portland cement in general, and we have  seen that cements differing in chemical composition and physical characte
ristics may  exhibit different properties when hydrated. It should thus be possible to select mixtures of  raw materials for the production of cements with various desired properties. In fact,  several types of Portland cement are available commercially and additional special cements  can be produced for specific uses. The various types of Portlan
d cement will now be  described. Several non-Portland cements will also be discussed.       Types of Portland Cement     In order to facilitate the discussion, a list of different Portland cements, together with the  American description where available, is given in Table  2.1. The ASTM composition  limits for some of these cements have already be
en listed (Table 1.9), and typical values of  compound composition are given in Table 2.2.    Table 2.1: Main Types of Portland Cement    Table 2.2: Typical Values of Compound Composition of Portland Cements of  Different  Types     Many of the cements have been developed to ensure good durability of concrete under a  variety of conditions. It has
 not been possible, however, to find in the constitution of  cement a complete answer to the problem of durability of concrete: the principle  mechanical properties of hardened concrete, such as strength, shrinkage, permeability,  resistance to weathering, and creep, are affected also by factors other than cement  constitution, although this deter
mines to a large degree the rate of gain of strength. Figure  2.1 shows the rate of development of strength of concretes made with cements of  different types: while the rates vary considerably, there is little difference in the 90-day  strength of cements of all types. The general tendency is for the cements with a low rate of  hardening to have 
a slightly higher ultimate strength. For instance, Fig. 2.1 shows that  Type IV cement has the lowest strength at 28 days but develops the second highest  strength at the age of 5 years. A comparison of Fig. 2.1 and Fig. 2.2 illustrates the fact that  differences between cement types are not readily quantified.    Fig. 2.1. Strength development of
 concretes containing 335 kg of cement per cubic metre  (565 lb / yd3) and made with cements of different types    Fig. 2.2. Strength development of concretes with a water / cement ratio of 0.49 made with  cements of different types     Also, the retrogression of strength of the concrete made with Type II cement is not  characteristic of this type
 of cement. The pattern of low early and high late strength agrees  with the influence of the initial framework of hardened cement on the ultimate  development of strength: the more slowly the framework is established the denser the gel  and the higher the ultimate strength. Nevertheless, significant differences in the important  physical properti
es of cements of different types are found only  in the earlier stages of hydration in well-hydrated pastes the differences are only minor.   The division of cements into different types is necessarily no more than a broad  classification and there may sometimes be wide differences between cements of nominally  the same type. On the other hand, th
ere are often no sharp discontinuities in the properties of different types of cement, and many cements can be classified as more than one type.   Obtaining some special property of cement may lead to undesirable features in another  respect. For this reason a balance of requirements may be necessary, and the economic  aspect of manufacture must a
lso be considered. Type II cement is an example of a  "compromise" all-round cement .   The methods of manufacture have improved steadily over the years, and there has been a  continual development of cements to serve different purposes with a corresponding  change in specifications.      Ordinary Portland Cement     This is by far the most common
 cement in use: about 90 per cent of all cement used in the  United States (total production of about 24 million tonnes per annum) and a like  percentage in the  United Kingdom (about 17 million tonnes per annum) is of the ordinary type.   Ordinary Portland cement (Type I) is admirably suitable for use in general concrete  construction when there 
is no exposure to sulphates in the soil or in ground water. The specification for this cement is given in BS 12: 1978. The limitations of chemical  composition are: the lime saturation factor is to be not greater than 1.02 and not less than  0.66. The factor is defined as -    (formula),    where each term in brackets denotes the percentage by wei
ght of the given compound  present in the cement.   The upper limit of the lime saturation factor ensures that the amount of  lime is not so  high as to result in free lime appearing at the clinkering temperature in equilibrium with the  liquid present. The unsoundness of cement caused by free lime was discussed in the  previous chapter, and is in
deed controlled by the Le Chatelier test. But the importance of  avoiding unsound cement is so great that in Britain the safeguard of controlled compound  composition is considered desirable. Nevertheless, the ASTM Standard and the majority of  European specifications for cement prescribe no limits of the lime content.  A further requirement of BS
 12: 1978 for the chemical composition of ordinary Portland  cement is that the magnesia content does not exceed 4.0 per cent. Formerly, the ratio  Al2O3 / Fe2O3 was specified as not less than 0.66. In addition, the insoluble residue must  not exceed 1.5 per cent and the loss on ignition is limited to 3 per cent in temperate  climates and 4 per ce
nt in the tropics. The maximum gypsum content is also specified (see  p 18).   Over the years, there have been some changes in the characteristics of ordinary Portland  cement. In particular, modern cements have a higher C3S content and a greater fineness  than 40 years ago. As a consequence, cements have nowadays a 28-day strength perhaps  25 MPa
 higher than in 1925, but the gain in strength between 28 days and 10 years is  unaltered: approximately 20 MPa (3000 psi) for continuously water-cured concrete with a  water / cement ratio of about 0.53. (See Fig. 2.3.)    Fig. 2.3. The rate of gain of strength of cements between 1916 and 1970 measured on  standard concrete cylinders with a water
 / cement ratio of 0.53     The German classification of cements is on the basis of the 28-day strength of 1 : 3  mortars with a water / cement ratio of 0.5: 35, 45, and 55 MPa.      Rapid Hardening Portland Cement     This cement is very similar to ordinary Portland cement, and is also covered by BS 12:  1978. Rapid hardening Portland cement (Typ
e III), as its name implies, develops strength  more rapidly, and should therefore be correctly described as high early strength cement.  The rate of hardening must not be confused with the rate of setting: in fact, the two  cements have similar setting times.   The strength developed by the rapid hardening Portland cement at the age of three days
 is  of the same order as the 7-day strength of ordinary Portland cement with the same  water / cement ratio but the British Standards no longer specify 7-day strength. The  expected rate of hardening is reflected in the minimum strengths specified by BS 12: 1978,  listed in Table 1.10. The increased rate of gain of strength of the rapid hardening
 cement is  achieved by a higher C3S content, sometimes as high as 70 per cent, and by finer grinding  of the cement clinker. BS 12:1978 prescribes a minimum fineness of 325m2 / kg, but as a  rule a higher fineness is encountered.   Ordinary Portland cement manufactured in Great Britain is invariably finer than the 225  m2 / kg prescribed by BS 12
:197 - often above 300m2 / kg. Many plants also produce  cement with a high C3S content so that sometimes in practice there is little difference  between rapid hardening and some ordinary cements; this cannot, however, be assumed to  be the rule (see p. 333).   The requirements of soundness and chemical composition are the same for rapid hardening
 as for  ordinary Portland cement and need not therefore be repeated.   The use of rapid hardening cement is indicated where a rapid strength development is desired,  e.g. when formwork is to be removed early for re-use, or where sufficient strength for further  construction is wanted as quickly as practicable. Rapid hardening cement is only about
 $4 (f2) per  tonne dearer than ordinary cement, and it is not surprising that rapid hardening cement is used  extensively, accounting for about 10 per cent of all cement manufactured in the United Kingdom. Since,  however, the rapid gain of strength means a high rate of heat development, rapid hardening Portland  cement should not be used in mass
 construction or in large structural sections. On the other hand, for  construction at low temperatures the use of cement with a high rate of heat liberation may prove a  satisfactory safeguard against early frost damage.      Special Rapid Hardening Portland Cements    There exist several specially manufactured cements which are particularly rapi
d hardening.  One of these, a so-called extra rapid hardening Portland cement, is obtained by intergrinding calcium  chloride with rapid hardening Portland cement. The quantity of calcium chloride should not exceed 2  per cent. Because calcium chloride is deliquescent it is vital to store extra rapid hardening cement under  dry conditions, and it 
should generally be used within one month of despatch from the cement plant.   Extra rapid hardening cement is suitable for cold weather concreting, or when a very high  early strength is required, but structural use with reinforcement is not permitted by the British Code of  Practice CP 110: 1972 because of the risk of corrosion, and the cement i
s no longer manufactured in the  United Kingdom. The strength of extra rapid hardening cement is about 25 per cent higher than that of  rapid hardening cement at 1 or 2 days and 10 to 20 per cent higher at 7 days. The setting time of extra  rapid hardening cement is short: depending on temperature it can be 5 to 30 minutes so that early  placing i
s essential. Shrinkage is rather higher than when rapid hardening Portland cement is used.   If extra rapid hardening cement is not available it is possible to use rapid hardening  Portland cement whose speed of hardening is increased by means of an addition of calcium chloride  immediately prior to mixing the concrete. In this case, calcium chlor
ide can be classified as an  accelerator; its effects are discussed on p. 101.   Another type of cement with very rapid hardening properties is the so-called ultra high  early strength Portland cement, marketed in Great Britain. This cement contains no admixture and is  therefore suitable for use in reinforced and prestressed concrete; the rapid s
trength development is due  to the very high fineness of the cement: 700 to 900 m2 / kg. Because of this, the gypsum content has to  be higher (4 per cent expressed as SO3) than in cements complying with BS 12: 1978, but in all other  respects the ultra high early strength cement satisfies the requirement of that standard. We may note  that the hi
gh gypsum content has no adverse effect on long-term soundness as the gypsum is used up  in the early reactions of hydration.   The cement is manufactured by separating fines from rapid hardening Portland cement by  a cyclone air elutriator. Because of its high fineness, the ultra high early strength cement has a low bulk  density and deteriorates
 rapidly on exposure. High fineness leads to rapid hydration, and therefore to a  high rate of heat generation at early ages and to a rapid strength development; for instance, the 3-day  strength of rapid hardening Portland cement is reached at 16 hours, and the 7-day strength at 24  hours. There is, however, little gain in strength beyond 28 days
. Typical strengths of 1 : 3 concretes  made with the ultra high early strength cement are given in Table 2.3.    Table 2.3: Typical Values of Strength of a 1 : 3 Concrete made with Ultra High Early Strength  Portland Cement     The cement has been used successfully in a number of structures where early prestressing  or putting in service is of im
portance. Shrinkage and creep are not significantly different from those  obtained with other cements when the mix proportions are the same; in the case of creep, the  comparison has to be made on the basis of the same stress / strength ratio (see p. 401). We should note,  however, that for the same mix proportions, the use of ultra high early str
ength cement results in lower  workability.   The ultra high early strength Portland cement is marketed as Swiftcrete. A somewhat less  fine cement is Speed cement, developed in Belgium. It contains no accelerator and has a specific surface  of 450 to 500 m2 / kg. The standard vibrated mortar cube test gives strengths of about 28 MPa (4000  psi) a
t 1 day, 48 MPa (7000 psi) at 3 days, and 68 MPa (9800 psi) at 28 days. The Speed cement is  suitable for winter concreting or for urgent jobs such as road repair, well-sealing, etc.   In some countries, e.g. Italy and Sweden, extremely high early strength cement is  manufactured by double burning in the kiln.   One more cement of the very high ea
rly strength cement variety should be mentioned.  This is the so-called regulated-set cement, or jet cement, developed in the U.S. The cement consists  essentially of a mixture of Portland cement and calcium fluoro-aluminate (C11A7.CaF2) with an  appropriate retarder (usually citric acid). The setting time of the cement can vary between 1 and 30  
minutes (the strength development being slower the slower the setting) and  is controlled in the manufacture of the cement as the raw materials are interground and burnt  together. Grinding is difficult because of hardness differences.   The early strength development is controlled by the content of calcium fluoro-aluminate:  when this is 5 per ce
nt, about 6 MPa (900 psi) can be achieved at 1 hour; a 50 per cent mixture will  produce 20 MPa (3000 psi) at the same time. These values are based on a mix with a cement content of  330 kg / m3 (560 lb / yd3). The later strength development is similar to that of the parent Portland  cement but at room temperature there is virtually no gain in str
ength between 1 and 3 days.   Regulated-set cement is manufactured in Japan, Austria and Germany, where it is known  as Schnellzement.      Low Heat Portland Cement    The rise in temperature in the interior of a large concrete mass due to the heat developed by the  hydration of cement can lead to serious cracking (see p. 424). For this reason, it
 is necessary to limit  the rate of heat evolution of the cement used in this type of structure: a greater proportion of the heat  can then be dissipated and a lower rise in temperature results.   Cement having such a low rate of heat development was first produced for use in large gravity  dams in the United States, and is known as low heat Portl
and cement (Type IV).   BS 1370 : 1979 limits the heat of hydration of this cement to 250 J per gram (60 cal / g) at the age  of 7 days, and 290 J per gram (70 cal / g) at 28 days.   The limits of lime content of low heat Portland cement, after correction for the lime combined  with SO3, are -    (formula).  The rather lower content of the more ra
pidly hydrating compounds, C3S and C3A, results in a  slower development of strength of low heat cement as compared with ordinary Portland cement,  but the ultimate strength is unaffected. In any case, to ensure a sufficient rate of gain of strength the  specific surface of the cement must be not less than 320 m2 / kg.   A low heat Portland blast-
furnace cement is also available; this is covered by BS 4246 : 1974. In  the United States, Portland-pozzolana cement Type P can be specified to be of the low heat  variety; the Type IP Portland-pozzolana cement can be required to have moderate heat of hydration.  ASTM Standard C 595-79 deals with these cements.   In some applications a very low e
arly strength may be a disadvantage, and for this reason a  so-called modified (Type II) cement was developed in the United States. This modified cement  successfully combines a somewhat higher rate of heat development than that of low heat cement with a  rate of gain of strength similar to that of ordinary Portland cement. Modified cement is reco
mmended  for use in structures where a moderately low heat generation is desirable or where moderate sulphate  attack may occur. This cement is extensively used in the United States.   Modified cement, referred to as Type II cement, and low heat cement (Type IV) are covered by  ASTM Specification C 150 -78a.      Sulphate-resisting Cement    In di
scussing the reactions of hydration of cement, and in particular the setting process,  mention was made of the reaction between C3A and gypsum (CaSO4.2H2O) and of the consequent  formation of calcium sulphoaluminate. In hardened cement, calcium aluminate hydrate can react with a  sulphate salt from outside the concrete in a similar manner: the pro
duct of addition is calcium  sulphoaluminate, forming within the framework of the hydrated cement paste. Since the increase in the  volume of the solid phase is 227 per cent, gradual disintegration of concrete results. A second type of  reaction is that of base exchange between calcium hydroxide and the sulphates, resulting in the  formation of gy
psum with an increase in the volume of the solid phase of 124 per cent.   These reactions are known as sulphate attack. The salts particularly active are magnesium  and sodium sulphate. Sulphate attack is greatly accelerated if accompanied by  alternate wetting and drying, as is the case for instance, in a marine structure in the zone  between the
 tides (see Chapter 7).   The remedy lies in the use of cement with a low C3A content, and such cement is known  as sulphate-resisting Portland cement. The British Standard for this cement, BS 4027 : 1980,  stipulates a C3A content of 3.5 per cent. The minimum fineness is 250 m2 / kg. In other respects,  sulphate-resisting cement is expected to co
nform to BS 12 : 1978 for ordinary Portland cement. In the  United States, sulphate-resisting cement is known as Type V cement and is covered by ASTM  Standard C 150-78a. This specification limits the C3A content to 5 per cent, and also restricts the total  content of C4AF plus twice the C3A content to 20 per cent. The magnesia content is limited 
to 6 per  cent.   The role of C4AF is not quite clear. From the chemical standpoint C4AF would be expected to  form calcium sulphoaluminate, as well as calcium sulphoferrite , and thus cause expansion. It seems,  however, that the action of calcium sulphate on hydrated cement is smaller the lower the  Al2O3 : Fe2O3 ratio. Some solid solutions are 
formed and they are liable to comparatively little attack.  The tetracalcium ferrite is even more resistant, and it may form a protective film over any free calcium  aluminate.   Since it is often not feasible to reduce the Al2O3 content of the raw  material, Fe2O3 may be  added to the mix so that the C4AF content  increases at the expense of C3A.
	An example of a cement with a very low Al2O3 : Fe2O3 ratio is the Ferrari cement, in whose  manufacture iron oxide is substituted for some of the clay. A similar cement is produced in Germany  under the name of Erz cement. The name of iron ore cement is also used for this type of cement.  The low C3A and comparatively low C4AF contents of sulph
ate-resisting cement mean that it has  a high silicate content and this gives the cement a high strength but, because C2S represents a high  proportion of the silicates, the early strength is low. The heat developed by sulphate-resisting cement is  not much higher than that of low heat cement. It could therefore be argued that sulphate-resisting  
cement is theoretically an ideal cement, but because of the special requirements for the composition of  the raw materials used in its manufacture, sulphate-resisting cement cannot be generally and cheaply  made.   Provision for a low-alkali sulphate-resisting cement is made in BS 4027 : 1980.      Portland Blast-furnace Cement    This type of cem
ent is made by intergrinding Portland cement clinker and granulated blast-furnace  slag, the proportion of the latter not exceeding 65 per cent of the weight of the mixture, as  prescribed by BS 146 : 1973.  Slag is a waste product in the manufacture of pig iron, the amounts of iron and slag obtained  being of the same order. The slag is a mixture
 of lime, silica, and alumina, that is the same oxides that  make up Portland cement, but not in the same proportions. While it is not possible to give ranges of  values it may be noted that a slag known to be satisfactory had the following composition: 42 per cent  lime, 30 per cent silica, 19 per cent alumina, 5 per cent magnesia, and 1 per cent
 alkalis.   Blast-furnace slag varies greatly in composition and physical structure depending on the  processes used and on the method of cooling of the slag. For use in the manufacture of blast-furnace  cement the slag has to be quenched so that it solidifies as a glass, crystallization being largely prevented.  This rapid cooling by water result
s also in a fragmentation of the material into a granulated form.   Slag can make a cementitious material in different ways. Firstly, it can be used together with  limestone as a raw material for the conventional manufacture of portland cement. Clinker made from  these materials is often used (together with slag) in the manufacture of portland bla
st-furnace cement.   The latter cement represents the second major use of slag. Dry  granulated slag is fed with  Portland cement clinker into a grinding mill, gypsum being added in order to control setting. It may be  noted that slag is harder than clinker so that intergrinding presents some difficulties. Portland  blast-furnace cement has been m
anufactured in Scotland for a number years, and is also made in the  United States (by intergrinding or by blending) where it is known as Type IS  cement, and is covered by ASTM C 595-79. Portland blast-furnace cement is used also in  Germany, under the name of Eisenportland (up to 35 per cent slag) and Hochofen cements (36 to 85 per  cent slag), 
and in France where the most common ones are ciment metallurgique mixte (50 per cent  slag) and ciment de haut fourneau (65 to 75 per cent slag). In the Netherlands, the slag content may be  as high as 85 per cent.   A Belgian development is the Trief process in which wet-ground granulated slag is fed in the form of a  slurry direct into the concr
ete mixer, together with Portland cement and aggregate. The cost of drying  the slag is thus avoided, and grinding in the wet state results in a greater fineness than would be  obtained with dry grinding for the same power input.   A variant used in Britain under the name of Cemsave and in South Africa known as Slagment is a  process where dry-gro
und granulated slag of the same fineness as cement is added at the mixer as  a partial replacement of Portland cement; Portland blast-furnace cement concrete is thus  manufactured in situ. Like Portland blast-furnace cement concrete, Cemsave concrete has a lower early  strength than when Portland cement only is used, but at later ages at least equ
al strengths are reached  However, with Cemsave the workability is somewhat higher so that some reduction in water / (cement  plus slag) ratio is possible compared with the water / cement ratio of a Portland cement mix with the  same aggregate content. Concrete  with Cemsave exhibits a significantly  smaller temperature rise during hydration. Also
, the coefficient of thermal expansion is  reduced by some 10 per cent compared with a similar Portland cement mix, possibly due to a lower  moisture content.   The exact nature of hydration of Portland blast-furnace cement is not quite clear. The  Portland cement component hydrates in the normal manner and it appears that the calcium hydroxide  t
hus liberated gives the correct alkalinity needed to provide a "starter" for the hydration of the  granulated slag. However, the further hydration of the slag is direct and does not depend on  combination with lime.   Portland blast-furnace cement is in many respects similar to ordinary Portland cement,  and BS 146 : 1973 requirements for fineness
, setting  times and soundness are the same for both  cements. In actual fact, the fineness of Portland blast-furnace cement tends to be higher, but even so the  rate of hardening of Portland blast-furnace cement is somewhat slower during the first 28 days, and  adequate curing is therefore of importance; the strength requirements of BS 146 : 1973
 are therefore  lower than for ordinary Portland cement the requirement for a 28-day strength is 34 MPa  (4900 psi) for  mortar cubes or 22 MPa (3200 psi) for concrete cubes. However, at later ages  there is little difference between the strengths of  Portland blast-furnace and ordinary  Portland cements. Fig. 2.4 shows typical strength - time cur
ves.    Fig. 2.4. Strength development of concretes made with Portland blast-furnace cement  (water / cement ratio = 0.6)     The heat of hydration of Portland blast-furnace cement is lower than  that of ordinary Portland cement so that the former can be used in mass concrete  structures. (The cement should then comply with BS 4246: 1974, which co
vers low-heat Portland  blast-furnace cement; this allows a longer final setting time and a lower strength than BS 1370:1979.)  However, in cold weather the low heat of hydration of Portland blast-furnace cement, coupled with a  moderately low rate of strength development, can lead to frost damage.  Because of its fairly high sulphate resistance, 
the C3A content being low, Portland blast-furnace  cement is frequently used in sea-water construction. Shrinkage and modulus of elasticity of concrete  made with Portland blast-furnace cement are the same as when Portland cement is used. Creep is also  unaffected, except that it is larger under drying conditions.   The relatively low energy requi
rement in the manufacture of Portland blast-furnace cement may be of  interest in these energy-conscious times, and this cement is already extensively used in many  countries which have a large production of slag. For instance, in the Netherlands, 60 per cent of all  cement used is of the Portland blast-furnace type.      Supersulphated Cement    
Because it is made from granulated blast-furnace slag, supersulphated cement will be  considered at this stage, even though it is not a Portland cement.   Supersulphated cement is made by intergrinding a mixture of 80 to 85 per cent of granulated  slag with 10 to 15 per cent of calcium sulphate (in the form of dead-burnt gypsum or anhydrite)  and 
about 5 per cent of Portland cement clinker. A fineness of 400 to 500 m2 / kg is usual.  The cement has to be stored under very dry conditions as otherwise it deteriorates rapidly.   Supersulphated cement is used extensively in Belgium, where it is known as ciment  metallurgique sursulfate, in France, and was previously manufactured in Germany (un
der  the name of Sulfathuttenzement). In the United Kingdom, the cement is covered by the British  Standard BS 4248: 1974, but, because of production difficulties, the manufacture of the cement  has been discontinued. The cement is highly resistant to sea water and can withstand the highest  concentrations of sulphates normally found in soil or gr
ound water, and is also resistant to peaty  acids and to oils. Concrete with a water / cement ratio not greater than 0.45 has been found not to  deteriorate in contact with weak solutions of mineral acids of pH down to 3.5. For these reasons,  supersulphated cement is used in the construction of sewers and in contaminated ground, although it  has 
been suggested that this cement is less resistant than sulphate-resisting Portland cement when the  sulphate concentration exceeds 1 per cent.   The heat of hydration of supersulphated cement is low: about 170 to 190 J / g (40 to 45  cal / g) at 7 days, and 190 to 210 J / g (45 to 50 cal / g) at 28 days. The cement is therefore  suitable for mass 
concrete construction but care must be taken if used in cold weather as the rate  of strength development is considerably reduced at low temperatures. The rate of hardening of  supersulphated cement increases with temperature up to about 50 C (122 F), but at higher  temperatures anomalous behaviour has been encountered. For this reason, steam curi
ng above  50 C (122 F) should not be used without prior tests. It may also be noted that supersulphated  cement should not be mixed with Portland cements because the lime released by the hydration of  an excessive amount of the latter interferes with the reaction between the slag and the calcium  sulphate.   Wet curing for not less than four days 
after casting is essential as premature drying  out results in a friable or powdery surface layer, especially in hot weather, but the depth of  this layer does not increase with time.    Table 2.4: Typical Values of Strength of Supersulphated Cement     Supersulphated cement combines chemically with more water than is required for the hydration of
	Portland cement, so that concrete with a water / cement ratio of less than	0.4 should not be made.  Mixes leaner than about 1 : 6 are not recommended. The decrease in strength with an increase in  the water / cement ratio has been reported to be smaller than in other cements but, since the early  strength development depends on the type of slag 
used in the manufacture of the cement, it is  advisable to determine the actual strength characteristics prior to use. Typical strengths  attainable are given in Table 2.4. It should be noted that for the concrete test BS 4248: 1974  prescribes a water / cement ratio of 0.55 instead of 0.60 used with other cements .      Portland-pozzolana Cements
 and Pozzolanas    The first of these is the name given to interground or blended mixtures of Portland cement and  pozzolana.   Pozzolana is a natural or artificial material containing silica in a reactive form. A more formal  definition of ASTM Specification C 618-78 describes pozzolana as a siliceous or siliceous and  aluminous material which in
 itself possesses little or no cementitious value but will, in finely  divided form and in the presence of moisture, chemically react with calcium hydroxide at ordinary  temperatures to form compounds possessing cementitious properties. It is essential that pozzolana  be in a finely divided state as it is only then that silica can combine with cal
cium hydroxide (liberated by  the hydrating Portland cement) in the presence of water to form stable calcium silicates which have  cementitious properties. Pozzolanic materials most commonly met with are: volcanic ash - the original pozzolana - pumicite, opaline shales and cherts, calcined diatomaceous earth, burnt clay, fly ash, etc.  In consider
ing pozzolanas in general, we should note that the silica has to be amorphous, as  crystalline silica has very low reactivity.   Some pozzolanas may create problems because of their physical properties; e.g. diatomaceous  earth, because of its angular and porous form, requires a high water content. Certain natural  pozzolanas improve their activit
y	by calcination in the range of 550 C to 1100 C, depending on  the material. Rice husks burnt at 450 C have been found to produce a pozzolana conforming to the  requirements of the earlier ASTM Standard C 618-72, but their use leads to increased shrinkage.   For an assessment of pozzolanic activity with cement, ASTM Standard C 618-78 prescribes t
he  measurement of a pozzolanic activity index. This is established by the determination of strength of  mixtures with a specified replacement of cement by pozzolana. There is also a pozzolanic activity  index with lime, which determines the total activity of a pozzolana. According to BS 4550: Part 2:1970,  pozzolanicity is assessed by comparing t
he quantity of Ca(OH)2 present in a liquid phase in contact  with the hydrated pozzolanic cement with the quantity of Ca(OH)2 capable of saturating a medium of  the same alkalinity. The underlying principle is that the pozzolanic activity consists of fixing of  Ca(OH)2 by the pozzolana so that the lower the resulting quantity of Ca(OH)2 the higher
 the  pozzolanicity.   Pozzolanicity is still imperfectly understood; specific surface and chemical composition are  known to play an important role but, since they are inter-related, the problem is complex.  It has been suggested that, in addition to reacting with Ca(OH)2, pozzolanas react also with  C3A or its products of hydration. A good revie
w	of the subject of pozzolanicity has been written by  Massazza and Costa.	Fly Ash    Fly ash, known also as pulverised-fuel ash (see p. 609), is the ash precipated electrostatically from  the exhaust fumes of coal-fired power stations, and is the most common artificial pozzolana; an  extensive review has been written by Berry and Malhotra. The
 fly ash particles are spherical  (which is advantageous from the water requirement point of view) and are of at least the same  fineness as cement (although with fewer very fine particles) so that the silica is readily available  for reaction.   The pozzolanic activity of fly ash is in no doubt, but it is essential that it has a constant fineness
	and a constant carbon content. The two are often interdependent since the carbon particles tend to  be coarser. Modern boiler plants produce fly ash with a carbon content of about 3 per cent but  much higher values are encountered in fly ash from older plants. British Standard BS 3892: 1965  lays down a maximum loss on ignition of 7 per cent, bu
t	probably a carbon content up to 12 per  cent is acceptable. It is not clear why carbon may be harmful but, of course, it is not pozzolanic and is  therefore no more than a filler. The importance of uniformity of properties of fly ash cannot be  overestimated, as otherwise it is difficult to maintain the uniformity of the resulting concrete. Brit
ish  Standard BS 3892: 1965 specifies three zones of fineness so that uniform supply can be assured. The  residue on the 45 um sieve is a convenient basis of classification of size. The main requirements of  ASTM Standard C 618-78 are: a minimum content of 70 per cent of silica, alumina and ferric oxide all  together, a maximum SO3 content of 5 pe
r	cent, a maximum loss on ignition of 12 per cent, and a  maximum alkali content (expressed as Na2O) of	1.5 per cent. The latter value is applicable only when  the fly ash is to be used with reactive aggregate. British Standard BS 3892: 1965 specifies a maximum  MgO content of 4 per cent and SO3 of 2 5 per cent.    It should be noted that fly ash 
may affect the colour of the resulting concrete.   Recently, high-lime ashes originating from lignite coal have entered the pozzolana market.  They do not satisfy the existing specifications and may have a lime content as high as 24 per  cent. High-lime ash has hydraulic properties of its own but, since its lime will combine with the  silica and a
lumina portions of the ash, there will be less of  these compounds to react with the  lime liberated by the hydration of cement. The carbon content is low, the fineness is high, and  the colour is light. However, the MgO content can be high, and some of the MgO as well as some of the  lime can lead to deleterious expansion.   The behaviour of high
-lime ashes is sensitive to temperature: specifically, in mass concrete when  a rise in temperature occurs, the products of reaction may not be of high strength. However, the  development of strength is not simply related to temperature, being satisfactory in the  region of 120 to 150 C (250 to 300 F) but not at about 200 C (about 400 F) when the 
products of  reaction are substantially different.   The testing of this type of ash is still being developed but future use is possible.    Use of Pozzolanas    It is not possible to make a generalized statement on the Portland-pozzolana cements because the  rate of strength development depends on the activity of the pozzolanas and on the proport
ion of  Portland cement in the mixture. As a rule, however, Portland-pozzolana cements gain strength  very slowly and require, therefore, curing over a comparatively long period, but their ultimate  strength is approximately the same as that of ordinary Portland cement alone. A typical strength curve  is shown in Fig. 2.5.    Fig. 2.5. Strength de
velopment of concrete made with Portland cement and fly ash     ASTM Standard C 595-79 describes Portland-pozzolana cement as Type IP for general  concrete construction and Type P for use when high strengths at early ages are not required, and  limits the pozzolana content to between 15 and 40 per cent of the weight of the Portland-pozzolana  ceme
nt. In the United Kingdom, Portland-pozzolana cement with up to 35 per cent of  fly ash appeared on the market in	1980.	Pozzolanas may often be cheaper than the Portland cement that they replace but their chief  advantage lies in slow hydration and, therefore, low rate of heat development: this is of great  importance in mass concrete constructi
on, and it is there that Portland-pozzolana cement or a  partial replacement of Portland cement by the pozzolana is mostly used.   Portland-pozzolana cements show also good resistance to sulphate attack and to some other  destructive agents. This is so because the pozzolanic reaction leaves less lime to be leached out  and also reduces that permea
bility of concrete. However, the resistance to freezing and thawing  cannot be developed until later ages when significant pozzolanic reaction has reduced the porosity  of the paste. It should be remembered that pozzolanas vary very considerably in their effects, both  good and bad, and it is advisable to test any untried pozzolanic material in co
mbination with the  cement and the aggregate which are to be used in the actual construction. The use of fly ash with  sulphate-resisting Portland cement is not allowed according to British Code of Practice CP 110: 1972  and BS 5328: 1976 when resistance to sulphate attack is required. This view is not necessarily correct  (see page 447).   When p
ozzolanas are used as a partial replacement for cement, the cement and the pozzolanas  are batched separately and mixed with the other ingredients in the concrete mixer. The required  properties (pozzolanas for such a purpose are prescribed by ASTM Standard 618-78.   Partial replacement of Portland cement by pozzolana has to be carefully defined, 
as the specific  gravity of pozzolanas is much lower than that cement; for instance, the specific gravity of fly ash  is	1.9 to 2.4, compared with 3.15 for cement. Thus replacement by weight results in a considerably  greater volume of the cementitious material in the mix. With replacement, concrete mixes have a  lower early strength than when Por
tland cement is used, but beyond about three months there is no  loss of strength. With lean mixes, there may even be a long-term gain of strength due to the  replacement (see Fig. 2.6). If equal early strength is required and pozzolana is to be used (e.g.  because of alkali - aggregate reactivity) then addition of pozzolana rather than partial re
placement of  cement is necessary. For  instance, when fly ash is used in lean mixes, almost 100 kg (220 lb)  pozzolana may be necessary to replace 50 kg (110 lb) of cement, but the amount of pozzolana is  lower in rich mixes. Because the continuing formation of hydrates fills the pores and also because  of the absence of free lime which could be 
reached out, partial replacement of Portland cement by  pozzolana reduces the permeability of concrete: a 7- to 10-fold reduction has been reported.    Fig. 2.6. Effect of partial replacement of  Portland cement by pozzolana on the strength  development of concrete made with 167 kg of "cement" material per cubic metre of  concrete (282 lb / yd3)  
	It must be remembered that, although pozzolana may be cheaper than Portland cement, the use of  an additional material on site (and especially of an extremely fine one) results in additional cost.  However, energy considerations as well as technical reasons have led to renewed interest in the  material. It is very likely that the 1980s will see
 an extensive use of portland-pozzolana cements.  Indeed, the U.S. government is proposing to specify the use of fly ash in all appropriate  construction involving Federal funds.      White Cement     For architectural purposes white concrete or, particularly in tropical countries, a pastel colour  paint finish are sometimes required. To achieve b
est results it is advisable to use white cement  with, of course, a suitable fine aggregate and, if the surface is to be treated, also coarse aggregate.  This type of cement has also the advantage that it is not liable to cause staining, since it has, low  content of soluble alkalis.   White Portland cement is made from raw materials containing ve
ry little iron oxide and  manganese oxide. China clay is generally used, together with chalk or limestone, free from  specified impurities. Oil is used as fuel for the kiln in order to avoid contamination by coal ash.  Since iron acts as a flux in clinkering, its absence necessitates higher kiln temperatures but  sometimes cryolite (sodium alumini
um fluoride) is added as a flux.   Contamination of the cement with iron during grinding has also to be avoided. For this reason,  instead of the usual ball mill, the rather inefficient pebble grinding or nickel and molybdenum alloy  balls are used. The cost of grinding is thus higher, and this, coupled with the more expensive raw  materials, make
s white cement rather expensive (about double the price of ordinary Portland cement).   Because of this, white cement concrete is often used in the form of a facing placed against ordinary  concrete backing, but great care is necessary to ensure full bond between the two concretes. To obtain  good colour, white concrete of rich-mix proportions is 
generally used, the water / cement ratio being not  higher than about 0.4. A possible saving in some cases can be achieved by a partial replacement of  white cement by blast-furnace slag, which has a very light colour.   When a pastel colour is required white concrete can be used as a base for painting. Alternatively,  pigments can be added to the
 mixer, but it is essential that the pigments do not affect adversely  the development of strength of the cement or affect air entrainment. For instance, carbon black  reduces the air content of the mix. For this reason, some pigments are marketed in the U.S.A.  with an interground air-entraining agent; it is of course essential to be aware of thi
s at the mix design  stage. Mixing of concrete with pigments is not common because it is rather difficult to maintain a  uniform colour of the resulting concrete. An improvement in the dispersion of the pigment can be  obtained by the use of superplasticizers (see p. 110).   A better way to obtain a uniform and durable coloured concrete is to use 
coloured cement. This  consists of white cement interground with 2 to 10 per cent of pigment. Specifications for the use  of this type of cement are  given by the individual manufacturers of this rather specialized product.  The specification for pigments is given in BS 1014 : 1975. Because the pigment is not  cementitious, slightly richer mixes t
han usual should be used.   A typical compound composition of white Portland cement is given in Table 2.5 but the C3S and  C2S contents may vary widely. White cement has a slightly lower specific gravity than ordinary  Portland cement, generally between 3.05 and 3.10. The strength of white Portland cement is  usually lower than that of ordinary Po
rtland cement but white cement nevertheless satisfies the  requirements of BS 12: 1978.    Table 2.5: Typical Compound Composition of white Portland Cement     White high-alumina cement is also made; this is considered on p. 100.      Other Portland Cements    Among the numerous cements developed for special uses, anti-bacterial cement is of inter
est. It is  a Portland cement interground with an anti-bacterial agent which prevents microbiological  fermentation. This bacterial action is encountered in concrete floors of  food processing plants  where the leaching out of cement by acids is followed by fermentation caused by bacteria in the  presence of moisture. Anti-bacterial cement can als
o be successfully used in swimming pools, public  baths and similar places where bacteria or fungi are present.   Another special cement is the so-called hydrophobic cement, which deteriorates very little during  prolonged storage under unfavourable conditions. This cement is obtained by intergrinding  Portland cement with 0.1 to 0.4 per cent of o
leic acid. Stearic acid or pentachlorophenol can also  be used. These additions increase the grindability of clinker, probably due to electrostatic forces  resulting from a polar orientation of the acid molecules on the surface of the cement particles.   Oleic acid reacts with alkalis in cement to form calcium and sodium oleates which foam, so tha
t  air-entraining results. When this is not desired a detraining agent, such as tri-n-butyl phosphate,  has to be added during grinding.   The hydrophobic properties are due to the formation of a water repellent film around each  particle of cement. This film is broken during the mixing of the concrete, and normal hydration  takes place but early 
strength is rather low.   Hydrophobic cement is similar in appearance to ordinary Portland cement but has a characteristic  musty smell. In handling, the cement seems more fluid than other Portland cements.   Hydrophobic cement should not be confused with waterproofed cements, which are claimed to  make a more impermeable concrete than ordinary Po
rtland cement. There is considerable  controversy about the effectiveness of these waterproofed cements.   Masonry cement, used in mortar in brickwork, is made by intergrinding  very finely ground  Portland cement, limestone and an air-entraining agent, or alternatively Portland cement and  hydrated lime, granulated slag or an inert filler, and an
 air-entraining agent. Masonry cements  make a more plastic mortar than ordinary Portland cement, they also have a greater  water-retaining power and lead to lower shrinkage. The strength of masonry cements is  lower than that of ordinary Portland cement, particularly since a high air content is  introduced, but this low strength is generally an a
dvantage in brick construction  Masonry cement must not be used in structural concrete.   Finally, mention should be made of blended Portland cements. It will be  recalled that, in most countries, specifications for Portland cement do not allow any  addition to clinker other than gypsum and water. With moves to save energy, the idea to  add some i
nert filler to Portland cement has been advanced; this then would be blended  Portland cement The most likely filler is limestone ground to the same fineness as  Portland cement, the proportion of the addition being 10 to 15 per cent of the total.   The filler has no cementitious value but it improves workability, and blended Portland  cements are
 likely to become extensively used in low-strength concrete. Indeed, one can  argue that, for many purposes, the existing high-quality Portland cements are too good,  so that there is intrinsic merit, and not only energy saving, in this development.   In a non-explicit way, some dilutants of Portland cement are allowed in the United  States, as AS
TM Standard C 465-74 allows processing additions, providing they do not  reduce the strength of the cement by more than 5 per cent. ASTM has also set up a  sub-committee to consider a possible standard for blended cements. Finland already allows  15 per cent of filler, and East Germany produces blended cements with 20 per cent of filler. In  Switz
erland, there is a blended cement with a maximum slag content of 5 per cent. In France, a 1979  specification provides for a ciment portland compose with up to 35 per cent of additive which, at least  in theory, can be entirely in the form of a filler; in practice, 10 to 15 per cent is used.  A filler is defined in this specification as a finely-g
round material which, owing to  its physical properties, influences some properties of cement, such as workability,  permeability, capillarity, and cracking. Appropriate specifications in other countries  are expected in the near future.   It may be noted that some of the inert components of blended cements are perhaps  not entirely inert. For ins
tance, it has been found by Zielinska  that CaCO3 reacts  with C3A and C4AF to produce 3CaO.Al2O3.CaCO3.11H2O.   There is one caveat about the use of blended cements: it is important to know  whether a given cement is "pure" or blended, and proper classification and labelling  are desirable. Moreover, we need reliable information on the durability
 of concrete  made with blended cements, especially under conditions of wetting and drying or  freezing and thawing: absorption of water by capillary action may play a role here.  Nevertheless, in the future, a more extensive use of blended cements in the lower  grades of concrete is likely.      Natural Cement    This is the name given to a cemen
t obtained by calcining and grinding a so-called cement  rock, which is a clayey limestone containing up to 25 per cent of argillaceous material.  The resulting cement is similar to Portland cement, and is really intermediate between  Portland cement and hydraulic lime.   Since natural cement is calcined at temperatures too low for sintering, it c
ontains  practically no C3S and is therefore slow hardening. Natural cements are rather variable  in quality as adjustment of composition by blending is not possible. Because of this, as  well as for economic reasons, natural cements are nowadays little used. In the United  States, natural cements represent no more than one per cent of the product
ion of all  Portland cements.      Expansive Cements   For many purposes it would be advantageous to use a cement which does not change in  volume owing to drying shrinkage or, in special cases, even expands on hardening.  Concrete containing such a cement expands in the first few days of its life, and a form  of pre-stress is obtained by restrain
ing this expansion with steel reinforcement: steel  is put in tension and concrete in compression. Restraint by external means is also  possible.   Cements of this type have been developed by H. Lossier in France, who used a  mixture of Portland cement, an expanding agent, and a stabilizer. The expanding agent is  obtained by burning a mixture of 
gypsum, bauxite, and chalk, which form calcium sulphate  and calcium aluminate (mainly C5A3)  In the presence of water, these compounds react to  form calcium sulphoaluminate hydrate (ettringite), with an accompanying expansion of the  paste. The stabilizer, which is blast-furnace slag, slowly takes up the excess calcium  sulphate and brings expan
sion to an end. Very careful proportioning of the cement  ingredients is necessary in order to obtain the desired expansion. Generally, about 8 to  20 parts of the "sulphoaluminate" clinker are mixed with 100 parts of Portland cement  and 15 parts of the stabilizer.   Another type of expanding cement, called high-energy expanding cement, is made b
y  intergrinding Portland cement clinker, high-alumina cement clinker, and gypsum,  approximately in the proportions 65 : 20 : 15. Expansion is due to the formation of calcium  sulphoaluminate, as in Lossier's cement, and takes place within 2 or 3 days of casting.  This is Type M cement in the classification of the ASTM Standard C 845-76T.   High-
energy expanding cement is quick-setting and rapid-hardening, reaching a strength  of about 7 MPa (1000 psi) in 6 hours, and 50 MPa (7000 psi) in 28 days. The cement has a  high resistance to sulphate attack.   A more recent development is expanding cement known as Type K in the ASTM  classification, developed in California. The ingredients of thi
s cement are similar to  those used by Lossier, but the material selection and clinkering conditions (maximum  temperature of about 1300 C) of the expanding agent lead to the formation of an anhydrous  calcium sulphoaluminate (C4A3.SO3). The expansion is due to the formation of hydrated  calcium sulphoaluminate (ettringite), as in Lossier's cement
, but the rate  and magnitude of expansion appear to be more reliable. A similar expanding agent is  manufactured in Japan, but intermixing with cement is done in the mixer.   There exists also a Type S cement which has a high C3A content and an amount of  interground calcium sulphate above that which is usual in Portland cement.   Since the devel
opment of expansion takes place only as long as the concrete is moist,  curing must be carefully controlled, and the use of expanding cement requires skill and  experience. The use of admixtures demands special care as they may affect expansion,  but there are no problems with air entrainment.   Strictly speaking, the use of expanding cement canno
t produce "shrinkless" concrete,  as shrinkage takes place after the moist curing has ceased, but the magnitude of  expansion can be adjusted so that the expansion and  the subsequent shrinkage are  numerically equal.   There are some practical difficulties still to be resolved in order to ensure   the required expansion under variable site condit
ions. The important requirement is  that CaO, SO3, and especially Al2O3, become available for  ettringite formation at the right time. Specifically, a major part of it must form  after a certain strength has been attained; otherwise, the expansive force will be  dissipated in the deformation of the still plastic concrete and no stress against the 
 restraint will result. On the other hand, if the ettringite continues to form rapidly for  too long, disruptive expansion may occur. The controlling factors are the presence of CaSO4  and the ratio of sulphate to aluminate in the paste. A complicating factor arises from the fineness of  the cement: for a given sulphate content, the greater  the f
ineness the smaller the  expansion: doubling the fineness halves the expansion. Also, clearly, the higher  the cement content of the concrete the higher its expansion. Likewise, expansion is greater the higher  the modulus of elasticity of the aggregate.   Concretes based on expanding cements have been classified by the American Concrete  Institut
e: two basic types are recognized. Shrinkage-compensating concrete  induces compressive stresses which approximately offset the tensile stresses induced by  shrinkage (about 0.2 to 0.7 MPa (30 to 100 psi)). Self stressing concrete is concrete in  which the induced compressive stresses are large enough to result in a significant  compressive stress
 after drying shrinkage has occurred (about 1 to 3.5 MPa (150 to 500  psi)). It is clear that is both these definitions involve a restraint of the expansion, usually in the form of  reinforcement, preferably triaxial. Indeed, it is not expansion that is of interest but  an induced compressive stress which can compensate for tensile stresses that w
ould  otherwise manifest themselves as tensile strains and possibly as cracks. A Standard  laying down the practice for the use of shrinkage-compensating concrete was issued by  the American Concrete Institute in 1977. It may be noted that many properties of  this concrete, such as strength, modulus of elasticity, inherent shrinkage, and  resistan
ce to freezing and thawing, are similar to those of corresponding Portland  cement concretes, but the loss of slump is faster. The resistance to sulphate attack  may be impaired, especially with types M and S.  The performance of these concretes on a limited scale has been found to be encouraging  and there is a good prospect of prestressing concr
ete by the use of expanding cement.  Nevertheless, further improvements in the cements are necessary before they can be more  widely employed.      High-alumina Cement    The search for a solution to the problem of attack of gypsum-bearing waters on Portland  cement concrete structures in France led Jules Bied to the development of a high-alumina 
 cement, at the beginning of this century. This cement is very different in its  composition and also in some properties from Portland cements so that its structural use  is severely limited, but the concreting techniques are similar. For full treatment of  the topic, the reader may consult a specialized book.    Manufacture    From the name of th
e cement - high-alumina or aluminous - it can be inferred that it  contains a large proportion of alumina: the cement consists, in fact, of approximately  equal parts, about 40 per cent each, of alumina and lime, with some ferrous and ferric  oxides, and up to about 8 per cent of silica.   The raw materials are limestone or chalk, and bauxite. Bau
xite is a residual deposit  formed by the weathering under tropical conditions of rocks containing aluminium, and  consists of hydrated alumina, oxides of iron and titanium, with small amounts of silica.  There are no bauxite deposits in Great Britain and it is imported from Greece and  France.   In the British process of manufacture of aluminous 
cement, bauxite is crushed into  lumps not larger than 100 mm (or 4 in.). Dust and small particles formed during this  fermentation are cemented into briquettes of similar size as dust would tend to damp the  furnace. the second main raw material is usually limestone, also crushed to lumps of  about 100 mm (or 4 in.).   Limestone and bauxite in th
e required proportions are fed into the top of a furnace which  is a combination of the cupola (vertical stack) and reverberatory (horizontal) types.  Pulverized coal is used for firing, its quantity being about 22 per cent of the weight  of the cement produced. In the furnace, the moisture and carbon dioxide are driven off  and the materials are 
heated by the furnace gases to the point of fusion at about  1600 C. The fusion takes place at the lower end of the stack so that the molten material  falls into the reverberatory furnace and thence through a spout into steel pans. The  melt is now solidified into pigs, fragmented in a rotary cooler, and then ground in a  tube mill. A very dark gr
ey powder with a fineness of 250 to 320 m2 / kg is produced.   Because of the high hardness of high-alumina cement clinker, the power consumption and  the wear of tube mills are considerable. This, coupled with the high prime cost of  bauxite and the high temperature of firing, leads to a high price of high-alumina  cement, compared with, say, rap
id hardening Portland cement. The price is, however,  compensated for by some valuable properties for specific purposes.   It may be noted that, unlike the case of Portland cement, the materials used in the  manufacture of high-alumina cement are completely fused in the kiln. This fact gave rise  to the French name Ciment Fondu, which is now used 
in Britain as a trade name. Because trade names  are so widely used the other names should also be mentioned: Lightning cement (in Britain) and  Lumnite (in the United States).    Composition    Table 2.6 gives typical values of oxide composition of high-alumina cement. A minimum  alumina content of 32 per cent is prescribed by BS 915: 1972, which
 requires also the  alumina / lime ratio to be between 0.85 and 1.3.    Table 2.6: Typical Oxide Composition of High-alumina Cement     Considerably less is known about the compound composition of high-alumina cement than of  Portland cement, and simple method of calculation is available. The main cementitious  compounds are calcium aluminates of 
low basicity: CA and C5A3. The latter is now  believed to be really C12A7. Other phases are also present: C6A4.FeO.S and an  isomorphous CgA4.MgO.S. C2S or C2AS does not account for more than a few per cent  and there are, of course, minor compounds present but no free lime can exist. Thus  unsoundness is never a problem in high-alumina cement alt
hough BS 915: 1972 prescribes  the conventional Le Chatelier test.    Hydration    The hydration of CA, which has the highest rate of strength development, results in the formation of CAH10, a small quantity of C2AH8, and of alumina gel (Al2O3.aq). With  time, these hexagonal CAH10 crystals, which are unstable both at normal and at higher  tempera
tures, become transformed into cubic crystals of C3AH6 and alumina gel. This  transformation is encouraged by a higher temperature and a higher concentration of lime  or a rise in alkalinity.   C5A3 is believed to hydrate to C2AH8. C2S forms CSHx, the lime liberated by hydrolysis  reacting with excess alumina; no Ca(OH)2, exists. The reactions of 
hydration of the other  compounds, particularly those containing iron, have not been determined with any degree  of certainty, but the iron held in glass is known to be inert.   The water of hydration of high-alumina cement is calculated to be up to 50 per cent of  the weight of the dry cement, which is about twice as much as the water required fo
r  the hydration of Portland cement, but mixes with a water / cement ratio as low as 0.35 are  practical and indeed desirable.    Resistance to Chemical Attack    As mentioned earlier, high-alumina cement was first developed to resist sulphate attack,  and it is indeed highly satisfactory in this respect. This resistance to sulphates is  due to th
e absence of Ca(OH)2 in hydrated high-alumina cement and also to the protective  influence of the relatively inert alumina gel formed during hydration. However, lean  mixes are very much less resistant to sulphates. Also, the chemical resistance  decreases drastically after conversion (see p. 92).   High-alumina cement is not attacked by CO2 disso
lved on pure water.  The cement is not acid-resisting but it can withstand tolerably well very dilute  solutions of acids (pH greater than 3.5 to 4 0) found in industrial effluents, but not of  hydrochloric, hydrofluoric or nitric acids. On the other hand, caustic alkalis, even in  dilute solutions, attack high-alumina cement with great vigour by 
dissolving the alumina  gel. The alkalis may have their origin outside (e.g. by percolation through Portland  cement concrete) or in the aggregate. The behaviour of this cement in the presence of  many agents has been studied by Hussey and Robson.   It may be noted that, although high-alumina cement stands up extremely well to sea  water, this wat
er should not be used as mixing water; the setting and hardening of the  cement are adversely affected, possibly because of the formation of chloroaluminates.  Likewise, calcium chloride must never be added to high-alumina cement.      Physical Properties of High-alumina Cement Concrete    Another outstanding feature of high-alumina cement is its 
very high rate of strength  development  About 80 per cent of its ultimate strength is achieved at the age of 24  hours, and even at 6 to 8 hours the concrete is strong enough for the side formwork to  be struck and for the preparation for further concreting to take place. The high rate of  gain of strength is due to its rapid hydration, which in 
turn means a high rate of heat  development. This can be as high as 38 J per gram per hour (9 cal / g / h) whereas for rapid  hardening Portland cement the rate is never higher than 15 J per gram per hour (3.5  cal / g /h). However, the total heat of hydration is of the same order for both types of  cement.   Concrete made with high-alumina cement
 and high-alumina cement clinker as aggregate,  with a water / cement ratio of 0.5, can reach a strength  of about 100 MPa (14 000 psi) in  24 hours, and 120 MPa (18000 psi) in 28 days at cool temperatures. This extremely high  strength development is due to the cementitious character of the aggregate but this  aggregate is, of  course, very expen
sive.   It should be stressed that the rapidity of hardening is not accompanied by rapid  setting. In fact, high-alumina cement is slow setting but the final set follows the  initial set more rapidly than is the case in Portland cement. Typical values for high-alumina  cement are: initial set - 4 to 5 hours, final set - 30 minutes later.   BS 915 
: 1972 requires the initial set to take place between two and six  hours after mixing, and the final set not more than two hours after the initial set.  Of the compounds present in the high-alumina cement, C5A3 sets in a few minutes while  CA is considerably more slow-setting, so that the higher the C : A ratio in the cement  the more rapid the se
t. On the other  hand, the higher the glass content of the cement  the slower the setting. It is likely that, because of its rapid setting properties,  C5A3 is responsible for the loss of workability of many high-alumina cement concretes,  which takes place within 15 or 20 minutes of mixing.   The setting time of high-alumina cement is greatly aff
ected by the  addition of plaster, lime, Portland cement and organic matter and for this  reason no additives should be used.    Fig. 2.7. Setting time of Portland - high-alumina cement mixtures     In the case of Portland cement - high-alumina cement mixtures, when either cement  constitutes between 20 and 80 per cent of the mixture, flash set ma
y occur. Typical data  are shown in Fig. 2.7 but actual values vary for different cements, and trial tests  should be made with any given cements. The accelerated setting is due to the formation of  a hydrate of C4A by the addition of lime from the Portland cement to calcium aluminate  from the high-alumina cement. Also, gypsum contained in the Po
rtland cement may react with  hydrated calcium aluminates, and as a consequence the now non-retarded Portland cement  may exhibit a flash set.   Mixtures of the two cements in suitable proportions are used when rapid setting is of  vital importance, e.g. for stopping the ingress of water, or for temporary construction  between the tides, but the u
ltimate strength of such pastes is quite low.   Because of the rapid setting just described, in normal concrete construction it is  essential to make sure that the two cements do not come in contact with one another.  Thus, placing concrete made with one type of cement against concrete made with the other  must be delayed by at least 24 hours if h
igh-alumina cement was cast first, or 3 to 7  days if the earlier concrete was made with Portland cement. Contamination through plant  or tools must also be avoided.  It may be noted that, for equal mix proportions, high-alumina cement produces a somewhat  more workable mix than when Portland cement is used. This may be due to the lower total  sur
face area of high-alumina cement particles which have a "smoother" surface than  Portland cement particles, since high-alumina cement is produced by complete fusion of  the raw materials.   Creep of high-alumina cement concrete has been found to differ little from the creep of  Portland cement concretes when the two are compared on the basis of th
e stress / strength  ratio.    Conversion of High-alumina Cement    The high strength of high-alumina cement concrete referred to on Page 90 (see also Fig.  2.8) is reached when the hydration of CA results in the formation of CAH10 with a small  quantity of C2AH8 and of alumina gel (Al2O3.aq). The hydrate CAH10 is, however,  chemically unstable bo
th at higher and normal temperatures and becomes transformed into  C3AH6 and alumina gel. This change is known as conversion, and, since the symmetry of  the crystal systems is pseudo-hexagonal for the decahydrate and cubic for the  sesquihydrate, one can refer to it as the change from the hexagonal to cubic form.    Fig. 2.8. Strength development
 of concrete cylinders with different water / cement ratios  made with high-alumina cement and cured at 18 C (64 F) and 95 per cent relative humidity     An important feature of hydration of high-alumina cement is that at higher temperatures  only the cubic form of the calcium aluminate hydrate can exist; at room temperature  either can, but the h
exagonal crystals spontaneously, albeit slowly, convert to the  cubic form. Because they undergo a spontaneous change, the hexagonal crystals can be  said to be unstable at room temperature, the final product of the reactions of  hydration being the cubic form. Application of heat speeds up the process. This then is  conversion: an unavoidable cha
nge of one form of calcium aluminate hydrate to another,  and it is only reasonable to add that this type of change is not an uncommon phenomenon  in nature.   Before discussing the significance of conversion we should briefly describe the  reaction. Conversion both of CAH10 and of C2AH8 proceeds direct; for instance:    (reaction).     It should 
be noted that, although water appears as a product of the reaction,  conversion can take place only in the presence of water and not in desiccated concrete  because redissolving and reprecipitation are involved. As far as neat cement paste is  concerned, it has been found that, in sections thicker than 25 mm, the interior of  the hydrating cement 
has an equivalent relative humidity of 100 per cent regardless of  the environmental humidity, so that conversion can take place. The influence of the  ambient humidity is thus only on concrete near the surface.   The cubic product of conversion, C3AH6 is stable in a solution of calcium hydroxide at  25 C but reacts with a mixed Ca(OH)2-CaSO4 solu
tion to form 3CaO.Al2O3.3CaSO4.31H2O both  at 25 C and at higher temperatures.  The degree of conversion is estimated from the percentage of C3AH6 present as a  proportion of the sum of the cubic and hexagonal hydrates together, i.e.    (formula).     The relative weights of the compounds are derived from the measurements of endothermic  peaks in 
a differential-thermal-analysis thermogram.   However, unless the determination can be made under CO2-free conditions, there is a risk  of decomposition of C3AH6 into AH3. The degree of conversion can be determined also in  terms of the latter compound because, fortuitously, the weights of C3AH6 and AH3  produced in conversion are not very differe
nt. Thus we can write:    (formula).     While the two expressions do not give exactly the same result, at high degrees of  conversion the difference is not significant. Most laboratories report the result to the nearest 5 per cent.  Concrete which has converted about 85 per cent would be considered as fully converted.   The rate of conversion dep
ends on temperature; some actual data are shown in Table  2.7. The relation between the time necessary for one-half of the CAH10 to convert  and the temperature of storage of 13 mm (1/2 in.) cubes of neat cement paste with a  water / cement ratio of 0.26 is shown in Fig. 2.9. It is likely that, for the more porous  concretes of practical mix propo
rtions, the periods are much shorter as full conversion  has been observed after some twenty years at 20 C or thereabouts Thus data on neat  cement pastes with very low water / cement ratios should be used circumspectly, but they  are nevertheless of scientific interest.    Table 2.7: Development of Conversion with Age    Fig. 2.9. Time for half-c
onversion of neat high-alumina cement pastes cured at various temperatures  (13 mm (1/2 in.) cubes)     In the past, the main concern with control of conversion lay in preventing a  temperature rise of concrete due to the rapid development of the heat of hydration. For  this reason, cooling of freshly-placed concrete was insisted upon. Later, atte
mpts were  made to show that conversion occurs only under certain conditions and that its effects  could under many circumstances be avoided. We can now state with reasonable certainty  that under practical conditions obtaining in temperate climates, and of course at  higher temperatures, conversion inevitably occurs. The rate of conversion is  ac
celerated by a higher temperature Even short periods of higher temperature increase  conversion and their effect is cumulative. It follows that buildings where higher  temperatures obtain only at times are also affected. The speed of conversion within a  given building may vary, being highest near sources of heat, especially with underfloor  heati
ng.   The practical interest in conversion lies in the fact that it leads to a loss of  strength of high-alumina cement concrete. The most likely explanation why this is so is  in terms of the densification of the calcium aluminate hydrates: typically, the density  would be 172 g / ml for CAH10 and 2.53 for  C3AH6. Thus, under conditions such that
 the  overall dimensions of the body are constant (as is the case in set cement paste),  conversion results in an increase in the porosity of the paste.   Numerous proofs of this are available, a recent and particularly convincing one being  the measurement of air permeability of converted compared with unconverted high-alumina  cement concrete. (
see Fig. 2.10).    Fig. 2.10. Airflow through concrete: (a) unconverted high-alumina cement concrete;  (b) converted high-alumina cement concrete; (c) Portland cement concrete (temperature 22 to  24 C (72 to 75 F), relative humidity 36 to 41 per cent; pressure difference 10.7 kPa)     As shown on page 271, the strength of hydrated cement paste or 
of concrete is very  strongly affected by its porosity; porosity of 5 per cent can reduce the strength by  more than 30 per cent, and a 50 per cent reduction in strength would be caused by a  porosity of the order of 8 per cent. This order of magnitude of porosity can be induced  by conversion in high-alumina cement concrete.   It follows that, si
nce conversion takes place in concretes and mortars of any mix  proportions, they lose strength when exposed to a higher temperature, and the general  pattern of the strength loss versus time curves is similar in all cases. However, the  degree of loss is a function of the water / cement ratio of the mix, as shown in Fig.  2.11. The mix proportion
s and percentage loss are given in Table 2.8. It is clear that  the loss, either in megapascals or as a fraction of the strength of cold-cured concrete,  is smaller in mixes with low water / cement ratios than in mixes with high water / cement  ratios.  Fig. 2.11. Influence of the water / cement ratio on the strength of high-alumina cement  concre
te cubes curved in water at 18 C and 40 C for 100 days	Table 2.8: Influence of Water / Cement Ratio on Loss of Strength on Conversion.   It may be observed that the shape of the strength versus water / cement ratio curves  for storage at 18 C (Fig. 2.11) is dissimilar from the usual curves for Portland cement  concretes. This is characteristic 
of concretes made with high-alumina cement, and has  been confirmed also for cylinders both of standard size and other height / diameter  ratios.   The values shown in Fig. 2.11 are no more than typical, and clearly some variation  would be found with different cements, but the pattern of  behaviour is the same in all  cases. It is important to no
te that the residual strength of mixes with moderate and  high water / cement ratios, say over 0.5, may be so low as to be unacceptable for most  structural purposes.   The relative difference in the loss of strength of high-alumina cement concrete  made with a high water / cement ratio and with a low water / cement  ratio should be noted. In the 
latter case, say at a water / cement ratio of 0.35, even  after full conversion the strength could be deemed to be adequate for all structural  purposes. Two caveats should, however, be expressed. First, under practical conditions  of manufacture of concrete, it is not possible to guarantee that the water / cement ratio will  not be occasionally e
xceeded by 0.05 or even by 0.10; this has been repeatedly demonstrated.  Second, converted high-alumina cement concrete, even if of adequate strength, is more porous and  therefore more liable to chemical attack than before conversion.   Not only the water / cement ratio but also the richness of the mix may affect the loss of  strength on conversi
on. For a given water / cement ratio, the leaner the mix the lower the  porosity of the cement paste and, consequently, the lower the relative loss in strength  with time.   The important point is that, despite some attempts to prove the contrary, all mixes  lose strength on conversion. Test results on the strength of 13 mm (1/2 in.) cubes of neat
	high-alumina cement paste show that even at a water / cement ratio of 0.30, the strength  at 50 C (122 F) is only 34 per cent of the strength at 18 C (64 F); at a water / cement ratio  of 0.50, the fraction is only 20 per cent. These values are much lower than those often  quoted for concrete but it must be stressed that the data refer to neat c
ement pastes.	The value of the minimum strength after conversion was also controversial, but there  are now available long-term data which give a clear and broad picture of the residual  strength which can be expected. Actual values may vary with the particular cement used,  but the following can be used as typical values:    (table).     The lo
ss of strength is often accompanied by a change in the colour of the cement paste  from blackish-grey to brown or yellow-brown. This happens probably because the increase  in porosity on conversion facilitates the oxidation of the ferrous compounds in the  cement paste. This is why conversion and change in colour may occur together, but  concrete 
which is porous due to bad mix proportions can also change colour in the absence  of conversion.   In view of the effects of conversion, high-alumina cement is no longer used in  structural concrete above or below ground, but it is a valuable material for repair work  of limited life and in temporary works.      Refractory Properties of High-alumi
na Cement    High-alumina cement concrete is one of the foremost refractory materials but it is  important to be clear about its performance over the full temperature range. Between  room temperature and about 500 C, high-alumina cement concrete loses strength to a greater  extent than Portland cement concrete, then up to 800 C the two are compara
ble, but above  about 1000 C high-alumina cement gives excellent performance.   Fig. 2.12 shows the behaviour of high-alumina cement concrete made with four different  aggregates over a temperature range up to 1100 C. The minimum strength varies between 5  and 26 per cent of the original value but, depending on the type of aggregate, above 700  to
 1000 C, there is a gain in strength due to the development of a ceramic bond. This  bond is established by solid reactions between the cement and fine aggregate, and  increases with an increase in temperature and with the progress of the reactions.    Fig. 2.12. Strength of high-alumina cement concretes made with different aggregates as a    func
tion of temperature.	  As a result, high-alumina cement concrete can withstand very high temperatures: with  crushed firebrick aggregate up to about 1350C , and with special aggregates, such as  fused alumina or carborundum, up to 1600 C. A temperature as high as 1800 C can be  withstood over prolonged periods of time by concrete made from spec
ial white calcium  aluminate cement with fused alumina aggregate. This cement is made using alumina as  raw material and contains 70 to 80 per cent of A12O3, 20 to 25 per cent lime, and only  about 1 per cent of iron and silica; the composition of the cement approaches C3A5. It  is appropriate to mention that the price of the cement is about $500 
(f250) per tonne.   Refractory concrete made with high-alumina cement has a good resistance to acid attack  (e.g. acids in flue gases), the chemical resistance being in fact increased by firing at  900 to 1000 C. The concrete can be brought up to service temperature as soon as it has  hardened, that is it does not have to be pre-fired. While refra
ctory brickwork expands  on heating and therefore needs expansion joints, high-alumina cement concrete can be  cast monolithically, or with butt joints only (at 1 m to 2 m), to exactly the required  shape and size. The reason for this is that the loss of water on first firing results in  a contraction approximately equal to the thermal expansion o
n	heating, so that the nett  dimensional change (depending on aggregate) is small. Upon subsequent cooling, for  instance during the shut-down of plants, butt joints would open slightly due to the  thermal contraction but they would close up again on re-heating. It is worth noting that  refractory high-alumina cement concrete can withstand a consi
derable thermal shock.  Refractory linings can be made by shotcreting high-alumina cement mortar.   For insulating purposes, when temperatures up to about 950 C are  expected, lightweight concrete can be made with high-alumina cement and lightweight  aggregate. Such concrete has a density of 500 to 1000 kg / m3 (30 to 60 lb / ft3) and a  thermal c
onductivity of	0.21 to	0.29 J / m2 s C / m (0.12 to 0.17 Btu / ft2 h F / ft).      Admixtures    Often, instead of using a special cement, it is possible to change some of the  properties of the cement in hand by the use of a suitable additive. A great number of  proprietary products is available: their effects are described by the manufacturers b
ut  the full details of the action of many of these additives, known as admixtures, are yet  to be determined, and the performance of any one admixture should be carefully checked  before it is used.   Admixtures may be classified according to the purpose for which they are used in  concrete; the approach of ASTM Standard C 494-79 can be used. In 
this chapter, we shall  consider only one well-tried accelerating admixture (Type C, according to the ASTM  classification) (calcium chloride), retarding admixtures or retarders.(Type B), water-reducing  admixtures (Type A), and also the so-called superplasticizers. The nomenclature  is somewhat confusing in that retardation refers to the setting 
of concrete while  acceleration primarily to the early strength development, i.e. to hardening (see p. 68),  more rapid setting being generally only coincidental. The classification of the British  Standard BS 5075 : Part 1 : 1974 is substantially similar; the standard lays down the  requirements for the various types of admixtures (Table 2.9).   
 Table 2.9: Specification for the Various Types of Admixtures According to BS 5075: Part 1: 1974     Additives inducing air entrainment are considered on p. 475. There exist also additives  for other purposes, such as air detrainment, fungicidal action, water-proofing, etc.,  but these are not sufficiently standardized. Here it suffices to say tha
t	water-proofing  admixtures (calcium stearate, butyl stearate, and calcium oleate) are supposed to repel  water by an electrostatic charge which they form after reacting with calcium ions on the  walls of the capillaries in the hydrated cement paste. It is doubtful whether the effect  persists over long periods. For fungicidal purposes, copper su
lphate and  pentachlorophenol have been suggested. These control the growth of algae or lichen on  hardened concrete but again their effectiveness is lost with time. Clearly, admixtures  which may prove toxic should not be used. Useful information about miscellaneous  admixtures is given in a guide of the American Concrete Institute.   The water-p
roofing admixtures should be distinguished from water repellents, based on  silicone resins, which are applied to the concrete surface. Waterproof membranes are  emulsion-based bitumen coatings, possibly with rubber latex, which produce a tough film  with some degree of elasticity. Their use requires specialist advice. Pigments can also  be consid
ered as an additive; they are mentioned on page 82.   An important feature of the majority of admixtures for concrete is that they are used  primarily on the basis of experience or ad hoc tests: theoretical information on a  scientific basis is generally not available to permit a reliable quantitative prediction  of behaviour in concrete under the
 various possible circumstances. This is due to the  marketing of admixtures largely as proprietary products. Acceptance requirements are  laid down by ASTM Standard C 494-79 and by British Standard BS 5075 : Part 1 : 1974.    Calcium Chloride    The addition of calcium chloride to the mix increases the rate of development of  strength, and this a
ccelerator is, therefore, sometimes used when concrete is to be  placed at low temperatures (in the region of 2 to 40 C (35 to 40 F)) or when urgent  repair work is to be done.   Calcium chloride increases the rate of heat liberation during the first few hours after  mixing, the action of CaCl2 being probably that of a catalyst in the reactions of
	hydration of C3S and C2S; it is possible that the reduction in the alkalinity of the  solution promotes the hydration of the silicates. The hydration of C3A is delayed  somewhat, but the normal process of hydration of cement is not changed.   Calcium chloride may be added to rapid hardening as well as to ordinary Portland  cement, and the more r
apid the natural rate of hardening of the cement the earlier  becomes apparent the action of the accelerator. Calcium chloride must not, however, be  used with high-alumina cement. With rapid hardening Portland cement the increase in  strength due to CaCl2 can be as much as 7 MPa (1000 psi) at 1 day while with ordinary  Portland cement this increa
se would be achieved only after 3 to 7 days. By the age of 28  days there is no difference between the strengths of rapid hardening cements with and  without CaCl2, but in the case of ordinary Portland cement the addition of CaCl2 would  still show an improved strength.   Hickey's results for cements of different types are shown in Fig. 2.13. The 
long-term  strength of concrete is believed to be unaffected by CaCl2. Calcium chloride is  generally more effective in increasing the early strength of rich mixes with a low  water / cement ratio than of lean ones.    Fig. 2.13. Influence of CaCl, on the strength of concretes made with different types of  cement (For Type V curve reference is 2.1
.)     The quantity of CaCl2 added to the mix must be carefully controlled. To calculate the  quantity required it can be assumed that the addition of 1 per cent of anhydrous CaCl2  (as a fraction of the weight of cement) affects the rate of hardening as much as a rise  in temperature of 6 C (11 F). A calcium chloride content of 1 to 2 per cent is
 generally  sufficient; the latter figure should not be exceeded unless a test with the cement to  be actually used in construction is made, as the effects of calcium chloride depend to a  certain degree on the composition of the cement. This applies in particular to the  setting time. Calcium chloride generally accelerates setting and an  excess 
of CaCl2 can cause a flash set. Typical figures showing the influence of CaCl2 on  the setting time are given in Table 2.10. This acceleration of setting makes the  addition of CaCl2 useful in repair work, for instance, when the ingress of water can be  prevented for a short time only.    Table 2.10: Influence of Calcium chloride on Setting Time  
		It is important that calcium chloride be uniformly distributed throughout the mix, and  this is best achieved by dissolving the additive in the mixing water before it enters  the mixer. It is convenient to prepare a concentrated aqueous solution of CaCl2, and  this is more easily obtained using commercial calcium chloride flakes rather than  gr
anular calcium chloride which dissolves very slowly. The flakes consist of CaCl2.2H2O  so that 1.37g of flakes is equivalent to 1g of anhydrous calcium chloride. When  different chlorides co-exist, calculation in terms of chloride ions is useful: 1.56g of  CaCl2 corresponds to 1 g of chloride ion.   In cases when the durability of concrete may be 
impaired by outside agencies the use of  calcium chloride may be inadvisable. For instance, the resistance of cement to sulphate  attack is reduced by the addition of CaCl2, particularly in lean mixes, and the risk of  an alkali - aggregate reaction, when the aggregate is reactive, is increased. However,  when this reaction is effectively controll
ed by the use of low-alkali cement and the  addition of pozzolanas, the effect of CaCl2 is very small. Another undesirable feature  of the addition of CaCl2 is that it increases the drying shrinkage usually by about 10  to 15 per cent, sometimes even more, and possibly increases also the creep.   Although the addition of CaCl2 reduces the danger o
f	frost attack during the first few  days after placing, the resistance of air-entrained concrete to freezing and thawing at  later ages is adversely affected. Some indication of this is given in Fig. 2.14.    Fig. 2.14. Resistance to freezing and thawing of concrete cured moist at 4 C (40 F)  for different contents of CaCl2     On the credit side
,	CaCl2 has been found to raise the resistance of concrete to erosion  and abrasion, and this improvement persists at all ages. When plain concrete is  steam cured, CaCl2 increases the strength of concrete and permits the use of a more  rapid temperature rise during the curing cycle (see p. 327).   The action of sodium chloride is similar to that 
of calcium chloride but is of lower  intensity. The effects of NaCl are also more variable and a depression in the heat of  hydration, with a consequent loss of strength at 7 days and later has been observed.  For this reason the use of NaCl is definitely undesirable. Barium chloride has been  suggested but it acts as an accelerator only under war
m conditions.   The possibility of corrosion of reinforcing steel by calcium chloride has for long been  a subject of controversy. The U.S. Bureau of Reclamation found no corrosion of  reinforcement when calcium chloride is used in correct proportions, but others in the  U.S. found that at high water / cement ratios some corrosion takes place, alt
hough the  attack is not progressive. However, the chloride ions in concrete are often non-uniformly  distributed and a number of cases of severe corrosion of reinforcement in structural  members exposed to the weather have been reported even when the quantity of calcium  chloride was apparently carefully controlled.   The situation can be summari
zed as follows. Calcium chloride must not be used in  concrete in contact with prestressing wires. Steam curing reinforced concrete with  calcium chloride leads to a serious risk of severe corrosion. In the United Kingdom,  since 1977, the use of calcium chloride in reinforced concrete or concrete containing  embedded metal is not permitted by the
 Code of Practice CP 110: 1972. The maximum  chloride ion content (as a percentage of weight of cement) arising from all sources,  including aggregate, is as follows: 0.1 for prestressed concrete, steam-cured reinforced  concrete; 0.2 for any concrete made with sulphate-resisting or supersulphated cement;  and 0.35 for all other concrete containin
g metal. The variability of the chloride ion content  is recognized by allowing the value of 0.35 to rise to 0.50 in 5 per cent of test results.  British Standard BS 5075: Part 1: 1974 prescribes a method of determining the chloride content  of admixtures and there is no difficulty in establishing the chloride content of aggregate  In the United S
tates, the American Concrete Institute proposes a revision of the Standard 201  to limit the total(i.e. from all sources) soluble chloride ion concentration as a percentage of  weight of cement to 0.06 for prestressed concrete, and 0.10 to 0.15 for other concretes.  Only concrete that will be permanently dry is exempt from this limitation. It may 
be  relevant to note that the use of calcium chloride in reinforced concrete was discontinued  in Germany around 1974.   In view of the above rules, it is necessary to determine the chloride content even of those  admixtures in which calcium chloride is only a minor component.   The limitations of the Code of Practice CP 110: 1972 may seem particu
larly severe, but  calcium chloride does contribute to the risk of corrosion of reinforcement, and the risk  can be avoided by the use of very rapid hardening cements or of chloride-free admixtures.  Most of these are based on calcium formate, which, being slightly acidic, accelerates the  process of hydration of cement. Sometimes, calcium formate
 is blended with corrosion  inhibitors such as soluble nitrites, benzoates and chromates. This type of admixture has a greater  accelerating effect at low temperatures than at room temperature and is reported not to increase  shrinkage to a significant extent. Calcium nitrite has also been suggested as an accelerator and so  have several other com
pounds.    Retarders    A delay in the setting of the cement paste can be achieved by the addition  to the mix of a retarding admixture. These admixtures generally slow down also the  hardening of the paste although some salts may speed up the setting but inhibit the  development of strength. Retarders do not alter composition or identity of produ
cts of  hydration.   Retarders are useful in concreting in hot weather, when the setting time is shortened  by the higher temperature and in preventing formation of cold joints. The delay in  hardening caused by the retarders can be used to obtain an architectural finish of exposed  aggregate: the retarder is applied to the interior surface of the
 formwork so that the  hardening of the adjacent cement is delayed. This cement can be brushed off after  formwork has been struck so that an exposed aggregate surface is obtained.   Retarding action is exhibited by sugar, carbohydrate derivatives, soluble  zinc salts, soluble borates, and others. In practice, retarders which are also  water-reduc
ing are more commonly used; these are described in the next section. Great care is  necessary in using retarders as in incorrect quantities they can totally inhibit the  setting and hardening of concrete. Cases are known of seemingly inexplicable results of  strength tests when sugar bags have been used for the shipment of aggregate samples to  th
e laboratory or when molasses bags have been used to transport freshly mixed  concrete. The effects of sugar depend greatly on the quantity used, and conflicting  results were reported in the past. It seems now that, used in a carefully controlled  manner, a small quantity of sugar (about 0.05 per cent of the weight of cement) will  act as an acce
ptable retarder; the delay in setting of concrete is about 4 hours.  However, the exact effects of sugar depend greatly on the chemical composition of  cement. For this reason, the  performance of sugar, and indeed of any retarder, should  be determined by trial mixes with the actual cement which is to be used in  construction.   A large quantity 
of sugar, say 0.2 to 1 per cent of the weight of cement, will  virtually prevent the setting of cement. Such quantities of sugar can therefore be used  as an inexpensive "kill", for instance when a mixer or an agitator has broken down and  cannot be discharged.   When sugar is used as a controlled set retarder, the early strength of concrete is  s
everely reduced but beyond about 7 days there is an increase in strength of several  per cent compared with a non-retarded mix. This is probably due to the fact that  delayed setting produces a denser gel (cf. p. 318).   It is interesting to note that the effectiveness of an admixture depends on the time  when it is added to the mix: a delay of ev
en 2 minutes after water has come into  contact with the cement increases the retardation; such a delay can be achieved by a  suitable sequence of feeding the mixer. The increased retardation occurs especially  with cements which have a high C3A content because, once some C3A has hydrated, it  does not absorb the admixture so that more of it is le
ft to retard the hydration of the  calcium silicates, which occurs through adsorption on to the calcium hydroxide nuclei.   Retarders tend to increase the plastic shrinkage because the duration of the plastic  stage is extended but drying shrinkage is not affected.   The mechanism of action of retarders has not been established with certainty. We 
know  that they modify the crystal growth or morphology and this results in a more efficient  barrier to further hydration than is the case without an admixture. The admixtures are  finally removed from solution by being incorporated into the hydrated material but this  does not necessarily mean the formation of different complex hydrates. This is
 also  the case with water-reducing and retarding admixtures (such as ligno-sulphonates):  Khalil and Ward showed that the linear relation between the heat of hydration and  weight of non-evaporable water (see Fig. 2.15) is unaffected by the addition of the  admixture.    Fig. 2.15. Relation between the non-evaporable water content of cement and h
eat of hydration    Water-reducing Admixtures    According to ASTM Standard C 494-79, admixtures which are water-reducing only are called  Type A, but if the water-reducing properties are associated with set retardation, the  admixture is classified as Type D. There exists also water-reducing and accelerating  admixtures (Type E) but these are of 
little interest. The two main groups of admixtures  of Type D are -  (a) Lignosulphonic acids and their salts (known as Class 1 in the ASTM nomenclature and  (b) hydroxylated carboxylic acids and their salts (known as Class 3).   The modifications and derivatives of these, known as Class 2 and 4 respectively, do not  act as retarders, and may even
 behave as accelerators (see Fig. 2.16): they are  therefore of Type A or E (see p. 100).     2.16 Effect of various water-reducing admixtures on the setting time of concrete     The principal active components of the admixtures are surface-active agents. These  are substances which are concentrated at the interface between two immiscible phases a
nd  which alter the physico-chemical forces acting at this interface. The substances are  adsorbed on the cement particles, giving them a negative charge which leads to repulsion  between the particles and results in stabilizing their dispersion; air bubbles are also  repelled and cannot attach to the cement particles. In addition, the charge caus
es the  development around each particle of a sheath of oriented molecules which prevent a  close approach of the particles to one another. The particles have, therefore, a greater  mobility, and water freed from the restraining influence of the flocculated system  becomes available to lubricate the mix so that the workability is increased.   One 
effect of dispersion is to expose a greater surface area of cement to hydration,  which progresses therefore at a higher rate in the early stages. For this reason, there  is an increase in the strength of concrete, compared with a mix of the same water / cement  ratio but without the admixture. A more uniform distribution of the dispersed cement  
throughout the concrete may also contribute to the improved strength. The increase  in strength is particularly noticeable in very young concretes but under certain  conditions persists for a long time.   The influence of admixtures on strength varies considerably with the composition of  cement, the greatest increase in strength occurring when us
ed with cements of low alkali  or low C3A content. With some cements, the influence of admixtures is very small, but in  general terms admixtures are effective with all types of Portland cement and also with  high-alumina cement. Some water-reducing admixtures are more effective when used in  mixes containing pozzolanas than in plain mixes.   The 
reduction in the quantity of mixing water that is possible owing to the use of  admixtures varies between 5 and 15 per cent (Table 2.11). A part of this is in many  cases (especially with Class 1 admixtures) due to the entrained air introduced by the  admixture. The actual decrease in the mixing water depends on the cement content,  type of aggreg
ate used, presence of air-entraining agents or pozzolanas. It is therefore  apparent that trial mixes, containing the actual materials to be used on the job, are essential  in order to determine the type and quantity of admixture to achieve optimum properties.  Data given by the manufacturers of admixtures cannot generally be unquestionably accept
ed.  It should also be noted that even though set is retarded, admixtures do not always reduce the rate of  loss of workability with time. A further aspect to be considered is the danger of segregation of the  concrete; this can be investigated by measurement of bleeding.     Table 2.11: Water Reduction Obtained with Different Water-reducing and S
et-retarding Admixtures     The quantity of admixture represents generally only a fraction of one per cent of the  weight of cement in the mix so that use of reliable dispensing equipment is essential.    Superplasticizers (High-range Water-reducers)    These are a modern type of water-reducing admixture, much more effective than the  admixtures d
iscussed in the preceding section. Chemically, they are sulphonated melamine  formaldehyde condensates and sulphonated napthalene formaldehyde condensates, the latter  being probably the somewhat more effective of the two in dispersing the cement and  generally having also some retarding properties. At a given water / cement ratio, this  dispersin
g action increases the workability of concrete, typically by raising the slump  from 75 mm (3 in.) to 200 mm (8 in.), the mix remaining cohesive (see Fig. 2.17).  (The improvement in workability is smaller at high temperatures.) The resulting concrete  can be placed with little or no compaction and is not subject to excessive bleeding or  segregat
ion. Such concrete is termed flowing concrete and is useful for placing in very  heavily reinforced sections, in inaccessible areas, in floor or road slabs and also  where very rapid placing is desired. Some comments on mix design of flowing concrete  are made on page 709. It should be remembered when designing formwork that flowing  concrete can 
exert full hydrostatic pressure.    Fig. 2.17. Relation between German flow table spread and water content of concrete  with and without superplasticizer     The second use of superplasticizers is in the production of concrete of normal  workability but with an extremely high strength owing to a very substantial reduction  in the water / cement ra
tio. Water / cement ratios down to 0.28 have been used with 28-day  strengths of the order of 100 MPa (15 000 psi). The long-term strength is unimpaired, test  results being available up to 13 years. Generally speaking, superplasticizers can reduce  the water content for a given workability by 25 to 35 per cent (compared with half that  value in t
he case of conventional water-reducing admixtures), and increase the 24-hour  strength by 50 to 75 per cent and even greater increase occurs at earlier ages.  Practical mixes with a strength of 30 MPa (4300 psi) at 7 hours have been obtained  (see Fig. 2.18). With steam-curing or high-pressure steam-curing, even higher strengths  are possible.    
Fig. 2.18. The influence of the addition of superplasticizer on the early strength of  concrete made with a cement content of 370 kg / m3 (630 lb / yd3) and cast at room  temperature. Type III cement; all concretes of the same workability     When the strength at later ages is of primary importance, superplasticizers can be used  in concrete with 
partial fly ash replacement of cement.   The plasticizing action of superplasticizers is of short duration (perhaps 10 minutes):  after some 30 to 90 minutes the workability returns to normal. For this reason, the  superplasticizer should be added to the mix immediately prior to placing; usually,  conventional mixing is followed by the addition of
 the superplasticizer and a short  period of additional mixing. In the case of ready-mixed concrete, a 2-minute re-mixing  period is essential. While re-tempering with an additional dose is not recommended  because of the risk of segregation; re-dosing to maintain workability up to 160 minutes  has been used. Superplasticizers can be used at high 
dosages. They do not markedly change  the surface tension of water, their action being the dispersion of cement agglomerates normally  found when cement is suspended in water. These admixtures are thought to be adsorbed on the  surface of cement and of other very fine particles, causing them to become mutually repulsive as a  result of the anionic
 nature of superplasticizers, which causes the cement particles to become  negatively charged.   Superplasticizers do not significantly affect the setting of concrete except that, when  used with cements having a very low C3A content, there may be excessive retardation.  They do not influence shrinkage, creep, modulus of elasticity or resistance t
o  freezing and thawing. Durability on exposure to sulphates is unaffected. The use  of superplasticizers with an air-entraining admixture requires caution as sometimes the  actual amount of entrained air is reduced by the superplasticizer. Specially modified  superplasticizers have been developed and these seem to produce satisfactory  air-entrai
ned concrete with conventional air-entraining agents. The only real disadvantage  of superplasticizers is their relatively high cost, which is between $2 and $6 (fl and  f3) per cubic metre of concrete. This is due to the high manufacturing cost of a product  with a high molecular weight.  12      Constituent Materials    Concrete is composed main
ly of three materials, namely, cement, water and aggregate, and  an additional material, known as an admixture, is sometimes added to modify certain of its  properties. Cement is the chemically active constituent but its reactivity is only brought  into effect on mixing with water. The aggregate plays no part in chemical reactions but its  usefuln
ess arises because it is an economical filler material with good resistance to volume  changes which take place within the concrete after mixing, and it improves the durability of  the concrete.   A typical structure of hardened concrete and the proportions of the constituent  materials encountered in most concrete mixes are shown in figure 12.1. 
In a properly  proportioned and compacted concrete the voids are usually less than 2 per cent. The  properties of concrete in its fresh and hardened state can show large variations depending  on the type, quality and proportions of the constituents, and from the discussion to follow  students should endeavour to appreciate the significance of thos
e properties of the  constituent materials which affect concrete behaviour.    Figure 12.1 Composition of concrete      12.1 Cement    The different cements used for making concrete are finely ground powders and all have the  important property that when mixed with water a chemical reaction (hydration) takes place  which, in time, produces a very 
hard and strong binding medium for the aggregate  particles. In the early stages of hydration, while in its plastic stage, cement mortar gives to  the fresh concrete its cohesive properties.   The different types of cement and the related British Standards, in which  certain physical and chemical requirements are specified, are given in figure 12.
2; methods  of testing cement for various properties are described in BS 4550. Of these, ordinary  Portland cement is the most widely used, the others being used where concretes with  special properties are required.    Figure 12.2 Different types of cement used for making concrete      Portland Cement    Portland cement was developed in 1824 and 
derives its name from Portland limestone in  Dorset because of its close resemblance to this rock after hydration has taken place. The  basic raw materials used in the manufacture of Portland cements are calcium carbonate,  found in calcareous rocks such as limestone or chalk, and silica, alumina and iron oxide  found in argillaceous rocks such as
 clay or shale. Marl, which is a mixture of calcareous  and argillaceous materials, can also be used.    Manufacture    Cement is prepared by first intimately grinding and mixing the raw constituents in certain  proportions, burning this mixture at a very high temperature to produce clinker, and then  grinding it into powder form. Since the clinke
r is formed by diffusion between the solid  particles, intimate mixing of the ingredients is essential if a uniform cement is to be  produced. This mixing may be in a dry or wet state depending on the hardness of the  available rock.   The wet process is used, in general, for the softer materials such as chalk or clay.  Water is added to the propo
rtioned mixture of crushed chalk and clay to produce a slurry  which is eventually led off to a kiln. This is a steel cylinder, with a refractory lining, which  is slightly inclined to the horizontal and rotates continuously about its own axis. It is  usually fired by pulverised coal, although gas or oil may also be used. It may be as large as  3.
5 m in diameter and 150 m long and handle up to 700 t of cement in a day. The slurry is  fed in at the upper end of the kiln and the clinker is discharged at the lower end where fuel  is injected. With its temperature increasing progressively, the slurry undergoes a number of  changes as it travels down the kiln. At 100 C the water is driven off, 
at about  850 C carbon dioxide is given off and at about 1400 C incipient fusion takes place in the  firing zone where calcium silicates and calcium aluminates are formed in the resulting  clinker. The clinker is allowed to cool and then ground, with 1 to 5 per cent gypsum, to  the required fineness. Different types of Portland cement are obtained
 by varying the  proportions of the raw materials, the temperature of burning and the fineness of grinding,  and in some cases by intergrinding the clinker with other  recognised materials such as pulverized-fuel ash (pfa), or granulated blastfurnace slag  (gbfs). Gypsum is added to control the setting of the cement, which would otherwise set  muc
h too quickly for general use. Certain additives may also be introduced for producing  special cements, for example, calcium chloride is added in the manufacture of  extra-rapid-hardening cement.   The dry or semi-dry process is used for the harder rocks such as limestone and  shale. The constituent materials are crushed into powder form and, with
 a minimum  amount of water, passed into an inclined rotating nodulising pan where nodules are  formed. These are known as raw meal. This is fed into a kiln and thereafter the  manufacturing process is similar to the wet process although a much shorter length of kiln  is used. It should be noted that the dry and semi-dry processes are more  energy
 efficient than the wet process.   The grinding of the clinker produces a cement powder which is still hot and this  hot cement is usually allowed to cool before it leaves the cement works.      Basic Characteristics of Portland Cements    Differences in the behaviour of the various Portland cements are determined by their chemical  composition an
d fineness. The effect of these on the physical properties of cement mortars and  concrete are considered here.    Chemical composition    As a result of the chemical changes which take place within the kiln several compounds are formed in  the resulting cement although only four (see table 12.1) are generally considered to be important. A  direct
 determination of the actual proportion of these principal compounds is a very tedious process  and it is more usual to calculate these from the proportions of their oxide constituents, which can be  determined more easily. A typical calculation using Bogue's method is shown in table 12.2. The  limitations on chemical composition specified in Brit
ish Standards for the various main Portland  cements are summarised in table 12.3 with lime saturation factor (LSF) defined as    (formula)    in which each oxide constituent in brackets is expressed as a percentage by weight of the cement.    TABLE 12.1 Main chemical compounds of Portland cements    TABLE 12.2 A typical chemical composition of or
dinary Portland cement    TABLE 12.3 British Standard requirements for the chemical composition of the principal Portland  cements     The two silicates, C3S and C2S, which are the most stable of these compounds, together form  70 to 80 per cent of the constituents in the cement and contribute most to the physical properties of  concrete. When cem
ent comes into contact with water, C3S begins to hydrate rapidly, generating a  considerable amount of heat and making a significant contribution to the development of the early  strength, particularly during the first 14 days. In contrast C2S, which hydrates slowly and is mainly  responsible for the development in strength after about 7 days, may
 be active for a considerable period  of time.  It is generally believed that cements rich in C2S result in greater resistance to chemical  attack and a smaller drying shrinkage than do other Portland cements.  It may be noted from table  12.2 that the C3S and C2S contents are interdependent. The hydration of C3A is extremely  exothermic and takes
 place very quickly, producing little increase in strength after about 24 hours. Of  the four principal compounds tricalcium aluminate, C3A, is the least stable and cements containing  more than 10 per cent of this compound produce concretes which are particularly susceptible to  sulphate attack. Tetracalcium aluminoferrite, C4AF, is of less impor
tance than the other three  compounds when considering the properties of hardened cement mortars or concrete.   From the foregoing, certain conclusions may be drawn concerning the nature of various  cements. The increased rate of strength development of rapid-hardening Portland cement arises from  its generally high C3S content and also from its i
ncreased fineness which, by increasing the specific  surface of the cement, increases the rate at which hydration can occur. The low rate of strength  development of low-heat Portland cement is due to its relatively high C2S content and low C3A and  C3S contents. An exceptionally low C3A content contributes to the increased resistance to sulphate 
 attack of sulphate-resisting cement. It should be noted that while there can be large differences in the  early strength of concretes made with different Portland cements, their final strengths will generally  be very much the same (see chapter 14).    Fineness    The reaction between the water and cement starts on the surface of the cement parti
cles and in  consequence the greater the surface area of a given volume of cement the greater the hydration. It  follows that for a given composition, a fine cement will develop strength and generate heat more  quickly than a coarse cement. It will, of course, also cost more to manufacture as the clinker must be  more finely ground. Fine cements, 
in general, improve the cohesiveness of fresh concrete and can be  effective in reducing the risk of bleeding (see chapter 13), but they increase the tendency for shrinkage  cracking.   Several methods are available for measuring the fineness of cement, for example BS 4550:  Part 3 prescribes a permeability method which is a measure of the resista
nce of a layer of cement to  the passage of air. The measured fineness is an over-all value known as specific surface and is  expressed in square metres per kilogram (m2 kg-1 ). The minimum fineness requirement for Portland  cements specified in British Standards ranges from 225 m2 kg-1 for ordinary Portland cement (OPC)  to 325 m2 kg-1 for rapid-
hardening Portland cement (RHPC). However, cements manufactured in the  United Kingdom are generally much finer than this, typical values for OPC and RHPC being around  325 and 385 m2 kg-1 respectively.    Hydration    The chemical combination of` cement and water, known as hydration, produces a very hard and  strong binding medium for the aggrega
te particles in concrete and is accompanied by the liberation of  heat, normally expressed as joules per gram. The rate of hydration depends on the relative properties  of` silicate and aluminate compounds, the cement fineness and the ambient conditions (particularly  temperature and moisture). The time taken by the main constituents of cement to 
attain 80 per cent  hydration is given in table 12.4 . Factors affecting the rate of hydration have a similar effect on the  liberation of heat. It can be seen from table 12.5 that the heat associated with the hydration of each of  the principal compounds of cement is very different and in consequence cements having different  compositions also ha
ve different heat characteristics (see figure 12.3).    TABLE 12.4 Time taken to achieve 80 per cent hydration of the main compounds  of Portland cement, based on Goetz (1969)    TABLE 12.5 Heat of hydration of the main chemical compounds of Portland cement,  based on Goetz ( 1969)     Figure 12.3 Typical results for the heat evolution at 20 C of 
different Portland cements:  (A) low heat, (B) ordinary and (C) rapid hardening, based on Lea (1970)     Concrete is a poor conductor of heat and the heat generated during hydration  can have undesirable effects on the properties of the hardened concrete as a result of microcracking of  the binding medium. The possible advantages associated with t
he increased rate of hydration may in  these circumstances be outweighed by the loss in durability of the concrete resulting from the  microcracking. Other factors which affect the temperature of the concrete are the size of the structure,  the ambient conditions, the type of formwork and the rate at which concrete is placed. It should be  noted t
hat it is the rate at which heat is generated and not the total liberated heat which in practice  affects the rise in temperature. The heat characteristics must be considered when determining the  suitability of a cement for a given job.    Setting and hardening    Setting and hardening of the cement paste are the main physical characteristics ass
ociated with  hydration of the cement. Hydration results in the formation of a gel around each of the cement  particles and in time these layers of gel grow to the extent that they come into contact with each other.  At this stage the cement paste begins to lose its fluidity. The beginning of a noticeable stiffening in  the cement paste is known a
s the initial set. Further stiffening occurs as the volume of gel increases  and the stage at which this is complete and the final hardening process, responsible for its strength,  commences is known as the final set. The time from the addition of the water to the initial and final  set are known as the setting times (BS 4550: Part 3) and the spec
ific requirements in this respect for  cements are given in the appropriate British Standards. The setting times for some of the more  important Portland cements are given in table 12.6. In practice, when mixes have a higher water  content than that used in the standard tests, the cement paste takes a correspondingly longer time to  set. Setting t
ime is affected by cement composition and fineness, and also, through its influence on the  rate of hydration, by the ambient temperature.    TABLE 12.6 Typical initial and final setting times for the main Portland cements     Two further phenomena are a flash set and a false set. The former takes place in cement with  insufficient gypsum to contr
ol the rapid reaction of C3A with water, this reaction generating a  considerable amount of heat and causing the cement to stiffen within a few minutes after mixing. This  can only be overcome by adding more water and reagitating the mix. The addition of water results in  a reduction in strength. A false set also produces a rapid stiffening of the
 paste but is not accompanied  by excessive heat. In this case remixing the paste without further addition of water causes it to regain  its plasticity and its subsequent setting and hardening characteristics are quite normal. False set is  thought to be the result of intergrinding gypsum with very hot clinker in the initial stages of the  manufac
ture of cement.    Strength    The strength of hardened cement is generally its most important property. The British Standard  strength requirements for Portland cements, obtained from mortar or concrete tests carried out in  accordance with BS 4550: Part 3, are summarised in table 12.7. It should be understood that cement  paste alone is not used
 for this test because of the unacceptably large variations of strength thus  obtained. Standard aggregates are used for making prescribed mortar or concrete test mixes to  eliminate aggregate effects from the measured strength of the cement.    TABLE 12.7 British Standard requirements for strength of the principal Portland cements    Soundness   
 An excessive change in volume, particularly expansions of a cement paste after setting indicates that  the cement is unsound and not suitable for the manufacture of concrete. In general, the effects of`  using unsound cement may not be apparent for some considerable period of` time but usually manifest  themselves in cracking and disintegration o
f` the surface of` the concrete.  One of` the methods for  testing the soundness of cements is that developed by Le Chatelier as described in BS 4550: Part 3.  The British Standard limitations specified for various Portland cements require that the measured  expansion in this test be not more than 10 mm.      Types of Cement    The different  type
s	of cement are shown in figure	12.2 and their main properties are summarised in  table 12.8.  A brief description of typical properties of each type of cement is given here. For more  detailed information the reader is referred to Neville (1986), Orchard (1979) and Harrison and  Spooner (1986).    TABLE 12.8 Main properties of different cements  
	Portland cements    Ordinary Portland cement has a medium rate of hardening, making it suitable for most concrete work.  It has, however, a low resistance to chemical attacks. Rapid-hardening Portland cement is in many  ways similar to ordinary Portland cement but produces a much higher early strength. The increased  rate of hydration is accompa
nied by a high rate of heat development which makes it unsuitable for  large masses of concrete, although this may be used to advantage in cold weather. Low-heat Portland  cement has a limited use but is suitable for very large structures, such as concrete dams, where the use  of ordinary cement would result in unacceptably large temperature gradi
ents within the concrete. Its  slow rate of hydration is accompanied by a much slower rate of increase in strength than for ordinary  Portland cement although its final strength is very similar. Its resistance to chemical attack is greater  than that of ordinary Portland cements. Sulphate-resisting Portland cement, except for its high  resistance 
to sulphate attack, has principal properties similar to those of ordinary Portland cement.  Calcium chloride should not be used with this cement as it reduces its resistance to sulphate attack.  Extra-rapid-hardening Portland cement is used when very high early strength is required or for  concreting in cold conditions. Because of its rapid settin
g	and hardening properties the concrete should  be placed and compacted within about 30 minutes of mixing. Since the cement contains approximately  2 per cent calcium chloride dry storage is essential. Its use in reinforced or prestressed concrete is not  recommended (BS 8110: Part 1). Ultra-high early-strength Portland cement, apart from its much
	greater fineness and larger gypsum content, is similar in composition to ordinary Portland cement.  Although the early development in strength is considerably higher than with rapid-hardening cement  there is little increase after 28 days. It is suitable for reinforced and prestressed concrete work. White  and coloured Portland cements are simil
ar in basic properties to ordinary Portland cement. White  cement requires special manufacturing methods, using raw materials containing less than 1 per cent  iron oxide. Coloured cements are produced by intergrinding a chemically inert pigment with ordinary  clinker. Because of its inert characteristics, the presence of a pigment slightly reduces
 the concrete  strength. These cements are used for architectural purposes. Hydrophobic Portland cement, owing to  the presence of a water-repellent film around its grain, can be stored under unfavourable conditions of  humidity for a long period of time without any significant deterioration The protective coating is  broken off during mixing and 
normal hydration then takes place. Waterproof and water-repellent  Portland cements produce a more impermeable fully compacted concrete than ordinary Portland  cement. Air-entraining Portland cement produces concrete with a greater resistance to frost attack. The  cement is produced by intergrinding an air-entraining agent with ordinary clinker du
ring  manufacture. However, in practice, it is more advantageous to add an air-entraining agent during  mixing since its quantity can be varied to meet particular requirements.    Slag cements    Different types of slag cement can be produced by intergrinding varying proportions of granulated  blastfurnace slag with activators such as ordinary Por
tland cement and gypsum. The chemical  composition of the granulated slag is similar to that of Portland cement but the proportions are  different. One general requirement is that the slag used must have a high lime content. Of the four  slag cements listed in figure 12.2 only Portland blastfurnace cement has been used to any great extent.   Portl
and blastfurnace cement is produced by mixing up to 65 per cent granulated blastfurnace  slag  with ordinary Portland cement. The basic characteristics are similar to those of ordinary  Portland cement although the rate of hydration is lower. It is particularly suited for structures  involving large masses of concrete. Its resistance to chemical a
ttack, particularly seawater, is  somewhat better than that of ordinary Portland cement. Low-heat Portland blastfurnace cement, except  for its greater slag content, is similar to Portland blastfurnace cement and can therefore be effectively  employed where a control in the rise of temperature is the main requirement. It has a greater resistance  
to chemical attack than Portland blastfurnace and ordinary Portland cements. Supersulphated cement  contains up to 85 per cent slag, 10 to 15 per cent gypsum and a small percentage of ordinary Portland  cement. Its resistance to chemical attack is similar to that of sulphate-resisting Portland cement.  Because of its low heat properties it can als
o	be used for mass concrete work. In order to prevent the  friable and dusty surface appearance often associated with this cement, careful initial curing is  required.    Pozzolanic cements    Pozzolanic cements are made by grinding together a pozzolanic material with ordinary Portland  cement clinker. Pozzolanic materials contain alumino-silicate
 glass, which reacts with the lime  (calcium hydroxide) released by the hydration of the cement and forms cementitious materials.  Pozzolanas occur naturally, for example, volcanic ash, or are obtained in other forms such as  pulverized-fuel ash (pfa), also known as fly ash. Pulverized-fuel ash is a finely divided powder  extracted from the flue g
ases produced in the pulverised-coal fired furnaces used for the generation of  electricity at modern power stations. It is now increasingly used in the manufacture of cements. In the  UK its use is permitted in the manufacture of Portland pulverized-fuel ash cement (BS 6588) with  between 15 to 35 per cent pfa content and Pozzolanic cement with p
ulverized-fuel ash as pozzolana  (BS 6610) with 35 to 50 per cent pfa content. The use of these cements reduces the water demand of a  concrete mix, when compared with that using the corresponding ordinary Portland cement. The rate  of gain in strength and liberation of heat is lower than for ordinary Portland cement and this can be  useful for ma
ss concrete work; the use of pozzolanic cements (BS 6610) is particularly beneficial in  this respect. However, an equal 28 day strength can generally be achieved by increasing the cement  content of a Portland pulverized-fuel ash cement concrete by approximately 10 per cent compared  with that of an ordinary Portland cement mix. Like sulphate res
isting Portland cement, the pozzolanic  cements have a high resistance to chemical attack.    High-alumina cement    This cement, which is manufactured by melting a mixture of limestone or chalk and bauxite  (aluminium ore) at about 1450 C and then grinding the cold mass, has different composition and  properties from those of Portland cements. Th
e high proportion of aluminate, about 40 per cent,  brings about an exceptionally high early strength and consequently it often becomes necessary to keep  concrete, in which this cement is used, continuously wet for at least 24 hours to avoid damage from  the associated heat of hydration.   Structural concrete made with high-alumina cement (common
ly known as HAC) presents a  serious problem however, as it can suffer a substantial reduction in its strength in most normal  environments, owing to the conversion of the hydrated cement to a more porous form. Thus, the use of  high-alumina cement in structural concrete is not permitted in BS 8110.   The conversion of high-alumina cement is parti
cularly sensitive to temperature, the  water - cement ratio and the richness of the mix at a given water - cement ratio, increasing as all these  factors increase. In certain circumstances, the residual strength of concrete can be considerably lower  than its one day strength. Water - cement ratios higher than 0.4 should generally be avoided.   Th
e exceptionally high early strength property of high-alumina cement makes it well suited  for repair work of limited life and for temporary works. It also has a wide application in refractory  concrete and its resistance to chemical attack, particularly sulphate attack, is greater than that of  Portland cements, although in its converted, more por
ous, state it is very susceptible to alkali and  sulphate attack.      1 2.2 Aggregate    Aggregate is much cheaper than cement and maximum economy is obtained by using as much  aggregate as possible in concrete. Its use also considerably improves both the volume stability and the  durability of the resulting concrete. The commonly held view that 
aggregate is a completely inert filler  in concrete is not true, its physical characteristics and in some cases its chemical composition  affecting to a varying degree the properties of concrete in both its plastic and hardened states.    Basic Characteristics of Aggregate    The criterion for a good aggregate is that it should produce the desired
 properties in both the fresh  and hardened concrete. In testing aggregates it is important that a truly representative sample is used.  The procedure for obtaining such a test sample is described in BS 812: Part 102.    Physical properties    The properties of the aggregate known to have a significant effect on concrete behaviour are its  strengt
h, deformation, durability, toughness, hardness, volume change, porosity, relative density and  chemical reactivity.   The strength of an aggregate limits the attainable strength of concrete only when its  compressive strength is less than or of the same order as the design strength of concrete. In practice  the majority of rock aggregates used ar
e usually considerably stronger than concrete. While the  strength of concrete does not normally exceed 80 N mm-2 and is generally between 30 and  50 N mm-2 the strength of the aggregates commonly used is in the range 70 to 350 N mm-2, In  general, igneous rocks are very much stronger than sedimentary and metamorphic rocks. Because of the  irregul
ar size and shape of aggregate particles a direct measurement of their strength properties is not  possible. These are normally assessed from compressive strength tests on cylindrical specimens taken  from the parent rock and from crushing value tests on the bulk aggregate. For weaker materials, that  is, those with crushing values greater than 30
, the crushing value may be unreliable and the load  required to produce 10 per cent fines in the crushing test should be used (BS 812). It should be noted  that the strength test on rock specimens was deleted from BS 812: Part 3 in 1975 and is consequently  less used than the tests on bulk aggregate. The results of these tests for the strength pr
operties of  aggregates are only a guide to aggregate quality, however, which may also be assessed from the  intensity of aggregate fracturing in ruptured concrete specimens.   The deformation characteristics of an aggregate are seldom considered in assessing its  suitability for concrete work although they can easily be determined from compressio
n tests on  specimens from the parent rock. In general, the modulus of elasticity of concrete increases with  increasing aggregate modulus. The deformation characteristics of the aggregate also play an  important part in the creep and shrinkage properties of concrete as the restraint afforded by the  aggregate to the creep and shrinkage of the cem
ent paste depends on their relative moduli of  elasticity.   A commonly used definition for aggregate toughness is its resistance to failure by impact and  this is normally determined from the aggregate impact test (BS 812: Part 3). The aggregate impact  value (AIV) has a close linear correlation with the aggregate crushing value (ACV) and can the
refore  be employed as the test for assessing aggregate strength (Spence et al ., 1974). The AIV test has  advantages over the ACV test of simplicity and cheapness in operation. It does not require the more  elaborate facilities of a testing laboratory, the equipment is portable and the small sample required  makes it a particularly useful test fo
r quality control purposes. Hardness is the resistance of an  aggregate to wear and is normally determined by an abrasion test (BS 812: Part 3). Toughness and  hardness properties of an aggregate are particularly important for concrete used in road pavements.   Volume changes due to moisture movements in aggregates derived from sandstones,  greywa
ckes and some basalts may result in considerable shrinkage of the concrete (BRE Digest 35,  1968, and Dhir et al ,1978). If the concrete is restrained this produces internal tensile stresses,  possible tensile cracking and subsequent deterioration of the concrete. If the coefficient of thermal  expansion of an aggregate differs considerably from t
hat of the cement paste this too may adversely  affect the concrete performance.   Aggregate porosity is an important property since it affects the behaviour of both freshly  mixed and hardened concrete through its effect on the strength, water absorption and permeability of  the aggregate. An aggregate with high porosity will tend to produce a le
ss durable concrete,  particularly when subjected to freezing and thawing, than an aggregate with low porosity. Direct  measurement of porosity is difficult and in practice a related property, namely, water absorption, is  measured. The water absorption is defined as the weight of water absorbed by a dry aggregate in  reaching a saturated surface-
dry state and is expressed as a percentage of the weight of the dry  aggregate. In general, sedimentary rock materials have the highest absorption values. A gravel  aggregate with a similar petrology to that of a crushed-rock aggregate will absorb more water because  of its greater weathering. It should be noted that, for a given rock type, absorp
tion can vary depending  on the way in which it is measured and also on the size of the aggregate particles.   The total moisture content, that is, the absorbed moisture plus the free or surface water, of  aggregates used for making concrete varies considerably. The water added at the mixer must be  adjusted to take account of this if the free wat
er content is to be kept constant and the required  workability and strength of concrete maintained. Concrete mix proportions are normally based on the  weight of the aggregates in their saturated surface-dry condition and any change in their moisture  content must be reflected in adjustments to the weights of the aggregates used in the mix. Sever
al  methods for the determination of moisture content and absorption are described in BS 812: Part 2.   The relative density of a material is the ratio of its unit weight to that of water. Since  aggregates incorporate pores the value of relative density varies depending on the extent to which the  pores contain (absorbed) water when the value is 
determined (BS 812: Part 2). For the purposes of  mix design the relative density on a 'saturated and surface-dry' basis is used. This is given by  A / (A - B) where A is the weight of the saturated surface-dry sample in air and B the weight of the  saturated sample in water. The relative density of most natural aggregates falls within the range  
2.5 - 3.0. For artificial aggregates the relative density varies over a much wider range. Aggregate is the  major constituent of concrete and as such its relative density is an important factor affecting the  density of the resulting concrete.    Shape and surface texture    Aggregate shape and surface texture can affect the properties of concrete
 in both its plastic and  hardened states. These external characteristics may be assessed by observation of the aggregate  particles and classification of their particle shape and texture in accordance with tables 12.9 and 12.10  from BS 812: Part 1. The classification is somewhat subjective , however, and the particle shape may  also be assessed 
by direct measurement of the aggregate particles to determine the flakiness,  elongation and angularity.    TABLE 12.9 Shape of aggregates    TABLE 12.10 Surface texture of aggregates     The angularity is expressed in terms of the angularity number. This is the difference between  the solid volume of rounded aggregate particles after compaction i
n a standard cylinder, expressed as  a percentage of the volume of the cylinder, and the solid volume of the particular aggregate being  investigated when compacted in a similar manner. It is a measure of the increased voids in the  compacted nonrounded aggregate, expressed as a percentage of the total volume. The angularity  number ranges from ze
ro for a perfectly rounded aggregate to about 12. Hughes ( 1966) proposed an  alternative method for determining shape based on an angularity factor. This is defined as the ratio of  the solid volume of loose single size spheres to the solid volume of a single-size aggregate of similar  nominal dimensions placed under identical conditions. The met
hod has been shown to be suitable for  both fine and coarse aggregates.    Grading    The grading of an aggregate defines the proportions of particles of different size in the aggregate. The  size of the aggregate particles normally used in concrete varies from 37.5 to 0.15 mm. BS 882 places  aggregates in three main categories: fine aggregate or 
sand containing particles the majority of which  are smaller than 5.00 mm, coarse aggregate containing particles the majority of which are larger than  5.00 mm and all-in aggregate comprising both fine and coarse aggregate.     The grading of an aggregate can have a considerable effect on the workability and stability of a  concrete mix (see chapt
er 13) and is a most important factor in concrete mix design. In the United  Kingdom, the grading of natural aggregates is generally required to be within the limits specified in  BS 882. The fine aggregates are generally required to conform to any one of three standard grading  limits, all of which are suitable for concrete provided suitable mix 
proportions are used.   A sieve analysis is used for determining the grading of an aggregate (BS 812: Part 103). The  aggregate sample must be air-dried and the weight of material retained on each sieve should not  exceed the specified maximum values as overloading will give erroneous results. Sieving  may be  performed by hand or machine and the 
results of a typical analysis are shown in table 12.11  from which it will also be seen that successive sieve sizes decrease by a factor of about two. A  convenient visual assessment of the particle size distribution can be obtained from a grading chart  (figure 12.4). Curve A represents a continuously graded aggregate and curve B, in which two of
 the  sizes are missing, represents a gap-graded aggregate.    TABLE 12.11 A typical example of calculation for aggregate grading    Figure 12.4 Typical aggregate grading curves: (A) continuously graded (see table 12.11) and  (B) gap-graded aggregates     In practice, fine and coarse aggregates are batched separately, their proportions being  gove
rned largely by their respective gradings. One method for determining the required aggregate  contents (see chapter 15) uses the grading modulus and the equivalent mean diameter of both fine and  coarse aggregate.   The grading modulus of an aggregate is the mean specific surface of spheres which pass  through the same sieve size as the actual agg
regate particles and whose size distribution, or grading,  corresponds to that of those aggregate particles.   For spheres with constant diameter D, the grading modulus G and the specific surface of a  sphere are identical, that is, (formula).   Consider now the grading modulus of aggregate particles lying wholly between successive  sieve sizes D1
 and D2 corresponding to the smallest and largest diameters of the associated spheres. If,  between successive sieve sizes, the proportion by volume of particles of diameter D is assumed to be  inversely proportional to D1 then the grading modulus is given by    (formula).    The grading moduli for particles between different sieve sizes are given
 in table 15.3.   The grading modulus of an aggregate whose particles cover a range of sieve  sizes is given by    (formula)    where Ga and Gb are the grading moduli of the coarse and fine aggregate respectively.   The equivalent mean diameter of particles lying wholly between successive  sieve sizes D1 and D2 is conveniently assumed (Hughes, 196
0) to be given by D = (D1 + D2) /2. The  equivalent mean diameters for particles between different sieve sizes are given in table 15.3 and  values for aggregates whose particles cover a range of sieve sizes are obtained in the same way as the  grading modulus (see chapter 1 5).      Types of Aggregate    In the previous sections discussion has bee
n mainly confined to rock aggregates. Although other types  of aggregate are used for making concrete their contribution is very small in comparison with rock  aggregates. The general classifications of aggregates and the related British Standards are shown in  figure 12.5. For a more detailed discussion the reader is referred to Taylor (1965), Sh
ort and  Kinniburgh (1978) and BRE Digests 111 and 123.    Figure 12.5 A classification of aggregates used for making concrete    Heavyweight aggregate    Heavyweight aggregates provide an effective and economical use of concrete for radiation shielding,  by giving the necessary protection against X-rays, gamma-rays and neutrons, and for weight co
ating  of submerged pipelines. The effectiveness of heavyweight concrete, with a density from 4000 to 8500  kg m-3, depends on the aggregate type, the dimensions and the degree of compaction. It is frequently  difficult with heavyweight aggregates to obtain a mix which is both workable and not prone to  segregation.    Normal aggregate    These ag
gregates are suitable for most purposes and produce concrete with a density in the range  2300 to 2500 kg m-3. Rock aggregates are obtained by crushing quarried rock to the required particle  size or by extracting the sand and gravel deposits formed by alluvial or glacial action. Some sands and  gravels are also obtained by dredging from sea and r
iver beds. Aggregates, in particular sands and  gravels, should be washed to remove impurities such as clay and silt. In the case of river and marine  aggregates the chloride content should generally be less than 1 per cent if these are to be used for  structural concrete. The properties of rock aggregates depend on their composition, grain size a
nd  texture. For example, granite has a low fire resistance because of the high coefficient of expansion of  its quartz content. Sandstone aggregates generally produce concretes with a high drying shrinkage  because of their high porosity. Crushed aggregates tend to be angular and gravels irregular or rounded  in shape.   Air-cooled blastfurnace s
lag aggregates produce concretes with similar strength to natural  aggregates but with improved fire resistance. Broken-brick aggregates are also very fire resistant, but  should not be used for normal concrete if its soluble sulphate content exceeds 1 per cent.    Lightweight aggregate    Lightweight aggregates find application in a wide variety 
of concrete products ranging from  insulating screeds to reinforced or prestressed concrete although their greatest use has been in the  manufacture of precast concrete blocks. Concretes made with lightweight aggregates have good fire  resistance properties. The most commonly used artificial lightweight aggregates in the United  Kingdom are sinter
ed shale (Aglite), foamed slag, expanded clay (Leca), sintered pulverised fuel ash  (Lytag and Taclite), pelletised expanded slag (Pellite and Lycrete) and sintered colliery shale (Sintag).  They are highly porous and absorb considerably greater quantities of water than do normal  aggregates. For this reason they should normally be batched by volu
me owing to the large variations  that can occur in their moisture content. Their bulk density normally ranges from 350 to 850 kg m-3  for coarse aggregates and from 750 to 1100 kg m-3 for fine aggregates. The methods for testing  lightweight aggregates are described in BS 3681.      12.3 Water    Water used in concrete, in addition to reacting wi
th cement and thus causing it to set and harden, also  facilitates mixing, placing and compacting of the fresh concrete. It is also used for washing the  aggregates and for curing purposes. The effect of water content on the properties of fresh and  hardened concrete is discussed in chapters 13 and 14. In general water fit for drinking, such as ta
p  water, is acceptable for mixing concrete. The impurities that are likely to have an adverse effect when  present in appreciable quantities include silt, clay, acids, alkalis and other salts, organic matter and  sewage. The use of seawater does not appear to have any adverse effect on the strength and durability  of Portland cement concrete but 
it is known to cause surface dampness, efflorescence and staining and  should be avoided where concrete with a good appearance is required. Seawater also increases the risk  of corrosion of steel and its use in reinforced concrete is not recommended. When the suitability of  mixing water is in question, it is desirable to test for both the nature 
and extent of contamination as  prescribed in BS 3148. The quality of water may also be assessed by comparing the setting time and  soundness of cement pastes made with water of known quality and the water whose quality is suspect.   The use of impure water for washing aggregates can adversely affect strength and durability if  it deposits harmful
 substances on the surface of the particles. In general, the presence of impurities in  the curing water does not have any harmful effects, although it may spoil the appearance of concrete.  Water containing appreciable amounts of acid or organic materials should be avoided.      12.4 Admixtures    Admixtures are substances introduced into a batch
 of concrete, during or immediately before its  mixing, in order to alter or improve the properties of the fresh or hardened concrete or both. Although  certain finely divided solids, such as pozzolanas and slags, fall within the above broad definition of  admixtures they are distinctly different from what is commonly regarded as the main stream o
f  admixtures and therefore should be treated separately . It should also be noted that the materials used  by the cement manufacturers to modify the properties of cement are normally described as additives.   In general, the changes brought about in the concrete by the use of admixtures are effected  through the influence of the admixtures on hyd
ration, liberation of heat, formation of pores and the  development of the gel structure. Concrete admixtures should only be considered for use when the  required modifications cannot be made by varying the composition and proportion of the basic  constituent materials, or when the admixtures can produce the required effects more economically.   S
ince admixtures may also have detrimental effects, their suitability for a particular concrete  should be carefully evaluated before use, based on a knowledge of their main active ingredients, on  available performance data and on trial mixes. The specific effects of an admixture generally vary with  the type of cement, mix composition, ambient co
nditions (particularly temperature) and its dosage.  Since the quantity of admixture used is both small and critical the required dose must be carefully  determined and administered. Where related British Standards exist (BS 5075) the admixtures should  comply with their specifications such as those for acceptance test requirements given in table 
12.12. It  should be remembered that admixtures are not intended to replace good concreting practice and should  not be used indiscriminately.    TABLE 12.12 Admixture acceptance test requirements as specifies in BS 5075      Types of Admixture    Several hundred proprietary admixtures are available and since a great many usually contain several  
chemicals intended simultaneously to change several properties of concrete, they are not easy to  classify. Moreover, as many of the individual constituents and their proportions are not widely known  the selection of an admixture must frequently be based on the information provided by the suppliers.  Different types of admixture are listed in tab
le 12.13 and for a more detailed discussion on these  admixtures the reader is referred to Rixom (1978) and the Concrete Society Technical Report No. 18  (1980).    TABLE 12.13 Types of admixture    Air-entraining agents    These are probably the most important group of admixtures. They improve the durability of concrete,  in particular its resist
ance to the effects of frost and deicing salts. The entrainment of air in the form  of uniformly dispersed, very small and stable bubbles of predominantly between 0.25 and 1 mm  diameter can be achieved by using foaming agents based on natural resins, animal or vegetable fats  and synthetic detergents which promote the formation of air bubbles dur
ing mixing or by using  gas-producing chemicals such as zinc or aluminium powder which react with cement to produce gas  bubbles. The first method is generally more effective and is the most widely used. The beneficial  effects of entrained air are produced in two ways: first, by disrupting the continuity of capillary pores  and thus reducing the 
permeability of concrete and second, by reducing the internal stresses caused by  the expansion of water on freezing.   Air-entraining agents also improve the workability and cohesiveness of fresh concrete and tend  to reduce bleeding and segregation; this is particularly useful when aggregates with poor gradings are  used. However, entrained air 
results in some reduction in concrete strength. Since improvements in  workability can permit a reduction in the water content the loss in strength can be minimised. The  amount of entrained air in concrete is dependent on the type of admixture and dosage used, as well as  on the cement type, aggregate type and grading, mix proportions, ambient te
mperature, type of mixer  and mixing time. Thus, it should only be used when adequate supervision is assured. The density of  air-entrained concrete is reduced in direct proportion to the amount of air entrained. In the United  Kingdom, air-entraining admixtures are required to comply with BS 5075: Part 2 (see table 12.12)  and a method for determ
ining the amount of air entrained in concrete is described in BS 1881: Part  106.    Accelerating agents    These can be divided into two groups, namely, setting accelerators and setting and hardening  accelerators. The first of these are alkaline solutions which can considerably reduce the setting time  and are particularly suitable for repair wo
rk involving water leakage. Because of their adverse effect  on subsequent strength development these admixtures should not be used where the final concrete  strength is an important consideration. Setting and hardening accelerators increase the rate of both  setting and early strength development. The most common admixture for this purpose is cal
cium  chloride. Since its use may result in several adverse effects such as increased drying shrinkage,  reduced resistance to sulphate attack and increased risk of corrosion of steel reinforcement, it should  only be used with extreme caution and in accordance with any relevant specifications. Indeed, in line  with many other countries, the use o
f calcium chloride in concrete containing steel reinforcement is  no longer permitted in the United Kingdom (BS 8110). This has resulted in a new breed of chloride-free  accelerating admixtures for use in reinforced concrete (McCurrich et al., 1979). These admixtures  are based on calcium formate, sometimes blended with corrosion inhibitors such a
s soluble nitrates,  benzoates and chromates. However, calcium chloride remains a most effective and economical  material for use in plain concrete, particularly under winter conditions, for emergency repair work, or  where early removal of formwork is required.  The accelerating admixtures are required to comply with BS 5075: Part 1 (see table 12
.12) and their  effect on concrete is dependent on dosage, cement type and mix proportions used and the ambient  temperature of concrete. There are many accelerators which can achieve an increase in 1 day strength  of up to 100 per cent over the corresponding normal mix, which is well in excess of the value  specified in BS 5075: Part 1 (see table
 12.12).    Retarders    Most admixtures in this group are based on lignosulphonic or hydroxylated-carboxylic acids and their  salts with cellulose or starch. They are used mainly in hot countries where high temperatures can  reduce the normal setting and hardening times. Other notable applications include situations where  large concrete pours, s
liding formwork, or ready-mixed concrete is used. The extended setting time  prevents the formation of cold joints, allows time for steel-fixing and sometimes for overnight  stopping of the formwork and compensates for the time lost in transit of ready-mixed concrete. A  slightly reduced water content may be used when using these retarding agents,
 with a corresponding  increase in final concrete strength. The lignin-based retarders result in some air-entrainment and tend  to increase cohesiveness and reduce bleeding although drying shrinkage may be increased . The  hydroxy-carboxylic retarders, however, tend to increase bleeding.   The use of retarders on their own is declining however, wi
th retarding water reducing  admixtures gaining popularity. The effect of retarders is dependent on dosage, cement type  and mix proportions used, as well as the time of addition and the ambient temperature. The  effectiveness of retarding water-reducing admixtures is also influenced by aggregate type and grading.  In the United Kingdom, both thes
e admixtures are required to comply with BS 5075: Part 1 (see table  12.12).    Water reducers or plasticizers    These admixtures are also based on lignosulphonic and hydroxylated-carboxylic acids. Their effect is  thought to be due to an increased dispersion of cement particles causing a reduction in the viscosity of  the concrete. This effect c
an be used in three ways. First, to increase concrete workability for a given  water - cement ratio and nominal strength; this allows easier placing and compaction of concrete.  Second, to increase concrete strength without the addition of further cement owing to the reduced  water requirement of a mix at a given workability; this allows the produ
ction of high strength  concrete. Third, to reduce cement content of a mix at a given workability and strength by reducing its  water content while maintaining the original water - cement ratio; this reduces the cost of a mix. The  reduction in cement content will result in lower maximum temperatures and hence reduce the risk of  shrinkage crackin
g.   The effectiveness of a water-reducing admixture at a given dosage is dependent on cement  type, aggregate type and grading, mix proportions, and ambient temperature. At a normal dose of  admixture, water and cement reductions up to about 10 per cent can be achieved. The British  specifications for this admixture are given in BS 5075: Part 1 (
see table 12.12).    Superplasticizers    These are a relatively new category of water-reducing admixtures, and are most effective in dispersing  cement particles and thus in increasing the concrete fluidity. The superplasticizers, which are mainly  based on sulphonated melamine formaldehyde condensates or sulphonated naphthalene formaldehyde  con
densates, do not have the problem of retardation and excessive air entrainment associated with  high rates of addition of normal plasticizers.  The superplasticizers are commonly used to produce flowing concrete, defined as having slump in  excess of 200 mm or a flow value within 510 to 620 mm (BS 5075: Part 3), without having to change  the origi
nal mix composition and without causing a strength reduction. The self-levelling property of  flowing concrete means that it can be placed with little or no compaction and is therefore particularly  suitable for heavily reinforced and inaccessible sections, or where rapid placement of concrete is  required. To avoid bleeding, segregation and other
 adverse effects which tend to occur in  high workability mixes it is important that the flowing concrete mix design, production and handling  are all properly controlled. It should also be noted that the flowing characteristics of a mix are  retained only for a short period of time, about 30 minutes after the addition of the superplasticizer,  an
d for this reason superplasticizers should be added to the mix immediately prior to its placing.  Although it is claimed that the properties of flowing concrete in its hardened state are comparable to  those of the corresponding original mix, recent research has shown that this is not always so (Dhir  and Yap, 1984) and the differences should be t
aken into account at the design stage of the structure.  An alternative use of superplasticizers is in the production of high strength concrete by reducing the  mix water content, and hence water - cement ratio, while maintaining the original workability. A  reduction in the water content up to 25 per cent is possible.    Bonding admixtures    The
se are organic polymer emulsions used to enhance the bonding properties of concrete,  particularly for patching and remedial work. The bonding admixtures are known also to  increase the abrasion resistance of concrete and its tensile strength, but some reduction in  compressive strength also occurs.    Water-repelling agents    These are the least
-effective of all admixtures and are based on metallic soaps or vegetable  or mineral oils. Their use gives a slight temporary reduction in concrete permeability.    Pigments    Colouring pigments, in powder form, are normally used in concrete for architectural  purposes and the best effect is produced when they are interground with the cement  cl
inker rather than added during mixing. Pigments used for this purpose are formulated  from both natural and synthetic materials and in the United Kingdom should conform to  BS 1014. Pigments do not normally affect the concrete properties although those based on  carbon may cause some loss of strength at early ages and can also reduce the effective
ness  of air-entraining admixtures.    Pore fillers    These are chemically inactive finely ground materials such as bentonite, kaolin or rock  flow. These admixtures are thought to improve workability, stability and impermeability of  concrete, and are used with poorly graded aggregates, but their use may result in some  reduction in concrete str
ength.    Pozzolanas    The most commonly used pozzolanas are pumicite and pulverized-fuel ash (pfa). Because  of their reaction with lime, which is liberated during the hydration of Portland cement,  these materials can improve the durability when added to concrete. Since they often retard  the rate of setting and hardening and thus the rate of h
eat evolution, they can be useful in  mass concrete work. Pulverized-fuel ash (see Pozzolanic cements), the most common  artificial pozzolana, can be divided into two generic types dependent on the parent coal  source. Lowlime pfa, generally containing less than 10 per cent CaO, is derived from  anthracite and bituminous coals, whereas high-lime p
fa, generally containing more than 10  per cent CaO, is derived from sub-bituminous and lignitic coals. In the UK, only low-lime  pfa is generally produced. The main difference between the two types of pfa is that the  high-lime pfa is intrinsically hydraulic, although in other respects its activity is similar to  the low-lime pfa. The use of pfa 
in concrete as a partial replacement of Portland  cement is now recognised by the various British Standards (for example, BS 5328 and BS  8110) and other specifying authorities such as the Department of Transport (Specification  for Highway Works, 1986). For use in concrete as a cementitious component in the UK,  however, pfa is required to confor
m to BS 3892: Part 1 (see table 12.14). It should be  noted that, while pfa conforming to BS 3892: Part 2 (table 12.14) may be used in  concrete, it is not considered as a part of the cementitious component.    TABLE 12.14 BS 3892 specification for PFA for use in concrete     The main benefits obtained from the inclusion of pfa in fresh concrete a
re  increased workability, decreased bleeding and reduced tendency to segregation. In the  hardened concrete, pfa promotes a denser matrix structure and gives good long-term  strength development. For detailed information on the mechanisms by which pfa affects the  fresh and hardened concrete, the reader is referred to Dhir (1986), which also prov
ides  information concerning the characterisation of pfa and the mix proportioning, engineering  properties and durability of concrete containing pfa. Briefly, a concrete  mix incorporating pfa can be designed for a given workability and strength comparable  with the corresponding mix in which Portland cement alone is used, provided the  variation
s in pfa characteristics (as in the case of Portland cement) are taken into account.  A concrete mix designed in this manner can generally be expected to be similar to or better  than the corresponding Portland cement concrete as a construction material.   The use of pfa in most structural grade concrete mixes can result in a saving in  the Portla
nd cement content of up to 100 kg m-3, depending upon the quality of pfa and  Portland cement. Since pulverized-fuel ash is considerably cheaper than the cement, its use  in concrete will result in direct economic savings, although indirect savings resulting from  the conservation of energy (owing to reduced cement production) and environment (owi
ng  to pfa consumption rather than disposal) is of greater national interest.  5 Strength of Concrete      Strength of concrete is commonly considered its most valuable property, although in many  practical cases other characteristics, such as durability and impermeability, may in  fact be more important. Nevertheless, strength usually gives an ov
erall picture of the  quality of concrete because strength is directly related to the structure of the  hardened cement paste.   The mechanical strength of cement gel was discussed on Page 35; below, some empirical  relations concerning the strength of concrete will be considered.      Water / Cement Ratio    In engineering practice, the strength 
of concrete at a given age and cured at  a prescribed temperature is assumed to depend primarily on two factors only: the  water / cement ratio and the degree of compaction. The influence of air voids on strength  was discussed on page 204, and at this stage we shall consider fully-compacted concrete only:  in practice this is taken to mean that t
he hardened concrete contains about 1 per cent of air  voids.   When concrete is fully compacted its strength is taken to be inversely proportional to the  water / cement ratio according to the "law" established by Duff Abrams in 1919. He found  strength to be equal to    (formula)    where w / c represents the water / cement ratio of the mix (ori
ginally taken by  volume), and K1, and K2 are empirical constants. A typical strength versus water / cement  ratio curve is shown in Fig. 5.1.    Fig. 5.1. The relation between strength and water / cement ratio of concrete     Abrams' "law", although established independently, is a special case of  a general rule formulated by Feret in 1896. This 
was in the form    (formula)    where f, is the strength of concrete, c, w, and a are the absolute volumes of cement,  water, and air respectively, and K is a constant.   It may be recalled that the water / cement ratio determines the porosity of the hardened  cement paste at any stage of hydration (see p. 32). Thus the water / cement ratio and th
e  degree of compaction both affect the volume of voids in concrete, and this is why the  volume of air in concrete is included in Feret's expression.   The relation between strength and the total volume of voids is not a unique property of  concrete but is found also in other brittle materials in which water leaves behind pores: for  instance, th
e strength of plaster is also a direct function of its void content (see Fig. 5.2).  Moreover, if the strength of different materials is expressed as a fraction of the strength at  a zero porosity, a wide range of materials conform to the same relation between relative  strength and porosity, as shown in Fig. 5.3 for plaster, steel, iron, alumina 
and zirconia.  This general pattern is of interest in understanding the role of voids in the strength of  concrete. Moreover, the relation of Fig. 5.3 makes it clear why cement compacts (see p.  30), which have a very low porosity, have a very high strength.   Strictly speaking, strength of concrete is probably influenced by the volume of all void
s in  concrete: entrapped air, capillary pores, gel pores, and entrained air if present. An  example of the calculation of the total air content may be of interest and is given in a  footnote.    Fig. 5.2. Strength of plaster as a function of its void content    Fig. 5.3. Influence of porosity on relative strength of various materials     The infl
uence of the volume of pores on strength can be expressed by a power function of  the type    (formula),    where  fc = strength of concrete with porosity p   fc,0 = strength at zero porosity   and n = a coefficient, which need not be constant.   However, the shape of the pores is also a factor. The shape of the solid  particles and their modulus 
of elasticity also influence the stress distribution,  and therefore stress concentration, within the concrete.   The possibility of obtaining very high strengths at extremely low  porosities was demonstrated by the application of high pressure (340 MPa  (50 000 psi)) with a simultaneous high temperature (250 C (480 F)) to  cement paste (see Fig. 
5.4(a)). Porosity of about 1 per cent was obtained.  The degree of hydration ranged between 0.29 and 0.37 but did not seem to be  related to strength. The specimens had a compressive  strength of about 660 MPa (95 000 psi) and tensile splitting strength of 64 MPa (9250  psi). These and other tests show that there is a direct relation between stren
gth and  porosity, although the exact form of this relation has not been established: specifically, it is  not clear whether the logarithm of porosity is linearly related to strength or to its logarithm  (compare Fig. 5.4(a) and (b)).     Fig. 5.4(a). Relation between compressive strength and logarithm of porosity of  cement paste compacts for var
ious treatments of pressure and high temperature    Fig. 5.4(b). Relation between logarithm of compressive strength and logarithm of porosity  of cement paste compacts for various treatments of pressure and high temperature (after  ref. 5.107)    Fig. 5.1 shows that the range of validity of the water / cement ratio rule is limited. At the  lower e
nd of the scale the curve ceases to be followed when full compaction is no longer  possible; the actual position of the point of departure depends on the means of compaction  available. It seems also that mixes with a very low water / cement ratio and an extremely  high cement content (470 to 530 kg / m3 (800 to 900 lb / yd3)) exhibit retrogressio
n of  strength, particularly when large size aggregate is used. Thus at later ages, in this type of  mix, a lower water / cement ratio would not lead to a higher strength. This behaviour may  be due to stresses induced by shrinkage, whose restraint by aggregate particles causes  cracking of the cement paste or a loss of the cement - aggregate bond
.   From time to time the water / cement ratio rule has been criticized as not being sufficiently  fundamental (see the following section). Nevertheless, in practice the water / cement ratio  is the largest single factor in the strength of fully compacted concrete. Perhaps the best  statement of the situation is that by Gilkey:   "For a given ceme
nt and acceptable aggregates, the strength that may be  developed by a workable, properly placed mixture of cement, aggregate, and water (under  the same mixing, curing, and testing conditions) is influenced by the:   (a) ratio of cement to mixing water   (b) ratio of cement to aggregate   (c) grading, surface texture, shape, strength, and stiffne
ss of aggregate particles   (d) maximum size of the aggregate"   We can add that factors (b) to (d) are of lesser importance than factor (a)  when usual aggregates up to 40 mm (1 1/2 in.) maximum size are employed They are,  nevertheless, present since, as pointed out by Walker and Bloem, the strength of concrete  results from: (1) the strength of
 the mortar; (2) the bond between the mortar and the  coarse aggregate; and (3) the strength of the coarse aggregate particle, i.e., its ability to  resist the stresses applied to it".   Fig. 5.5. shows that the graph of strength versus water / cement ratio is   approximately in the shape of a hyperbola. This applies to concrete made with any give
n  type of aggregate and at any given age. It is a geometrical  property of a hyperbola  y = k / x that y against 1 / x plots as a straight line. Thus the relation between the strength  and the cement / water ratio is approximately linear in the range of cement / water ratios  between about 1 2 and 2 5. This linear relationship is clearly more con
venient to use than  the water / cement ratio curve, particularly when interpolation is desired. Fig. 5.6 shows  the data of Fig. 5.5 plotted with the cement / water ratio as abscissa. The values used apply  to the given cement only, and in any practical case the actual relation between strength and  cement / water ratio has to be determined.    F
ig. 5.5. Relation between 7-day strength and water / cement ratio for concrete made with  a rapid-hardening Portland cement    Fig. 5.6. A plot of strength against cement / water ratio for the data of Fig. 5.5    It must be admitted that the relations discussed here are not precise, and other  approximations can be made. For instance, it has been 
suggested that as an approximation  the relation between the logarithm of strength and the natural value of the water / cement  ratio can be assumed to be linear (cf. Abrams' expression). As an illustration, Fig. 5.7 gives  the relative strength of mixes with different water / cement ratios, taking the strength at the  water / cement ratio of 0 4 
as unity.   The pattern of strength of high-alumina cement concrete is somewhat different from that  of concrete made with Portland cement, in that strength increases with the cement / water  ratio at a progressively decreasing rate.    Fig. 5.7. Relation between logarithm of strength and water / cement ratio      Gel / Space Ratio     The influen
ce of the water / cement ratio on strength does not truly constitute a law as the  water / cement ratio rule does not include many  qualifications necessary for its validity. In  particular, strength at any water / cement ratio depends on the degree of hydration of  cement  and its chemical and physical properties; the temperature at which hydrati
on takes place;  the air content of the concrete; and also the change in the effective water / cement ratio  and the formation of fissures due to bleeding.   It is more correct, therefore, to relate strength to the concentration of the solid products  of hydration of cement in the space available for these products; in this connection it may  be r
elevant to refer again to Fig. 1.9. Powers has determined the relation between the  strength development and the gel / space ratio. This ratio is defined as the ratio of the  volume of the hydrated cement paste to the sum of the volumes of the hydrated cement  and of the capillary pores.   On page 27 it was shown that cement hydrates to occupy mor
e than twice its original  volume; in  the following calculations the products of hydration of 1 ml of cement will be assumed to  occupy 2.06 ml; not all the hydrated material is gel, but as an approximation we can  consider it as such.    Let c = weight of cement,   vc = specific volume of cement,   wo = volume of mixing water,   and a = the frac
tion of cement that has hydrated.  Then, the volume of gel = 2 06 cvca, and the total space available to the gel is cvca + wo.  Hence, the gel / space ratio is    (formula).     Taking the specific volume of dry cement as 0 319 ml / g, the gel / space ratio becomes    (formula).     The compressive strength of concrete tested by Powers was found t
o be 234 x3 MPa  (34 000 x3 psi), and is independent of the age of the concrete or its mix proportions. The actual  relation between the compressive strength of mortar and the gel / space ratio is shown in Fig. 5.8: it  can be seen that strength is approximately proportional to the cube of the gel / space ratio, and the  figure 234 MPa (34 000 psi
) represents the intrinsic strength of gel for the type of cement and of  specimen used. Numerical values differ little for the usual range of Portland cements  except that a higher C3A content leads to a lower strength at a given gel / space ratio.   If A cm3 of air are present in the set paste, the ratio wo / c in the above  expression would be 
replaced by    (formula) (See Fig. 5.9).    Fig. 5.8. Relation between the compressive strength of mortar and gel / space ratio    Fig. 5.9. Relation between the compressive strength of mortar and gel / space ratio  modified to include entrapped air voids     The resulting formula for strength would be similar to that of Feret but the ratio used h
ere  involves a quantity proportional to the volume of hydrated cement instead of the total  volume of cement and is thus applicable at any age.   The expression relating strength to the gel / space ratio can be written in a number of ways.  It may be convenient to utilize the fact that the volume of non-evaporable water, wn, is  proportional to t
he volume of the gel; and also that the volume of mixing water, wo, is related  to the space available for the gel. The strength, fc, in pounds per square inch, for f, greater than about  2000 psi, when the relation is approximately linear, can then be written in the form -    (formula).    Alternatively, the surface area of gel, Vm, can be used. 
Then    (formula).    Fig. 5.10 shows Powers' actual data for cements with low C3A contents.    Fig. 5.10. Relation between the strength of cement paste and the ratio of surface area of  gel Vm, to the volume of mixing water wo     The above expressions have been found to be valid for many cements but the numerical  coefficients may depend on the 
intrinsic strength of the gel produced by a given cement. In  other words, the strength of the paste depends primarily on the physical structure of the gel but  the effects of the chemical composition of cement cannot be neglected; however, at later ages  these effects become minor only. Another way of recognizing the properties of the gel is to s
ay  that strength depends primarily on porosity but also on the ability of the material to resist crack  propagation, which is a function of bonding. Poor bond between two crystals can be  considered to be a crack.      Effective Water in the Mix    The relations discussed so far in this chapter involve the quantity of water in the mix. This  need
s a more careful definition. We consider as effective that water which occupies space outside  the aggregate particles when the gross volume of concrete becomes stabilized, i.e., approximately at  the time of setting. Hence the terms effective or net water / cement ratio.   Generally, water in concrete consists of that added to the mix and that he
ld by the aggregate at the  time when it enters the mixer. A part of the latter water is absorbed within the pore structure of the  aggregate while some exists as free water on the surface of the aggregate and is therefore no different  from the water added direct into the mixer. Conversely, when the aggregate is not saturated and  some of its por
es are therefore air-filled, a part of the water added to the mix will be absorbed by  the aggregate during the first half hour or so after mixing. Under such circumstances the  demarcation between absorbed and free water is a little difficult.   On a site, the aggregate is as a rule wet, and only the water in excess of that required to  bring it 
to a saturated and surface-dry condition is considered to be the effective water of the mix.  For this reason, the strength curves of Road Note No. 4 (see Fig. 10.2) are based on the water  in excess of that absorbed by the aggregate. On the other hand, McIntosh and Erntroy's  data refer to the total water added to a dry aggregate. This condition 
of aggregate is frequently met in  the laboratory. Care is therefore necessary in translating laboratory results into mix proportions to be  used on a site and all reference to water / cement ratio must make it clear if total rather than effective  water is considered.      Nature of Strength of Concrete    The paramount influence of voids in conc
rete on its strength has been repeatedly mentioned, and  it should be possible to relate this factor to the actual mechanism of failure. For this  purpose, concrete is considered as a brittle material, even though it exhibits a small amount of  plastic action, as fracture under static loading takes place at a moderately low total strain; a strain 
 of 0 001 to 0 005 at failure has been suggested as the limit of brittle behaviour.    Strength in Tension    The actual (technical) strength of cement paste or of similar brittle materials such as stone is very  much lower than the theoretical strength estimated on the basis of molecular cohesion, and  calculated from the surface energy of a soli
d assumed to be perfectly homogeneous and flawless.  The theoretical strength has been estimated to be as high as 10.5 GPa (1.5 x 10'6 psi).   This discrepancy can be explained by the presence of flaws postulated by Griffith. These  flaws lead to high stress concentrations in the material under load so that a very high stress  is reached in very s
mall volumes of the specimen with a consequent microscopic fracture,  while the average (nominal) stress in the whole specimen is comparatively low. The flaws vary in  size and it is only the few largest ones that cause failure: the strength of a specimen is thus a problem  of statistical probability, and the size of the specimen affects the proba
ble nominal stress at  which failure is observed.   Cement paste is known to contain numerous discontinuities - pores, fissures and voids -  but the exact mechanism through which they affect the strength is not known. The voids  themselves need not act as flaws, but the flaws may be cracks in individual crystals associated with the  voids - or cau
sed by shrinkage or bad bond. In unsegregated concrete, voids are distributed in a  random manner, a condition necessary for the application of Griffith's hypothesis. While we do  not know the exact mechanism of rupture of concrete this is probably related to the bond  within the cement paste and between the paste and the aggregate.   Griffith's h
ypothesis postulates microscopic failure at the location of a flaw, and it is usually  assumed that the "volume unit" containing the weakest flaw determines the strength of the  entire specimen. This statement implies that any crack will spread throughout the section of the  specimen subjected to the given stress, or, in other words, an event taki
ng place in an  element is identified with the same event taking place in the body as a whole.   This behaviour can be met with only under a uniform stress distribution, with the additional  proviso that the "second weakest" flaw is not strong enough to resist a stress of n / n-1 times the  stress at which the weakest flaw failed, where n is the n
umber of elements in the section under load,  each element containing one flaw.   While local fracture starts at a point and is governed by the conditions at that point, the  knowledge of stress at the most highly stressed point in the body is not sufficient to  predict failure. It is necessary to know also the stress distribution in a volume suff
iciently  extended round this point as the deformational response within the material, particularly near failure,  depends on the behaviour and state of the material surrounding the critical point, and the possibility of  spreading of failure is strongly affected by this state. This would explain, for instance, why  the maximum fibre stresses in f
lexure specimens at the moment of incipient failure are higher  than the strength determined in uniform direct tension: in the latter case the propagation of fracture  is not blocked by the surrounding material. Some actual data on the relation between the  strength in flexure and in direct tension are given in Fig. 8.10.   We can see then that in
 a given specimen different stresses will produce fracture at different  points, but it is not possible physically to test the strength of an individual element without  altering its condition in relation to the rest of the body. If the strength of a specimen is governed  by the weakest element in it, the problem becomes that of the proverbial wea
kest link in a  chain. In statistical terms we have to determine the least value (i.e. the strength of the most  effective flaw) in a sample of size n, where n is the number of flaws in the specimen. The chain  analogy may not be quite correct because in concrete the links may be arranged in parallel as well  as in series but computations on the b
asis of the weakest link assumption yield results of the  correct order. It follows that the strength of a brittle material such as concrete cannot be described  by an average value only: an indication of the variability of strength must be given, as well as  information about the size and shape of the specimens. These factors are discussed in  Ch
apter 8.    Cracking and Failure in Compression    Griffith's hypothesis applies to failure under the action of a tensile force but it can be extended to  fracture under bi- and triaxial stress and also under uniaxial compression. Even when two  principal stresses are compressive, the stress along the edge of the flaw is tensile at some points,  s
o	that fracture can take place. Orowan calculated the maximum tensile stress at the tip of the  flaw of the most dangerous orientation relative to the principal stress axes as a function of  the two principal stresses P and Q. The fracture criteria are represented graphically in Fig. 5.11,  where K is the tensile strength in direct tension. Fractu
re occurs under a combination of P and Q such  that the point representing the state of stress crosses the curve outwards on to the shaded side.    Fig. 5.11. Orowan's criteria of fracture under biaxial stress     From Fig. 5.11 it can be seen that fracture may occur when uniaxial compression is applied; this  has in fact been observed in tests on
 concrete compression test specimens. The nominal  compressive strength in this case is 8K, i.e. eight times the tensile strength determined in a direct  tension test. This figure is in good agreement with the observed values of the ratio of the  compressive to tensile strengths of concrete. There are, however, difficulties in reconciling  certain
 aspects of Griffith's hypothesis with the observed direction of cracks in compression  specimens. It is possible, though, that failure in such a specimen is governed by the lateral  strain induced by Poisson's ratio. The order of values of Poisson's ratio for concrete is such that,  for elements sufficiently removed from the platens of the testin
g	machine, the resulting lateral  strain can exceed the ultimate tensile strain of concrete. Failure occurs then by splitting  at right angles to the direction of the load and this has been frequently observed, especially in  specimens whose height is greater than breadth.   There are strong indications that it is not a limiting stress but a limit
ing tensile strain that  determines the strength of concrete under static loading:  this is usually assumed to be between  100 x 10'-6 and 200 x 10'-6 The failure criterion of limiting tensile strain is supported by an  analysis recently advanced by Lowe. It has been found that at the point of  initial cracking the  strain on the tension face of a
 beam in flexure and the  lateral tensile strain in a cylinder in  uniaxial compression are of the same magnitude.     The tensile strain in a beam at cracking is    tensile stress at cracking / E,    where E is the modulus of elasticity of concrete over the linear range of  deformation.  Now, the lateral strain in a compression specimen when crac
king is first observed is    u x compressive stress at cracking / E,    where u is the static Poisson's ratio, and E is the same as above. From the  observed equality of the two strains it would appear that    u = tensile stress at cracking in flexure / compressive stress at cracking in a compression specimen.     Poisson's ratio varies generally 
between about 0 11 for high strength concrete and 0 21 for weak  mixes (see p. 370), and it is significant that the ratio of the nominal tensile and compressive  strengths for different concretes varies in a similar manner and between approximately the  same limits. There is thus a possibility of a connection between the ratio of nominal strengths
	and Poisson's ratio, and there are good grounds for suggesting that the mechanism producing  the initial cracks in uniaxial compression and in flexure tension is the same. The nature of  this mechanism has not been established, but cracking is probably due to local breakdowns in  bond between the cement and the aggregate.  The definition of fail
ure of concrete is not obvious. One proposal is to associate failure with the  so-called discontinuity point, defined as the point at which the volumetric strain stops decreasing  and Poisson's ratio starts to increase sharply. At this stage, extensive mortar cracking  starts to develop (see p. 285). This is the beginning of instability, and susta
ined loading above this  point will lead to failure. The lateral tensile strain at discontinuity depends on the level of axial  compression and is greater for stronger concrete; Carino and Slate observed an average value  of about 300 x 10'-6 at a stress of 7 5 MPa (1100 psi). It should be stressed, however, that other  workers consider that concr
ete is damaged progressively and do not recognize the discontinuity  point as a significant feature .   The ultimate failure under the action of a uniaxial compression is either a tensile failure of cement  crystals or of bond in a direction perpendicular to the applied load, or is a collapse caused by the  development of inclined shear planes. It
 is probable that ultimate strain is the criterion of failure,  but the level of strain varies with the strength of concrete: the higher the strength the lower the  ultimate strain. Some typical, but by no means general, values are given below:   Under triaxial compression, failure must take place by crushing: the mechanism is, therefore,  quite d
ifferent from that described above. An increase in lateral compression increases the  axial load that can be sustained, as shown for instance in Fig. 5.12. With very high lateral  stresses extremely high strengths have been recorded (Fig. 5.13). It may be noted that if the  development of pore water pressure in concrete is limited by allowing the 
displaced pore water to  escape through the loading platens, then the apparent strength is higher.    Fig. 5.12. Influence of lateral stress on the axial stress at failure of neat cement paste and  of mortar    Fig. 5.13. Influence of high lateral stress on the axial stress at failure of concrete     A lateral tensile stress has a similar influenc
e but, of course, in the opposite direction.  This behaviour agrees well with the theoretical considerations of page 539.   In practice it is likely that failure of concrete takes place over a range of stress rather than  as an instantaneous phenomenon, so that ultimate failure is a function of the type of loading. This is of special interest when
 repeated loading is applied - a condition frequently met with in practice.  Fatigue strength of concrete is considered on page 338.     Microcracking    Investigations have shown that very fine cracks at the interface between  coarse aggregate and cement paste exist in fact even prior to application of the load on  concrete. These cracks remain s
table up to about 30 per cent  or more of the ultimate load and then begin to increase in length, width,  and number. The overall stress under which they develop is sensitive to the  water / cement ratio of the paste. This is the stage of slow crack propagation.  At 70 to 90 per cent of the ultimate strength, cracks open through the mortar  (cemen
t paste and fine aggregate); they bridge the bond cracks so that a continuous  crack pattern is formed. This is the fast crack stage and, if  the load is sustained, failure may take place with time.   Interesting results of measurement of crack length are shown in Fig. 5.14. It can be seen that  there was very little increase in the total length b
etween the beginning of loading and a stress  equal to about 0.85 of the prism strength (100 mm by 100 mm by 300 mm (4 by 4 by 12 in.)  prisms). A further increase in stress resulted in a large increase in the total length of cracks.  At a stress / strength ratio of about 0.95 (determined on prisms), not only interface cracks but also  mortar crac
ks were present, and many cracks tended to become oriented roughly parallel to the  direction of the applied load. Once the specimen reached the descending part of the stress / strain  curve, the rate of increase in the crack length and width became large.    Fig. 5.14. Relation between the observed length of cracks in an area of 100 mm2 and the  
stress / strength ratio in compression (based on prisms)    Fig. 5.14 also shows the crack development under a cyclic stress alternating between zero and  0.85 of the prism strength. Immediately prior to failure, the cracks became longer and wider.  Likewise, sustained loading at a stress / strength ratio of 0.85 led to an increase in cracking pri
or to  failure.   It may be proper at this stage to consider the development of failure of concrete, although general  agreement upon it is still remote. One aspect that is not controversial is  the influence of the heterogeneity of concrete on its failure behaviour. While in a homogeneous  body under a simple state of stress, the stress trajector
ies are straight or simple curves,  this cannot be the case in concrete owing to the presence of aggregate: the properties of  the material vary from point to point, and the aggregate - paste bond  surfaces may form all the possible angles with the direction of the external force. As a  result, the local stress varies substantially above and below
 the  nominal applied stress.    Influence of Coarse Aggregate on Strength    Vertical cracking in a specimen subjected to uniaxial compression starts under a load equal to 50  to 75 per cent of the ultimate load. This has been determined from measurements of the velocity  of sound transmitted through the concrete, and also using ultrasonic pulse 
velocity techniques.  The stress at which the cracks form depends largely on the properties of the coarse aggregate:  smooth gravel leads to cracking at lower stresses than rough and angular crushed rock, probably  because mechanical bond is influenced by the surface properties and, to a certain degree, by the  shape of the coarse aggregate.   The
 properties of aggregate affect thus the cracking, as distinct from the ultimate, load in  compression and the flexural strength in the same manner, so that the relation between the  two quantities is independent of the type of aggregate used. Fig. 5.15 shows Jones and  Kaplan's results, each symbol representing a different type of coarse aggregat
e. On the other hand,  the relation between the flexural and compressive strengths depends on the type of coarse  aggregate used (see Fig. 5.16) since (except in high strength concrete) the properties of aggregate,  especially its shape and surface texture, affect the ultimate strength in compression very much less  than the strength in tension or
 the cracking load in compression. In experimental concrete, entirely  smooth coarse aggregate led to a lower compressive strength, typically by 10 per cent, than when  roughened.    Fig. 5.15. Relation between flexural strength and compressive stress at cracking for  concretes made with different coarse aggregates    Fig. 5.16. Relation between c
ompressive strength and indirect - tensile strength for concretes  of constant workability made with various aggregates (water / cement ratio between 0.33 and  0.68, aggregate / cement ratio between 2.8 and 10.1)     The influence of the type of coarse aggregate on the strength of concrete varies in magnitude and  depends on the water / cement rat
io of the mix. For water / cement ratios below 0.4, the use of  crushed aggregate has resulted in strengths up to 38 per cent higher than when gravel is used. The  behaviour at a water / cement ratio of 0.5 is shown in Fig. 5.17. With an increase in the  water / cement ratio, the influence of aggregate falls off, presumably because the strength of
 the  paste itself becomes paramount, and at a water / cement ratio of 0.65 no difference in the  strengths of concretes made with crushed rock and gravel has been observed.    Fig. 5.17. Relation between compressive strength and age for concretes made with  various aggregates (water / cement ratio = 0.5)     The influence of aggregate on flexural
 strength seems to depend also on the moisture  condition of the concrete at the time of test.  The shape and surface texture of coarse aggregate affect also the impact strength of  concrete, the influence being qualitatively the same as on the flexural strength (see p. 695).   Kaplan observed that the flexural strength of concrete is generally lo
wer than the  flexural strength of corresponding mortar. Mortar would thus seem to set the upper limit  to the flexural strength of concrete and the presence of the coarse aggregate generally  reduces this strength. On the other hand, the compressive strength of concrete is higher  than that of mortar, which, according to Kaplan, indicates that th
e mechanical  interlocking of the coarse aggregate contributes to the strength of concrete in  compression. This behaviour has not, however, been confirmed to apply generally, and the  question of the influence of aggregate on strength is considered further in the next  section.    Influence of Richness of the Mix on Strength    The anomalous beha
viour of extremely rich mixes was mentioned on p. 272, but the  aggregate / cement ratio affects the strength of all medium and high-strength concretes,  i.e. those with a strength of about 35 MPa (5000 psi) or more. There is no doubt that the  aggregate / cement ratio is only a secondary factor in the strength of concrete but it has been  found t
hat for a constant water / cement ratio a leaner mix leads to a higher strength (see Fig. 5.18).	Fig. 5.18. Influence of the aggregate / cement ratio on strength of concrete     This behaviour is probably associated with the absorption of water by the aggregate: a  larger amount of aggregate absorbs a greater quantity of water, the effective  w
ater / cement ratio being thus reduced. It is probable, however, that other factors also  play a part: for instance, the total water content per cubic metre of concrete is lower  in a leaner mix than in a rich one. As a result, in a leaner mix the voids form a  smaller fraction of the total volume of concrete, and it is these voids that have an  a
dverse effect on strength.   Recent studies on the influence of aggregate content on the strength of concrete with a  given quality of cement paste indicate that when the volume of aggregate is increased  from zero to 20, there is a gradual decrease in compressive strength, but between 40 and  80 per cent there is an increase. The pattern of behav
iour is shown in Fig. 5.19. The  reasons for this effect are not clear, but it is the same at all water / cement ratios. The  influence of the volume of aggregate on tensile strength is broadly similar (Fig. 5.20).    Fig. 5.19. Relation between the compressive strength of cylinder (100 mm diameter,  300 mm in length) and volume of aggregate at a 
constant water / cement ratio of 0.50    Fig. 5.20. Relation between direct tensile strength and volume of aggregate at a  constant water / cement ratio of 0.50     The effects are smaller in cubes than in cylinders or prisms. In consequence, the ratio  of cylinder strength to cube strength (cf. p. 543) decreases as the volume of aggregate  increa
ses from zero to 40 per cent. The explanation lies probably in the greater  influence of the aggregate on the crack pattern when the end effect of platens is  absent (see p. 539).      Equation of Strength    Strength of concrete is an inherent property of the material but, as measured in  practice, is also a function of the stress system which is
 acting. Mather pointed out  that, ideally, it should be possible to express the failure criteria under all possible  stress combinations by a single stress parameter, such as strength in uniaxial tension.  However, such a solution has not yet been found.   To develop equations for concrete strength, Berg considered the stress, ocr, at the  initia
tion of crack propagation, the maximum nominal stress, opr, on a concrete prism,  and the cleavage strength, ocl, of concrete in the direction normal to the applied  compressive stress. The cleavage strength is approximately equal to the strength in  axial tension. He showed that for the general case of a principal stress system o1, o2, and o3  (w
here o1 > o2 > o3 and compression is positive)    (formula).     This equation can be used for an analytical evaluation of the failure of concrete under  combined states of stress with the parameters opr, ocr, and ocl determined from tests on  uniaxial compression and uniaxial tension. Some empirical data for specimens of  various types are shown 
in Fig.	5.21. The criterion ceases to apply, however, when o2  and o3 have values such that the transverse cleavage strength cannot be overcome. The  behaviour of concrete is then no longer brittle but plastic. Some data on the strength  of concrete when o2 = o3 and o1 > o2 are shown in Fig. 5.13.    Fig. 5.21. Berg's equation for strength of conc
rete	A general biaxial stress interaction curve is shown in Fig. 5.22. A large interaction  is observed when there is a considerable frictional restraint at the platens, but when  the end restraint of the specimen is effectively eliminated (e.g. by the use of steel  brush platens, see p. 537), the effect is much smaller. It can be seen from Fi
g. 5.22  that under a biaxal stress o1 = o3, strength is only 16 per cent higher than in uniaxial compression;  biaxial tensile strength is no different from uniaxial tensile strength. These findings were  confirmed by other workers. Experimental data on interaction are plotted in Fig. 5.23;  these were obtained with brush platen loading and by th
e use of fluid  membranes and solid platens. Some contradictory data of other investigators can be  explained by the use of uncertain end restraints.    Fig. 5.22. Interaction curve for biaxial stress when the end restraint is effectively eliminated  (o1 and o2 are the biaxial stresses applied)    Fig. 5.23. Strength of concrete under multiaxial s
tress as measured by various investigators.  Wet or air-dried concrete (fc = compressive strength )     The level of uniaxial compressive strength virtually does not affect the shape of the  curve or the magnitude of the values given by it; the prism strength range tested  was 19 to 58 MPa (2700 to 8350 psi) and both the water / cement ratio and c
ement content  varied widely. However, in compression - tension and in biaxial tension, the relative  strength at any particular biaxial stress combination decreases as the level of uniaxial  compressive strength increases. This accords with the general observation that the  ratio of uniaxial tensile strength to uniaxial compressive strength decre
ases as the  compressive strength level rises (see p. 301); in these tests, the ratio was 0.11, 0.09,  and 0.08 at a uniaxial compressive strength level of 19, 31, and 58 MPa (2700, 4450, and  8350 psi) respectively.   Generally, triaxial compression increases the strength of weaker or leaner concrete  relatively more than that of stronger or rich
er concrete. For the range of  conventional concretes, Hobbs found that under triaxial compression, the major principal  stress at failure, ol, can be expressed as    (formula),    where o3 = minor principal stress,  and fcyl = cylinder strength,  all values being in MPa.   The limited information on lightweight aggregate concrete suggests that th
e influence  of o3 is not as large as with normal aggregates; therefore the coefficient 4.8 in the  above equation can be reduced to about 3.2.  The combined strength results for concretes in triaxial compression and in biaxial  compression plus tension, may be represented by the equation    (formula),    all values again being in MPa and compress
ion being taken as positive.   The values given in (1) and (2) apply to conventional concretes only and  not to neat cement pastes or mortars.   Substituting equations (1) and (2) into equation (3) yields the failure  criterion for conventional concretes:    (formula).     This equation is plotted in Fig. 5.24 for various values of the cylinder st
rength, fcyl.  The generality of this equation should not be overestimated for, as Hobbs points out,  the tensile strength and compressive strength of concrete are not equally affected by  the aggregate type and grading and by the direction of the applied stress relative to  the direction of casting. In each case, the tensile strength is more sens
itive.    Fig.	5.24. Failure stresses in concrete under biaxial stress     Recognizing that the strength of concrete cannot be predicted by considering  limitations on the compressive, tensile, and shearing stresses independently of each  other, Bresler and Pisters proposed a general equation of strength of concrete. The  equation uses criteria in
dependent of chosen cartesian co-ordinates, i.e. so-called  invariant terms, specifically    (formula)    with the generalized failure criterion F (I1, I2, I3) = 0.   Obviously, if one of the principal stresses is zero, I3 = 0.   Bresler and Pisters considered, after Nadai, failure to be defined by     (formula),  where o0 and t0 are the normal an
d shearing octahedral stresses given by    (formula).	Octahedral stresses are so named because they occur on the sides of an octahedral  element formed by planes whose normals make equal angles with the principal stress axes.   To avoid the influence of the properties of the specimen on the compressive strength of  concrete, the octahedral str
esses are rendered dimensionless by dividing them by the  nominal uniaxial compressive strength of the specimen used, fc,. An equation of the type    (formula)    is obtained, and this gives the combination of stresses that will just cause failure. A  typical curve is shown in Fig. 5.25. The effect of I3 has been found to be significant  so that t
he intermediate stress (e.g. in biaxial compression - tension) cannot be  neglected. For this reason, some of the classical failure theories are not considered  properly applicable to concrete. While the octahedral stress failure theory has met with  considerable success it does not give the same equation for bi- and triaxial stress so  that furth
er development of an equation of strength is still needed. Such information  will be of value in the design of structures such as shells, plates, pressure vessels  and even in parts of flexural members.    Fig. 5.25. Relation between octahedral shear stress and octahedral normal stress at  failure for 20 MPa (3000 psi) concrete subjected to uniaxi
al compression and to  biaxial compression - tension	Effect of Age on Strength of Concrete    The relation between the water / cement ratio and the strength of concrete applies to  one type of cement and one age only. On the other hand, the  strength versus gel / space  ratio relationship has a more general application because the amount of g
el present in  the cement paste at any time is itself a function of age and type of cement. In other  words, different cements require a different length of time to produce the same  quantity of gel.   The rate of gain of strength of different cements was discussed in Chapter 2, and Figs.  2.1 and 2.2 show typical strength - time curves. The influ
ence of the curing conditions on  the development of strength is considered later in this chapter, but here we are  concerned with the practical problem of strength of concrete tested at different ages.  In the majority of cases, the tests are made at the age of 28 days when the strength of  concrete is considerably lower than its long-term streng
th. In the past, the gain in  strength beyond the age of 28 days was regarded merely as contributing to an increase in  the factor of safety of the structure, but since 1957 the codes of practice for  reinforced and prestressed concrete allow the gain in strength to be taken into account  in the design of structures that will not be subjected to l
oad until a later age except  when no-fines concrete is used; with some lightweight aggregates, verifying tests are  advisable. The values of strength given in the British Code of Practice CP 110 : 1972,  based on the 28-day compressive strength, are given in Table 5.1, but they do not, of  course, apply when accelerators are used.    Table 5.1: B
ritish Code of Practice CP 110 : 1972 Factors for Increase in Compressive  Strength of Concrete with Age (Average Values)     The rate of gain is the strength of concrete is of interest also in connection with  testing. It is often desirable to check the suitability of a mix long before the results  of the 28-day test are available. However, even 
if the curing conditions are carefully  controlled, the prediction of the 28-day strength from that measured at the age of 7  days is difficult, mainly because of the variation in the intrinsic rate of hardening of  commercial cements. Furthermore, mixes with a low water / cement ratio gain strength,  expressed as a percentage of long-term strengt
h, more rapidly than mixes with higher  water / cement ratios (Fig. 5.26). This is because in the former case the cement grains are closer to  one another and a continuous system of gel is established more rapidly. For this reason,  a general extrapolation of, say, the 7-day strength to the 28-day strength is not easy,  even when dealing with one 
cement only.	Fig. 5.26. Relative gain of strength with time of concretes with different water / cement  ratios, made with ordinary Portland cement     When no specific data on the materials used are available, the 28-day strength may be  assumed to be 1.5 times the 7-day strength and, as an alternative to the specified  28-day strength of test 
cubes, the Code of Practice CP 114 (1969) accepts a 7-day strength  equal to not less than 3 of the required 28-day strength. Tests have shown that for  concretes made with British ordinary Portland cement the ratio of the 28-day to 7-day  strengths lies generally between 1.3 and 1.7 but the majority of the results fall above  1.5. The extrapolati
on of 7-day strength according to the Code of Practice is therefore  quite reliable. According to ASTM Standard, American Type 1 cements have a lower ratio  of 28-day to 7-day strength: 57 per cent of them have a ratio of at least 1.32 but only 2  per cent higher than 1.50.   It should be noted that in a hot climate the early strength gain is high
 and the ratio  of the 28-day to 7-day strengths tends to be lower than in cooler weather. This is also  the case with some lightweight aggregate concretes.   In Germany, the relation between the 28-day strength, fc,28 and the 7-day strength,  fc,7 is often taken to lie between    (formula),    fc being expressed in MPa. For strengths expressed in
 psi the two constant terms become  respectively 150 and 850.   Hummel recommends the use of an approximately linear relation between the strength  and the logarithm of age within the range of 3 days to 2 months. Thus if the strength is  determined at 3 and 7 days, it is possible to estimate the 28-day strength by  extrapolation. The validity of t
his approach between about 28 and 250 days is shown in  Fig. 5.27.    Fig. 5.27. Gain of strength of air-stored concrete with time (Austrian Portland cement;  cement and aggregate from one source; water / cement ratio = 0.65; cement content = 260 kg / m3  (440 lb / yd3); curing 7 days in wet sand, then in air at room temperature)     Pioneiro sugg
ested an expression of the type    (formula),   where fc,7 and fc,28 are strengths at 7 and 28-days respectively, and k1 and k2 are coefficients,  different for each cement and curing condition used. The value of k1 ranges from about 0.3 to  0.8, and that of k2 from 3 to 6.  All the expressions mentioned here apply only to concrete made with ordin
ary Portland  cement. Many of the other cements gain strength at different rates, and when they are  used the prediction of strength should be based on experimental results.   As far as the really long-term strength is concerned, American Portland cements made at  the beginning of the century (which had a high C2S content and a low specific surfac
e)  led to an increase in the strength of concrete stored outdoors which was proportional to  the logarithm of age up to 50 years. The 50-year strength was typically 2.4 times the  28-day strength. However, cements made since the 1930s (with a lower C2S content and  a higher specific surface) reach their peak strength between 10 and 25 years and  
thereafter undergo some retrogression of strength. German Portland cements made  in 1941, when used in concrete stored outdoors, led after 30 years to a strength 2.3 times  the 28-day strength. The relative increase in strength was greater at higher  water / cement ratios. By comparison, Portland blast-furnace cement led to a 3.l-fold  increase.  
	Autogenous Healing    Fine cracks in fractured concrete, if allowed to close without tangential displacement,  will heal completely under moist conditions. This autogenous healing is due to the  hydration of the hitherto unhydrated cement, and may also be aided by carbonation. The  younger the concrete, i.e. the more unhydrated cement it conta
ins, the higher the  regain of strength, but healing without a loss of strength has been observed at ages up to  three years. The application of pressure across the crack assists in healing.      Relation between Compressive and Tensile Strengths    From the discussion on the strength of compression and tension (both direct and flexure)  test spec
imens it would be expected that the two types of strength are closely  related. This is indeed the case but there is no direct proportionality, the ratio of the  two strengths depending on the general level of strength of the concrete. In other  words, as the compressive strength, fc, increases, the tensile strength, ft, also increases  but at a d
ecreasing rate.  A number of factors affect the relation between the two strengths. The beneficial effect  of crushed coarse aggregate on flexural strength was discussed on page 287, but it seems  that the properties of fine aggregate also influence the ft / fc ratio. The ratio is  furthermore affected by the grading of the aggregate. This is prob
ably due to the different  magnitude of the wall effect in beams and in compression specimens: their surface / volume  ratios are dissimilar so that different quantities of mortar are required for full  compaction.   Age is also a factor in the relation between ft and fc: beyond about one month the  tensile strength increases more slowly than the 
compressive strength so that the ratio  ft / fc decreases with time. This is in agreement with the general tendency of the  ratio to decrease with an increase in fc.   The tensile strength of concrete is more sensitive to inadequate curing than the  compressive strength, possibly because the effects of non-uniform shrinkage of flexure  test beams 
are very serious. Thus air-cured concrete has a lower ft / fc ratio than  concrete cured in water and tested wet.   Air entrainment affects the ft / fc ratio because the presence of air lowers the  compressive strength of concrete more than the tensile strength, particularly in the  case of rich and strong mixes. The influence of incomplete compac
tion is similar to  that of entrained air (see Fig. 5.28).    Fig. 5.28. Relation between compressive and flexural strengths of incompletely  compacted concrete. Compressive strength determined on equivalent cubes;  modulus of rupture determined under third-point loading     Lightweight concrete conforms broadly to the pattern of the relation  bet
ween ft and fc for ordinary concrete. At very low strengths (say, 2 MPa  (300 psi)) the ratio ft / fc can be as high as 0.3, but at higher strengths it  decreases to values comparable with those for ordinary concrete. In the  latter case, the ratio varies generally between 0.16 and 0.07 when the cube  crushing strength is taken as fc, and the modu
lus of rupture under  third-point loading as ft. As shown on page 547, the tensile strength of  concrete depends on the type and method of test, so that the means of  determining ft must be clearly stated.   A number of empirical formulae connecting ft and fc have been suggested, many of them of  the type -    (formula),    where k and n are coeff
icients. Tests made at the Building Research Station on 100 mm (4  in.) cubes and 100 x 100 x 400 mm (4 x 4 x 16 in.) beams centrally loaded have yielded  values of n = 1/2 and average k = 8.3. The latter actually varies between 6.2 for some  gravels and 10.4 for crushed rock. Other experiments indicate that n may vary  between 1/2 and 3/4. The fo
rmer value is used by the American Concrete Institute, but  Gardner and Poon found a value nearer the latter, cylinders being used in both  cases.   Comite Europeen du Beton assumes that the mean direct tensile strength is related to  the characteristic compressive strength of cylinders (see p. 656) by the expression    (formula),    the strengths
 being expressed in megapascals. Such a relation may be of use in design  but does not properly describe the properties of the material. Another formula was  suggested at the University of Illinois -    (formula),    where ft is the modulus of rupture and fcyl is determined on standard test cylinders,  both expressed in pounds per square inch.   I
n view of the numerous factors influencing the ratio of the strengths it is not  surprising that no simple relation is generally applicable. Data obtained at the  laboratories of the Portland Cement Association are given in Table 5.2, and Fig. 5.29  shows the results of Walker and Bloem's tests.    Fig. 5.29. Relation between compressive and flexu
ral strengths of concrete (both air- and  non-air-entrained)    Table 5.2: Relation between Compressive and Tensile Strengths of Concrete      Bond between Concrete and Reinforcement    Since structural concrete is in the vast majority of cases used with steel  reinforcement, the strength of bond between the two materials is of considerable  inter
est. Bond arises primarily from friction and adhesion between concrete and steel,  and may also be affected by the shrinkage of concrete relative to the steel. Bond  involves, however, not only the properties of concrete but also the mechanical  properties of steel and its position in the concrete member. For this reason the subject  of bond is la
rgely outside the scope of this book.	In general terms, bond is related to the quality of the concrete, and bond strength is  approximately proportional to the compressive strength up to about 20 MPa (3000 psi).  For higher strengths of concrete, the increase in bond strength becomes progressively  smaller and eventually negligible (see Fig. 5.3
0). This is why most codes of practice  restrict the permissible value of bond in high-strength concrete. For instance, the  British Code of Practice for the Structural Use of Concrete CP110: 1972 gives the values  listed in Table 5.3.    Fig. 5.30. Influence of the strength of concrete on bond determined by pull-out tests    Table 5.3: Maximum Va
lues of Flexural Bond Stress in Concrete according to the  British Code of Practice for the Structural Use of Concrete CP110 : 1972     A rise in temperature reduces the bond strength of concrete: at 200 to 300 C (400 to  570 F) there may be a loss of one-half of the bond strength at room temperature.   Galvanizing and other protective treatment o
f reinforcement may reduce the bond  strength, probably because in treated steel the good bond of a rusty surface is absent.  However, data of the Building Research Establishment show that bond of galvanized  reinforcement is at least as good as that of ordinary steel bars and wires. The  explanation lies in the development, on initial attack of z
inc by alkalis, of a coating  of a complex calcium zincate compound, which is durable in an alkaline environment. It  should be noted, though, that, with white cements and some others, the absence of traces  of soluble chromium can lead to the evolution of hydrogen gas by the reaction of zinc  with alkalis, and this can reduce bond.   The thicknes
s of the galvanized coating is usually between 0.03 and 0.10 mm (0.001 and  0.004 in.). It is advantageous in that it gives protection to steel even when cover to  it is reduced (perhaps by up to 25 per cent); it permits use of lightweight aggregate  concrete without increased cover; and it gives a better protection in a marine  environment or whe
n chlorides are present. Against this, the higher cost has to be  considered.   The real benefit of galvanized reinforcement is when carbonation can reach the  reinforcement in lightweight concrete and in concrete of poor quality. In good quality  concrete, there appears to be no benefit. Conversely, under extreme conditions of high  chloride cont
ent, galvanized steel as well as uncoated steel is attacked, but there  is also evidence of superiority of galvanized reinforcement.   Galvanizing can also be applied to prestressing wire when severe attack is expected.      Curing of Concrete    In order to obtain good concrete the placing of an appropriate mix must be followed by  curing in a su
itable environment during the early stages of hardening. Curing is the name  given to procedures used for promoting the hydration of cement and consists of a control  of temperature and of the moisture movement from and into the concrete. The  temperature factor is dealt with on page 318.   More specifically, the object of curing is to keep concre
te saturated as possible,  until the originally water-filled space in the fresh cement paste has been filled to the  desired extent by the products of hydration of cement. In the case of site concrete,  active curing stops nearly always long before the maximum possible hydration has taken  place. The order of influence of moist curing on strength 
can be gauged from Fig. 5.31,  obtained for concrete with a water / cement ratio of 0.50. The loss of strength due to  inadequate curing is more pronounced in thinner elements but is smaller in lightweight  aggregate concrete. Tensile and compressive strengths are affected in a similar  manner; in both cases, richer mixes are slightly more suscept
ible.    Fig. 5.31. Influence of moist curing on the strength of concrete with a water / cement  ratio of 0.50	The necessity for curing arises from the fact that hydration of cement can take place  only in water-filled capillaries. This is why a loss of water by evaporation from the  capillaries must be prevented. Furthermore, water lost inter
nally by self-desiccation  has to be replaced by water from outside, i.e. ingress of water into the concrete must  be made possible.   It may be recalled that hydration of a sealed specimen can proceed only if the amount  of water present in the paste is at least twice that of the water already combined.  Self-desiccation is thus of importance in 
mixes with water / cement ratios below about  0.5; for higher water / cement ratios the rate of hydration of a sealed specimen equals  that of a saturated specimen. It should not be forgotten, however, that only half the  water present in the paste can be used for chemical combination;  this is so even if the total amount of water present is less 
than the water required for  combination. This statement is of considerable importance as it was formerly thought  that, provided a concrete mix contained water in excess of that required for the  chemical reactions with cement, a small loss of water during hardening would not  adversely affect the process of hardening and the gain of strength. It
 is now known that  hydration can take place only when the vapour pressure in the capillaries is  sufficiently high, about 0.8 of the saturation pressure. Hydration at a maximum rate can  proceed only under conditions of saturation. Fig. 5.32 shows the degree of hydration  after six months' storage at different relative humidities, and it is clear
 that below a  vapour pressure of 0.8 of the saturation pressure the degree of hydration is low, and  negligible below 0.3 of the saturation pressure.    Fig. 5.32. Water taken up by dry cement exposed for six months to different vapour pressures     It must be stressed that for a satisfactory development of strength it is not necessary  for all t
he cement to hydrate, and indeed this is only rarely achieved in practice: as  shown earlier, the quality of concrete depends primarily on the gel / space ratio of the  paste. If, however, the water-filled space in fresh concrete is greater than the volume  that can be filled by the products of hydration, greater hydration will lead to a higher  s
trength and a lower permeability.   Evaporation of water from concrete soon after placing depends on the temperature and  relative humidity of the surrounding air and on the velocity of wind which effects a  change of air over the surface of the concrete. An indication of the influence of these  three factors can be obtained from Figs. 5.33, 5.34,
 and 5.35, based on Lerch's  results. The difference between the temperatures of concrete and of air also affects the  loss of water, as shown in Fig. 5.36. Thus concrete saturated in day-time would lose  water during a cold night, and this would also be the case with  concrete cast in cold weather, even in saturated air. The examples quoted are m
erely  typical as the actual loss of water depends on the surface /  volume ratio of the  specimen.    Fig. 5.33. Influence of relative humidity of air on the loss of water from  concrete in  the early stages after placing (air temperature 21 C (70 F); wind velocity 4.5 m / s (10 mph))    Fig. 5.34. Influence of temperature of air and concrete on 
the loss of water from  concrete in the early stages after placing (relative humidity of air 70 per cent; wind  velocity 4.5 m / s (10 mph))     Fig. 5.35. Influence of wind velocity on the loss of water from concrete in the early  stages after placing (relative humidity of air 70 per cent, temperature 21 C (70 F))    Fig. 5.36. Influence of tempe
rature of concrete (at an air temperature of 4.5 C  (40 F)) on the loss of water from concrete in the early stages after placing (relative  humidity of air 100 per cent, wind velocity 4.5 m / s (10 mph))      Methods of Curing    No more than an outline of the different means of curing will be given here  as the actual procedures used vary widely,
 depending on the conditions on the site and  on the size, shape, and position of the member.   In the case of members with a small surface / volume ratio, curing may be aided by oiling  and wetting the forms before casting; the forms may also be wetted during hardening, and  after stripping the concrete should be sprayed and wrapped with polythen
e sheets or  other suitable covering.   Large surfaces of concrete, such as road slabs, present a more serious problem. In  order to prevent crazing of the surface on drying out, loss of water must be prevented  even prior to setting. As the concrete is at that time mechanically weak it is necessary  to suspend a covering above the concrete surfac
e. This protection is required only in  dry weather but may also be useful to prevent rain marring the surface of fresh  concrete.   Once the concrete has set, wet curing can be provided by keeping the concrete in  contact with a source of water. This may be achieved by spraying or flooding (ponding),  or by covering the concrete with wet sand or 
earth, sawdust or straw. Periodically-wetted  hessian or cotton mats may be used, or alternatively an absorbent covering with  access to water may be placed over the concrete. A continuous supply of water is  naturally more efficient than an intermittent one, and Fig. 5.37 compares the strength  development of concrete cylinders whose top surface 
was flooded during the first 24  hours with that of cylinders covered with wet hessian. The difference is greatest at low  water / cement ratios where self-desiccation operates rapidly. The influence of curing  conditions on strength is lower in  the case of air-entrained than non-air-entrained  concrete.    Fig. 5.37. Influence of curing conditio
ns on strength of test cylinders     Another means of curing is to use an impermeable membrane or  waterproof paper. A  membrane, provided it is not punctured or damaged, will effectively prevent evaporation  of water from the concrete but will not allow ingress of water to replenish that lost  by self-desiccation. The  membrane is formed by seali
ng compounds applied after free  water has disappeared from the surface of the concrete but before the pores in  concrete can absorb the compound. The membrane may be clear, white or black. The  opaque compounds have the effect of shading the concrete, and  a light colour leads to a lower  absorption of the heat from the sun, and  hence to a small
er rise in the temperature of  the concrete. The effectiveness (as measured by strength) of a white membrane and of  white translucent sheets of polyethylene is the same. Details of the curing  compounds are outside the scope of this book. ASTM Method C 156-74  prescribes tests  for the efficiency of curing compounds and ASTM Standard C 171-69 (re
approved 1975)  prescribes sheet materials for curing.   Except when used on concrete with a high water / cement ratio, sealing compounds  reduce the degree and rate of hydration compared with efficient wet curing. However, wet  curing is often applied only intermittently so that in practice sealing may lead to  better results. Waterproof paper, o
nce removed, does not interfere with the adhesion of  the next lift of concrete, but the effects of membranes have to be ascertained in each  case. However, sheeting can cause discoloration or mottling because of non-uniform  condensation of water on the underside of the paper.   The period of curing cannot be prescribed simply but it is usual to 
specify a minimum  of seven days for ordinary Portland cement concrete. With slower-hardening cements a  longer curing period is desirable. The temperature also affects the length of the  required period of curing and the British Code of Practice for the Structural Use of  Concrete CP 110: 1972 lays down the normal curing periods for different cem
ents and  exposure conditions in terms of maturity of concrete (see p. 314); however, the  quality of concrete is ignored. The American Concrete Institute Standard 308-71  (reaffirmed in 1978) gives extensive information on curing. Striking times for formwork  are given in a British publication.   High-strength concrete should be cured at an early
 age as partial hydration may make  the capillaries discontinuous: on renewal of curing, water would not be able to enter  the interior of the concrete and no further hydration would result. However, mixes with  a high water / cement ratio always retain a large volume of capillaries so that curing can  be effectively resumed at any time. No loss o
f strength is caused by delaying the  curing, as shown for instance by Fig. 5.38 for a 1:3.4 mortar with a water / cement ratio  of 0.70, but in practice early drying may lead to shrinkage and cracking, and delaying  curing is inadvisable.    Fig. 5.38. Relation between total curing time in water and tensile strength of mortar  briquettes      Mat
urity of Concrete    So far we have considered only the time aspect of curing but, as mentioned earlier, the  temperature during curing also controls the rate of progress of the reactions of  hydration and consequently affects the development of strength of concrete. This  influence is shown, for instance, in Fig. 5.39 obtained from tests on speci
mens cast,  sealed and cured at the indicated temperatures. The effect of the temperature at the  time of setting is considered on page 318.    Fig. 5.39. Ratio of strength of concrete cured at different temperatures to the 28-day  strength of concrete cured at 21 C (70 F) (water / cement ratio = 0.50; the specimens  were cast, sealed, and cured a
t the indicated temperature)	Since strength of concrete depends on both age and temperature we can say that strength  is a function of C(time interval x temperature), and this summation is called maturity.  The temperature is reckoned from an origin found experimentally to be between -12 and  -10 C (11 and 14 F). This is because at temperature
s below the freezing point of water  and down to about - 12 C (11 F) concrete shows a small increase in strength with time  but the low temperature must not be applied, of course, until after the concrete has set  and gained sufficient strength to resist damage due to the action of frost; a "waiting  period" of 24 hours is usually required. Below 
-12 C (11 F) concrete does not appear  to gain strength with time.   Maturity is measured in C-hours ( F-hours) or C-days (F-days). Fig. 5.40 and Fig.  5.41 show that compressive and tensile strengths plotted against the logarithm of  maturity give a straight line. It is, therefore, possible to express strength at any  maturity as a percentage of 
strength of concrete at any other maturity; the latter is  often taken as 19 800 Ch (35 600 Fh), being the maturity of concrete cured at 18 C (64 F)  for 28 days. The ratio of strengths, expressed as a percentage, can then be written  as -    (formula).    Fig. 5.40. Relation between logarithm of maturity and compressive strength of cubes    Fig. 
5.41. Relation between logarithm of maturity and splitting strength (tests carried  out at 2, 13, and 23 C (35, 55, and 73 F) up to 42 days)     Strictly speaking, the linearity applies only above a certain minimum maturity, as  shown in Fig. 5.42.    Fig. 5.42. Relation between compressive strength of ordinary Portland (Type I)  cement concrete a
nd maturity for the data of Gruenwald as treated by Lew and Reichard     The values of the coefficients A and B depend on the strength level of concrete; those  suggested by Plowman are given in Table 5.4. It can be seen that the strength - maturity  relation depends on the properties of the cement and on the general quality of the  concrete, and 
is valid only within a range of temperatures. This is apparent, for  instance, from Fig. 5.43 obtained by Klieger who tested ordinary Portland cement  concrete with a water / cement ratio of 0.43 and an air content of 4.5 per cent, cured at  23 C (73 F) from the age of 28 days onwards. A further complication arises from the  fact that the effects 
of a period of exposure to a higher temperature are not the same  when this occurs immediately after casting or later in the life of the concrete.  Specifically, early high temperature leads to a lower strength for a given total  maturity than when heating is delayed for at least a week or is absent. This is of  interest in connection with steam c
uring.    Table 5.4: Plowman's Coefficients for the Maturity Equation    Fig. 5.43. Influence of the temperature during the first 28 days after casting on the  strength - maturity relation     Nevertheless, the maturity rule applies fairly well when the initial temperature of  concrete is between 16 and 27 C (60 and 80 F) (although probably a much
 lower  minimum is acceptable) and no loss of moisture by drying takes place during the period  considered. It is these limitations that have prevented a more widespread acceptance  of the maturity rule in concrete practice, especially in the United States, although  recently the Federal Highway Administration has investigated the use of the matur
ity  rule in the determination of the potential strength of concrete from early  tests.   An excellent review of the maturity concept has been prepared by Malhotra.      Influence of Temperature on Strength of Concrete    We have seen that a rise in the curing temperature speeds up the chemical reactions of  hydration and thus affects beneficially
 the early strength of concrete without any ill-effects  on the later strength. However, a higher temperature during placing and setting,  although it increases the very early strength, may adversely affect the strength from  about 7 days onwards. The explanation is that a rapid initial hydration appears to form  products of a poorer physical stru
cture, probably more porous so that a large proportion  of the pores will always remain unfilled. It follows from the gel / space ratio rule that  this will lead to a lower strength compared with a less porous, though slowly hydrating,  paste in which a high gel / space ratio will eventually be reached. This explanation of  the adverse effects of 
a high early temperature on later strength has been extended by  Verbeck and Helmuth who suggest that the rapid initial rate of hydration at higher  temperatures retards the subsequent hydration and produces a non-uniform distribution of  the  products of hydration within the paste. The reason for this is that at the high  initial rate of hydratio
n there is insufficient time available for the diffusion of the  products of hydration away from the cement grain and for a uniform precipitation in the  interstitial space (as is the case at lower temperatures). As a result, a high  concentration of the products of hydration is built up in the vicinity of the hydrating  grains, and this retards t
he subsequent hydration and adversely affects the long-term  strength.   In addition, the non-uniform distribution of the products of hydration per se adversely  affects the strength because the gel / space ratio in the interstices is lower than would  be otherwise the case for an equal degree of hydration: the local weaker areas lower the  streng
th of the paste as a whole .   Fig. 5.44 shows Price's data on the effect of the temperature during the first two  hours after mixing on the developing of strength of concrete with a water / cement ratio  of 0.53. The range of temperatures investigated was 4 to 46 C (40 to 115 F), and  beyond the age of two hours all specimens were cured at 21 C (
70 F). The specimens  were sealed so as to prevent movement of moisture. Tests on cylinders moist-cured during  the first 24 hours at 2 C (36 F) and at 18 C (64 F), and thereafter at 18 C (64 F)  have shown that, at 28 days, the former are 10 per cent stronger than the latter.     Fig. 5.44. Effect of temperature during the first two hours after c
asting on the  development of strength (all specimens sealed and after the age of 2 hours cured at  21 C (70 F))     Some field tests have confirmed the influence of temperature at the time of placing on  strength: typically, for an increase of 5 C (9 F) there is a decrease in strength of  1 9 MPa (270 psi).   The influence of curing temperature o
n strength of concrete tested after cooling at 1  and 28 days is shown in Fig. 5.45. However, the temperature at the time of testing  also appears to be a factor, at least in the case of neat (ordinary Portland) cement  paste compacts with a water / cement ratio of 0.14. The temperature was kept constant  from the initiation of hydration. When tes
ted (at 64 and 128 days) at the curing  temperature, the specimens had a lower strength at higher temperatures (Fig. 5.46); but,  if cooled to 20 C (68 F) over a period of 2 hours prior to testing, only temperatures  above 65 C (150 F) had a deleterious effect (Fig. 5.47).    Fig. 5.45. Influence of curing temperature on compressive strength at 1 
and 28 days  (specimens tested after cooling to 23 C (73 F) over a period of two hours)    Fig. 5.46. Relation between compressive strength and curing time of neat cement  paste compacts at different curing temperatures. The temperature of the specimens  was kept constant up to and including the period of testing    Fig. 5.47. Relation between com
pressive strength and curing time of neat cement  paste compacts at different curing temperatures. The temperature of the specimens  was moderated to 20 C at a constant rate over a two-hour period prior to testing  (water / cement ratio = 0.14; Type I cement)     Tests have also been made on concretes stored in water at different temperatures for 
a  period of 28 days, and thereafter at 23 C (73 F). As in Price's tests, a higher  temperature was found to result in a higher strength during the first few days after  casting, but beyond the age of one to four weeks the situation changed radically (see  Fig. 5.43). The specimens cured at temperatures between 4 and 23 C (40 and 73 F) up to  the 
age of 28 days all showed a higher strength than those cured at 32 to 49 C (90 to  120 F). Among the latter, retrogression was greater the higher the temperature, but in  the lower range of temperatures there appeared to be an optimum temperature that yielded  the highest strength. It is interesting to note that even concrete cast at 4 C (40 F)  a
nd stored at as low a temperature as -4 C (25 F) for four weeks and then at 23 C (73 F),  is from the age of 3 months onwards stronger than similar concrete stored  continuously at 23 C (73 F). Fig. 5.48 shows typical curves for concrete containing  307 kg of ordinary Portland cement per cubic metre (517 lb / yd3) of concrete with 4.5 per  cent of
 entrained air. Similar behaviour has been observed when rapid hardening Portland  and modified cement are used. When calcium chloride is added to the mix the adverse  effects of a high temperature during setting are attenuated.    Fig. 5.48. Effect of temperature during the first 28 days on the strength of concrete  (water / cement ratio = 0.41; 
air content = 4.5 per cent; ordinary Portland cement)     The increase in strength caused by the addition of calcium chloride depends on the  temperature of the concrete and is proportionately greater at lower temperatures. For  instance, at 13 C (55 F) the addition of 2 per cent of CaCl2 increases the one-day  strength by about 140 per cent, but 
the relative increase in the same mix at 49 C (120 F)  is only 50 per cent. This behaviour is, indeed, to be expected, as the rate of  hydration at higher temperatures is high even without an accelerator so that little  scope is left for the action of CaCl2. Calcium chloride should ordinarily be used only  at normal or low temperatures but, in any
 case, its influence on the corrosion of  reinforcing steel should not be forgotten (see p. 105).   Klieger's tests indicate that there is an optimum temperature during the early life  of concrete that will lead to the highest strength at a desired age. For laboratory-made  concrete, using ordinary or modified Portland cement, the optimum temperat
ure is  approximately 13 C (55 F); for rapid hardening Portland cement it is about 4 C (40 F).  It must not be forgotten, however, that beyond the initial period of setting and  hardening the influence of temperature (within limits) accords with the maturity rule: a  higher temperature accelerates the development of strength.   The tests described
 so far were all made in the laboratory, and it seems that the  behaviour on the site in a hot climate may not be the same. Here, there are some  additional factors acting: ambient humidity, direct radiation of the sun, wind velocity,  and method of curing. It should be remembered also that the quality of concrete depends  on its temperature and n
ot on that of the surrounding atmosphere, so that the size of  the member also enters the picture. Likewise, curing by flooding in windy weather  results in a loss of heat due to evaporation so that the temperature of concrete is  lower than when a sealing compound is used. Shalons found evaporation immediately  after casting to be beneficial, pos
sibly because water is drawn out of the concrete  while capillaries can still collapse, so that the effective water / cement ratio is  decreased. If, however, evaporation is allowed to lead to the drying of the surface,  plastic shrinkage and cracking may result.   Shalon's tests have shown that under hot and dry conditions, such as are  encounter
ed in the desert, strength decreases with an increase in temperature down to a  critical value at about 30 C (86 F), but between 30 C and 45 C (86 F and 113 F)  there may be a slight recovery or no further loss. This behaviour has been observed  using concrete without entrained air and stored at a relative humidity of between 20 and  70 per cent. 
It is possible that the presence or absence of entrained air is  responsible, at least in part, for the difference between Klieger's and Shalon's  results. It seems then that we are not yet aware of all the factors involved in the  problem, and careful site tests should be made before construction is begun in an  unknown climate.   In general term
s, however, concrete cast in summer can be expected to have a lower  strength than a similar mix cast in winter. Indeed, on many construction sites the  strength of test specimens has been found to be lower during hot weather even though  from the time of stripping the moulds at the age of 24 hours the specimens were cured in  water at 18 C (64 F)
. Similarly, in tropical countries an apparently lower strength of  concrete has been observed (Fig. 5.49).    Fig. 5.49. Comparison of strengths of concrete with a water / cement ratio of 0.40 cast in  Lagos, Nigeria and in England      Steam Curing at Atmospheric Pressure    Since an increase in the curing temperature of concrete increases its r
ate of  development of strength, the gain of strength can be speeded up by curing concrete in  steam. When steam is at atmospheric pressure, i.e. the temperature is below 100 C (212 F),  the process can be regarded as a special case of moist curing. High-pressure steam curing  (autoclaving) is an entirely different operation and is considered in t
he next section. Steam  curing has been used successfully with different types of Portland cement, but must  never be used with high-alumina cement because of the adverse effect of hot-wet  conditions on the strength of that cement. Concrete with a lower water / cement ratio  responds to steam curing much better than a weaker mix.   The primary ob
ject of steam curing is to obtain a sufficiently high early  strength so that the concrete products may be handled soon after casting: the moulds can  be removed, or the prestressing bed vacated, earlier than would be the case with ordinary moist curing,  and less curing storage space is required; all these mean an economic advantage. For many app
lications,  the long-term strength of concrete is of lesser importance.   Because of the nature of the operations involved in steam curing, the process is used mainly with  precast products. Low-pressure steam curing is normally applied in special chambers or in tunnels  through which the concrete members are transported on a conveyor belt. Altern
atively, portable boxes  or plastic covers can be placed over precast members, steam being supplied through flexible  connections.   Owing to the influence of temperature during the early stages of hardening on the later  strength, a compromise between the temperatures giving a high early and a high late  strength has to be made. Fig. 5.50 shows t
ypical values of strength of concrete made  with modified cement and a water / cement ratio of 0 55; steam curing was applied  immediately after casting .    Fig. 5.50. Strength of concrete cured in steam at different temperatures (water / cement  ratio = 0.50; steam curing applied immediately after casting)     One factor in the retrogression of 
strength of steam-cured concrete is the  pressure developed in the air pores during heating, since air has a higher coefficient  of thermal expansion than the surrounding solid material. Some of this effect can be  avoided by heating in closed formwork or in pressure chambers. With short-term curing  periods (2 to 5 hours) and moderate temperature
s, there is probably little real  retrogression of strength, and the apparent low strength at later ages is due to the  absence of prolonged wet curing.   A related problem is the rate of increase in temperature at the commencement of steam curing. It has  been found that if 49 C (120 F) is reached in a period shorter than about 2 to 3 hours, or 9
9 C (210 F) in  less than 6 to 7 hours from the time of mixing, the gain of strength beyond  the first few hours is affected adversely; such a rapid rise in temperature must not  be permitted. The values quoted are merely a guide, but excessively rapid heating can  cause a loss of strength at later ages of as much as one-third compared with wet cu
ring  at room temperature. The adverse effect of the rapid rise is more pronounced the higher  the  water / cement ratio of the mix, and is also more noticeable with rapid  hardening than with ordinary Portland cement. Saul found that, when the rate of rise  in the temperature of concrete does not exceed the values mentioned earlier, its  strength
 will differ only little from the strength of normally cured concrete, and will  fall within zone A of Fig. 5.51. By contrast, the strength of concrete heated too  rapidly will lie within zone B of the same Figure.    Fig. 5.51 Gain of strength of steam-cured concrete with time (water / cement ratio =  0.50, rapid-hardening Portland cement)     Si
nce it is the temperature at the time of setting that has the greatest influence on  the strength at later ages, a delay in the application of steam curing is advantageous.  Some indication of the influence of the delay in heating on strength can be obtained  from Fig. 5.52 plotted by Saul from the data of Shideler and Chamberlin. The  concrete us
ed was made with modified cement, and had a water / cement ratio of 0.6. The  solid line shows the gain in strength of moist-cured concrete at room temperature  plotted against maturity. The dotted lines refer to different curing temperatures  between 38 and 85 C (100 and 185 F), and the figure against each point denotes the  delay in hours before
 the higher curing temperature was suddenly applied.    Fig. 5.52. Effect of delay in steam curing on the early gain of strength with maturity     From Fig. 5.52 it can be seen that for each curing temperature there is a  part of the curve showing a normal rate of gain in strength with maturity. In other  words, after a sufficient delay, rapid hea
ting has no adverse effect. This delay is  approximately 2, 3, 5 and 6 hours respectively for 38, 54, 74 and 85 C (100, 130, 165  and 185 F). If, however, concrete is exposed to the higher temperature with a smaller  delay, the strength is adversely affected, as shown by the right-hand portion of each  dotted curve; this effect is more serious the
 higher the curing temperature.   Fig.	5.52 shows also that within a few hours of casting the rate of gain in strength is  higher than would be expected from the maturity calculations. This confirms the earlier  observation that the age at which a higher temperature is applied is a factor in the  maturity rule.   Practical curing cycles are chosen
 as a compromise between the early and late strength  requirements but are governed also by the time available (e.g. length of work shifts).  Economic considerations determine whether the curing cycle should be suited to a given  concrete mix or alternatively whether the mix ought to be chosen so as to fit a  convenient cycle of steam curing. Whil
e details of an optimum curing cycle depend on the  type of concrete product treated, a typical satisfactory cycle would consist of the  following: a delay period of 3 to 5 hours, heating at the rate of 22 to 33 C  per hour (40 to 60 F per hour) up to a maximum temperature of 66 to 82 C (150 to  180 F), then storage at maximum temperature, possibl
y followed by a period of "soaking"  when no heat is added but the concrete takes in the residual heat and moisture, and  finally a cooling period (at a moderate rate), the total cycle (exclusive of the delay  period) occupying preferably not more than 18 hours. Lightweight aggregate concrete can  be heated up to between 82 and 88 C (180 and 190 F
), but the optimum cycle is no  different from that for concrete made with normal weight aggregate.   The temperatures quoted are those of steam but not necessarily the same as those of the  concrete being processed. During the first hour or two after placing in the curing  chamber the temperature of the concrete lags behind that of the air but la
ter on, owing  to the heat generated by the reactions of hydration, the temperature of the concrete is  higher than that of the surrounding air. Maximum use can be made of the heat stored in  the curing chamber if steam is shut off early and a prolonged cooling period is allowed.  Thus an efficient curing programme would include a period of slowly
 rising temperature,  a period at the maximum temperature, and a period of cooling.   The temperature history of the interior of the concrete being cured is not the same as  that at the surface. The rise in temperature at the centre is slower but the rate of  cooling is lower, too. Thus the area under the temperature - time curve is approximately 
 the same for the interior and for points near the surface of the concrete block, so that  all parts of the concrete have the same maturity. This was demonstrated by Ross, and  Fig. 5.53 shows the computed temperature - time curves for a long block of  concrete subjected to a variation in surface temperature, the effects of the heat of  hydration 
being ignored.	Fig. 5.53. Ross's example of two-dimensional flow of heat in concrete with a  diffusivity of 0.0037 m2 / h (0.04 ft2 / h) (by computation)     In passing, it may be observed that, when concrete is cured at high temperatures, the  heat of hydration of cement is evolved rapidly so that the rise in temperature is  increased even in 
small specimens. On the other hand, with curing at ordinary  temperatures the effects of the heat of hydration are significant only in mass  structures.   It may be worth emphasizing that a longer period of curing at a lower temperature leads  to a higher optimum strength than when high temperature is applied for a shorter time.  For any one perio
d of curing, there is a temperature which leads to an optimum strength.  Also, for a given set of materials, it is possible to draw a line between the optimum  strength at various curing periods and the curing temperature; this is shown in Fig.  5.54.    Fig. 5.54. Strength development of concrete at different curing temperatures for  various peri
ods of curing	In addition to steam curing, other means of applying higher temperature to the concrete  have also been used. For instance, some attempts have been made to cure concrete blocks  by combustion gases, with or without humidification; less drying after curing is  required than with steam curing but the colour of blocks is affected by
 the gases. A  totally different approach uses electrical methods of heating by passing current  through the reinforcement (or through the prestressing steel in the long line process)  or directly through the concrete; this has been found successful but is generally  expensive. The current must be alternating as direct current would lead to hydrol
ysis of  the cement paste. The temperature rise should be gradual, about 10 C (18 F) per hour, up  to not more than 80 C (176 F). A period of several hours at this temperature should be  followed by slow cooling. Yet another method uses infra-red radiation of concrete slabs  slowly moving through a tunnel equipped with infra-red generators. With a
ll these methods,  it is important to avoid excessive drying of the concrete.      High-pressure Steam Curing (Autoclaving)    This process is quite different from curing in steam at atmospheric pressure, both in  the method of execution and in the nature of the resulting product.   Since pressures above atmospheric are involved, the curing chambe
r must be of the  pressure vessel type with a supply of wet steam; excess water is necessary as  superheated steam must not be allowed to come into contact with the concrete. Such a  vessel is known as an autoclave, and in the American literature high-pressure steam  curing is often referred to as autoclaving.   High-pressure steam curing was firs
t employed in the manufacture of sand-lime brick,  and is still extensively used for that purpose. In the field of concrete, high-pressure  steam curing is usually applied to precast products  (made both of ordinary and lightweight concrete) when any of the following  characteristics are desired:   (a) high early strength: with high-pressure steam
 curing, the 28-day strength on normal  curing can be reached in about 24 hours;   (b) high durability: high-pressure steam curing improves the resistance of concrete to  sulphates and to other forms of chemical attack, also to freezing and thawing, and  reduces efflorescence; and   (c) reduced drying shrinkage and moisture movement.   The optimum
 curing temperature has been found experimentally to be about 177 C (350 F)  which corresponds to a steam pressure of 0.8 MPa (120 psi) above atmospheric  pressure.   High-pressure steam curing is most effective when finely ground silica is added to the  cement, owing to the chemical reactions between the silica and Ca(OH)2 released on  hydration 
of C3S (see Fig.	5.55). Cements rich in C3S have a greater capacity for developing high  strength when cured at high pressure than those with a high C2S content, although for short periods of  high-pressure steam curing cements with a moderately low C3S / C2S ratio give good results.    Fig. 5.55. Influence of pulverized silica content on the stre
ngth of high-pressure  steam-cured concrete (age at commencement of curing, 24 hours; curing temperature,  177 C (350 F)     The fineness of the silica should be of the same order as that of the cement, and the  two materials must be intimately mixed before they are added to the mixer. The optimum  amount of silica depends on the mix proportions b
ut is generally between 0 4 and 0 7 of  the weight of cement.  This makes the lime / silica ratio of the mixture equal to approximately 1. The high  temperature during curing affects also the reactions of hydration of the cement itself.  For instance, some of the C3S may hydrate to C3SHx.   High temperature produces a hydrated cement paste of low 
specific surface, about 7000  m2 / kg. This means that the products of hydration are coarse and largely  microcrystalline. Since the specific surface of high-pressure steam-cured paste is only  about 1/20 of that of cement cured at ordinary temperature it appears that no more than 5  per cent of the high-pressure cured paste can be classified as g
el.   Because of the microcrystalline character of the cement paste, high-pressure steam-cured  concrete has a considerably reduced shrinkage, about 1/6 to 1/3 of that of concrete cured at  normal temperatures. When silica is added to the mix, shrinkage is higher, but still  only about one-half of the shrinkage of normally cured concrete. By contr
ast, since  low-pressure steam curing does not produce a microcrystalline paste, no reduction in  shrinkage is obtained. Creep is also significantly reduced by high-pressure steam  curing.   The products of hydration of cement subjected to high-pressure steam curing, as well as  those of the secondary lime - silica reactions, are stable, and there
 is no retrogression  of strength. At the age of one year the strength of normally cured concrete is  approximately the same as that of high-pressure steam-cured concrete of similar mix  proportions. The water / cement ratio affects the strength of high-pressure steam-cured  concrete in the usual manner, but the actual values of early strength dif
fer, of course,  from those for ordinary curing. The coefficient of thermal expansion and the modulus of  elasticity of concrete seem unaffected by high-pressure steam curing.   High-pressure steam curing improves the resistance of concrete to sulphate attack. This  is due to several reasons, the main one being the formation of aluminates more sta
ble in  the presence of sulphates than those formed at lower temperatures. For this reason, the  relative improvement in resistance to sulphate attack is greater in cements with a high  C3A content than in cements relatively resistant to sulphates. Another important factor  is the reduction in lime in the cement paste as a result of the lime - sil
ica reaction.  Further improvement in sulphate resistance is due to the increased strength and  impermeability of the steam-cured concrete, and also to the existence of hydrates in a  well-crystallized form.   High-pressure steam curing reduces efflorescence as there is no lime left to be leached  out.   On the debit side, high-pressure steam curi
ng reduces the strength in bond with  reinforcement by about one-half compared with ordinary curing so that the application of  high-pressure steam to reinforced concrete members is considered inadvisable.  High-pressure steamed concrete tends also to be rather brittle. On the whole, high-pressure  steam curing produces good quality, dense and dur
able concrete. It is whitish in  appearance as distinct from the characteristic colour of normally cured Portland cement  concrete.   It is essential that the rate of heating during high-pressure steam curing is not too  high, as interference with the setting and hardening processes may occur in a manner  similar to that discussed in connection wi
th steam curing at atmospheric pressure. A  typical steaming cycle consists of a gradual increase to the maximum temperature of  182 C (360 F) (which corresponds to a pressure of 1 MPa (140 psi)) over a period of 3  hours. This is followed by 5 to 8 hours at this temperature, and then a release of  pressure in about 20 to 30 min. A rapid release a
ccelerates the drying of the  concrete so that shrinkage in situ will be reduced. At each temperature there is an  optimum period of curing (see Fig. 5.54).   The details of the steaming cycle depend on the plant used and also on the size of the  concrete members being cured. The length of the period of normal curing preceding  placing in the auto
clave does not affect the quality of the steam-cured concrete, and  the choice of a suitable period is governed by the stiffness of the mix, which must be  strong enough to withstand handling. In the case of lightweight concretes, the details  of the steaming cycle have to be determined experimentally to suit the materials used.   Steam curing sho
uld be applied to concretes made with Portland cement only: high-alumina and  supersulphated cement would be adversely affected by the high temperature.   Within the Portland group, the type of cement affects the strength but not necessarily  in the same way as at normal temperatures; no systematic studies have, however, been  made. It is known, t
hough, that granulated slag may cause trouble if it has a high  sulphur content. High-pressure steam curing accelerates the hardening of concrete  containing calcium chloride, but the relative increase in strength is less than when no  calcium chloride is used.      Variation in Strength of Cement    Up to now we have not considered the strength o
f cement as a variable in the strength of  concrete. By this we do not mean the differences in the strength-producing properties of  cements of different types, but the variation between cements of nominally the same  type: they vary fairly widely, and it is this variation that is considered in this  section.   The scatter of strength of ordinary 
and rapid hardening Portland cements is illustrated  in Fig. 5.56, which shows also that there is a considerable overlap in the strengths of  the two cements. However, these cements nearly always comply with the minimum  requirements of ASTM Standard C 15-78a or of BS 12: 1978, the strength of some of the  cements sometimes reaching values as high
 as twice the prescribed minimum.    Fig. 5.56. Histograms for vibrated mortar cubes made with ordinary and rapid hardening  Portland cements as defined in BS 12: 1958     Although on a site it is difficult to isolate the influence of cement, there is no doubt  that the inherent variation in the strength of cement is reflected in the variation in 
 the strength of concrete. Thus the variability of cement is of considerable practical  importance, particularly since a concrete mix has to be designed so as to give a  satisfactory strength even when a low-strength batch of cement is encountered: no  advantage can be taken of the much higher mean strength. If cement had a lower  variability (but
 the same mean strength) the cement content of the mix could be reduced:  there may thus be some advantage in strength grading of cement even if this were  accompanied by a higher price of the material.   The variation in strength of cement is due largely to the lack of uniformity in the raw  materials used in its manufacture, not only between dif
ferent sources of supply, but  also within a pit or a quarry. Furthermore, differences in details of the processes of  manufacture and above all the variation in the ash content of coal used to fire the kiln  contribute to the variation in the properties of commercial cements.   The magnitude of the variation in the strength of cement can be judge
d from Fig. 5.57,  which shows the results of tests on ASTM standard mortar made with samples of cement  obtained from the same plant in the United States at two-week intervals. The strength is  expressed as a percentage of the mean strength of all the samples from the given works,  and each curve is an average of strength ratios obtained at 3, 7 
and 28 days. The variation in strength  due to testing per se is indicated by the dotted lines which show the strength ratios for tests made at the  same time using a control "stock" cement. The testing error accounts generally  of between 2 1/2 and 4 per cent. ASTM Standard C 917-80 gives a method of assessing the  uniformity of cement from a sin
gle source.    Fig. 5.57 Variation in strength of cement from the same plant     Studies of cement plants in California have shown that the coefficients of variation  of standard mortar cubes made with cement from any one given plant are as follows:-   20 per cent of plants have a coefficient of variation smaller than 6 per cent,   72 per cent - s
maller than 10 per cent,   28 per cent - 10 per cent or greater.   It may be interesting to quote some German data on the variability of cement as  tested by spot (grab) sampling. Usually 200 to 300 samples are tested in one plant in  any year, six cubes (of 1 : 3 mortar with a water / cement ratio of  0 5) per sample being crushed. At a typical p
lant, the  ordinary Portland cement (minimum 28-day strength of 35 MPa (5000 psi))  had an average 28-day strength of 49.5 MPa (7200 psi) with a standard deviation of 22  MPa (320 psi). This value of scatter includes the testing error, which may well account  for some 1.5 MPa (220 psi). It should be noted that German standards lay down a maximum  
strength of cement, this being 20 MPa (3000 psi) above the specified minimum. An appendix to  ASTM Standard C 917-80 gives data on the typical variability of cement made in U.S. and Canadian  plants.   The standard deviation of the strength of concrete due to the variation in  cement does not increase with age; indeed, German data suggest a  decre
ase of up to 1 MPa (150 psi). Since strength increases with age, the  coefficient of variation of strength becomes smaller the older the concrete at  the time of testing. This behaviour is not surprising because a large part  of the variation in the strength of cement is due to the differences in fineness and in  the C3S content: the effects of th
ese factors are greatest at early ages and with time  cease to be significant. By contrast, the standard deviation within batches increases  with an increase in mean strength. Table 5.5 gives Wright's data on the variability  of cement at different ages at test, and Table 5.6 gives approximate values of the  standard deviation of strength of site-
made concrete.    Table 5.5: Variation in the Strength of Mortar and Concrete Cubes made with Different  Batches of Cement from One Plant    Table 5.6: Standard Deviations of Strength of Concrete resulting from Variations in Cement     From the figures quoted it is easy to see that the use of cement from different  sources, or even the use of diff
erent batches of cement from one plant, leads to an  appreciable variation in the strength of concrete. This effect may be of considerable  importance on a large job: the use of cement all from one batch can result in a decrease  in cement content of up to 10 per cent. It must not be forgotten, however, that  variation in cement accounts, at the m
ost, for one-half of the variation in the strength  of site test specimens; U.S. Bureau of Reclamation data suggest a typical value of  one-third. A major part of this variation can be eliminated by the use of cement from  one silo at the plant. The variation in the strength of site cubes is discussed on page  653.   Finally it should be stressed 
that the variation in cement affects to the greatest  extent the early strength of concrete, i.e. the strength most often  determined by test but not necessarily the strength of greatest practical significance.  Furthermore, strength is not the only important characteristic of concrete: from  considerations of durability and impermeability, a ceme
nt content in excess of that  needed for strength may well be required, in which case the variability of cement  becomes unimportant.      Fatigue Strength of Concrete    In this chapter, we have considered so far only the strength of concrete under static  loading. In many structures, however, repeated loading is applied, and when a material  fai
ls under a number of repeated loads, each smaller than the static compressive  strength, failure in fatigue is said to take place. Both concrete and steel possess the  characteristics of fatigue failure but in this book the behaviour of concrete alone is  dealt with.   Let us consider a concrete specimen subjected to alternations of compressive st
ress  between values ol (> = 0) and oh (> ol). The stress - strain curve varies with the number of  load repetitions, changing from concave toward the strain axis (with a hysteresis loop  on unloading) to a straight line, which shifts at a decreasing rate (i.e., there is some  irrecoverable set) and eventually to concave toward the stress axis. Th
e degree of this  latter concavity is an indication of how near the concrete is to failure. For practical  loadings, failure will, however, take place only above a certain limiting value of oh,  known as fatigue limit or endurance limit. If oh is below the fatigue limit the  stress - strain curve will indefinitely remain straight, and failure in f
atigue will not take  place. The changes in the stress - strain curve with the number of applied cycles are  illustrated in Fig. 5.58.    Fig. 5.58. Stress - strain relation of concrete under cyclic compressive loading     If the stress - strain curve for unloading is also drawn, a hysteresis loop in each cycle  can be seen. The area of this loop 
decreases with each successive cycle and then eventually increases  prior to fatigue failure. There does not seem to be such an increase in specimens which do not fail in  fatigue. If we plot the area of each successive hysteresis loop as a percentage of the area of the first  loop, the variation with the number of cycles is as shown in Fig. 5.59.
	Fig. 5.59. Variation in the area of the hysteresis loop as a percentage of the first loop with the  number of cycles     The interest in the hysteresis loop arises from the fact that its area represents an  irreversible energy of deformation, and is manifested by a rise in temperature of the  specimen. The irreversible deformation involved may
 be either in the form of crack  formation or irreversible deformation without a loss of continuity, such as viscous  flow. Pulse velocity measurements have shown that it is the development of cracks  that is responsible for the change in behaviour near failure, but it is possible that  even earlier there is a slow and stable propagation of one or
 more cracks, although  there is some non-elastic deformation without disruption of continuity of the material.   Fatigue cracking is extensive and the observed strain at failure is much larger than in  static failure. The non-elastic strain increases with the number of cycles, as shown in  Fig. 5.60 and can sometimes be as high as 4000 x 10'-6 af
ter 13 million cycles at 200  cycles per minute. Generally, the specimen with a longer fatigue life has a higher  non-elastic strain at failure (Fig. 5.61).    Fig. 5.60. Variation in non-elastic strain with the number of cycles    Fig. 5.61. Relation between non-elastic strain near failure and number of cycles at failure     The elastic strain al
so increases progressively with cycling. This is shown in Fig.  5.62 by the reduction in the secant modulus of elasticity with an increase in the  percentage of the "fatigue life" used up. This relation is independent of the level of  stress in the fatigue test and is therefore of interest in assessing the remaining  fatigue life of a given concre
te.    Fig. 5.62. Relation between the ratio of the secant modulus of elasticity at the given  instant (E) to the modulus at the beginning of cycling (E0) and the percentage of  fatigue life used up     The lateral strain is also affected by the progress of cyclic loading, the Poisson's  ratio decreasing progressively.   Cyclic loading below the f
atigue limit improves the fatigue strength of  concrete, i.e. concrete loaded a number of times below its fatigue limit  will, when subsequently loaded above the limit, exhibit a higher fatigue strength than  concrete which had never been subjected to the initial cycles. The former concrete also  exhibits a higher static strength by some 5 to 25 p
er cent. It is possible that this  increase in strength is due to a densification of concrete caused by the initial  low-stress level cycling, in a manner similar to improvement in strength under moderate  sustained loading. This property is akin to strain hardening in metals, and is of  particular interest since concrete under static loading is a
 strain-softening rather  than strain-hardening material.   Strictly speaking, concrete does not appear to have a fatigue limit, i.e. a fatigue  strength at an infinite number of cycles (except when stress reversal  takes place). It is therefore to refer to fatigue strength at a very large  number of cycles, such as 10 million.   The fatigue stren
gth can be represented by means of a modified Goodman diagram (see  Fig. 5.63). The ordinate from a line at 45 through the origin shows the range of  stress (oh - ol) for a given number of cycles; ol is generally greater than zero  arising from the dead load, while oh is due to the dead plus live (transient) load. Thus  the range of stress that a 
given concrete can withstand a specified number of cycles can  be read off the diagram. For a given ol, the number of cycles is very sensitive to the  range of stress. For instance, an increase in range from 57.5 to 65 per cent of the  ultimate static strength has been found to decrease the number of cycles by a factor of  40.    Fig. 5.63. Modifi
ed Goodman diagram for concrete in compression fatigue (N is  number of repetitions)     The modified Goodman diagram (see Fig. 5.63) shows that for a constant range of stress,  the higher the value of the minimum stress the lower the number of cycles that a given  concrete can withstand. This is of significance in relation to the dead load of a c
oncrete member  which is to carry a transient load of a certain magnitude.   From the fact that the lines of Fig. 5.63 rise to the right it can also be seen that  the fatigue strength of concrete is lower the higher the ratio oh / ol. Some tests have  shown that lateral pressure increases the fatigue life of concrete, but not at very high stresses
.   The frequency of the alternating load, at least within the limits of 70 to 2000 cycles  per minute, does not affect the resulting fatigue strength, higher frequency is of  little practical significance. This applies both in compression and in flexure-tension,  the similarity between fatigue behaviour in the two types of loading, as well as in 
 splitting tension, suggesting that the failure mechanism is the same. In fact, the  fatigue behaviour in flexure parallels closely that in compression (Fig. 5.64). The  fatigue strength for 10 million cycles is 55 per cent of the static strength. By  comparison, in compression fatigue, the fatigue strength in flexure is between 60 and 64  per cen
t after the same number of cycles. Sufficiently accurate test results are not  available to state with certainty that these two limits (in compression and in flexure)  differ significantly from one another. Tests have shown that the moisture condition of  concrete prior to loading affects its fatigue strength in flexure: oven-dried specimens  show
 the highest strength and partially dried ones the lowest; wet specimens are in  between (Fig. 5.65). The explanation of this behaviour lies in differential strains  induced by the moisture gradients.    Fig. 5.64. Modified Goodman diagram for concrete in flexure fatigue    Fig. 5.65. Effect of moisture condition on fatigue performance of concrete
 specimens     As strength increases with age, fatigue strength both in compression and in flexure  also increases. The important point is that at a given number of cycles, fatigue  failure occurs at the same fraction of ultimate strength, and is thus independent of the  magnitude of this strength (both in compression and in splitting tension) and
 of age of  concrete, although some tests suggest an increase in fatigue life with age. It can  thus be seen that a single parameter is critical in fatigue failure. Murdock expressed  the view that the deterioration of bond between the cement paste and aggregate is  responsible for this failure. Tests have shown that fatigue specimens had fewer br
oken  aggregate particles than specimens which failed in a static test. Thus failure at the  bond interface is probably dominant in fatigue; in mortar, fatigue failure is believed  to take place at the interface of the fine aggregate particles. It is likely that a  smaller maximum size of aggregate leads to a higher fatigue strength, probably  bec
ause of greater homogeneity of concrete.   Air-entrained concrete and lightweight aggregate concrete have the same fatigue  behaviour as concrete made with ordinary aggregate. Fatigue in concrete  cylinders occurs in the same way as in large specimens subjected to low-frequency  loading.   The fatigue strength of concrete is increased by rest peri
ods (this does not apply when  there are stress reversals), the increase being proportional to their duration between 1  and 5 minutes; beyond the 5-minute limit there is no further increase in strength. With  the rest periods at their maximum effective duration, their frequency determines the  beneficial effect. The increase in strength caused by
 rest periods is probably due to  relaxation of concrete (primary bonds, which remained intact, restoring the internal  structure to its original configuration), as evidenced by a decrease in the total  strain; this decrease occurs rapidly after the cessation of cycling.   Murdock suggested that fatigue failure occurs at a constant strain, indepen
dent of  the applied stress level or of the number of cycles necessary to produce failure. This  fatigue ultimate strain is larger than the static failure strain. Nevertheless, this  behaviour of concrete would add further support to the concept of ultimate strain as  failure criterion.   In an eccentrically loaded specimen, there are elements str
essed to levels lower than  the maximum stress, and therefore a greater number of cycles can be sustained than in a  uniformly stressed specimen subjected to the same maximum stress. An increase of 17 per  cent was observed when the eccentricity was 25 mm (1 in.) in a 150 mm (6 in.) face.   Most fatigue tests are conducted under cyclic loading of 
the same shape. Variable  amplitude loading is more harmful than constant amplitude loading so that Miner's  hypothesis of linear accumulation of damage does not seem to apply to concrete.  The stress level and the sequence of cycles appear to be the significant factors.   Finally, it may be of interest to observe that the behaviour of concrete un
der fatigue  loading is not closely related to behaviour under impact loading, although of course  both may occur simultaneously in practice. Failure under fatigue loading occurs at a  strength below the static compressive strength while under impact the static strength is  not impaired.   While this book is not concerned with the fatigue behaviou
r of reinforced and  prestressed concrete, we should note that fatigue cracks in concrete act as  stress-raisers, thus magnifying the vulnerability of the steel to fatigue failure (if the  stress in it is in excess of its critical fatigue stress value).      Impact Strength    Impact strength is of importance primarily in connection with pile driv
ing and with  foundations for machines exerting impulsive loading, but accidental impact (e.g. during  handling of precast members) is also of interest. Extensive tests on impact strength of  concrete were made by Green. As principal criteria he considered the ability of a  specimen to withstand repeated blows and to absorb energy. In particular, 
he studied the  number of blows which the concrete can withstand before reaching the no-rebound  condition, this stage indicating a definite state of damage.   Impact tests, when conducted with a relatively small hammer (25 mm (1 in.) diameter  face) lead to a greater scatter of results than tests on static compressive strength of  the concrete. T
his arises from the fact that in the standard compression test some  relief of a highly stressed weak zone is possible owing to creep, while in the impact  test no redistribution of stresses is possible during the very short period of  deformation. Hence, local weaknesses have a greater influence on the recorded strength  of a specimen .   In gene
ral, Green found that the higher the static compressive strength of the concrete  the lower the energy absorbed per blow before cracking, but the impact strength of  concrete increases with its compressive strength (and therefore age) at a progressively  increasing rate (Fig. 5.66). The  relation is different for each coarse aggregate and  storage
 condition of the concrete. For the same compressive strength, the impact  strength is greater  for coarse aggregate of greater angularity and surface roughness.  This observation was confirmed by Dahms and supports the suggestion that  impact strength is more closely related to the tensile strength of concrete than to its  compressive strength. T
hus concrete made with a  gravel coarse aggregate has a low  impact strength, failure taking place owing to insufficient bond between mortar and  coarse aggregate. On the other hand, when the surface of the aggregate is rough, the  concrete is able to develop the full strength of much of the aggregate in the region of  failure.    Fig. 5.66. Relat
ion between compressive strength and number of blows to  "no-rebound" for concretes made with different aggregates and Type I cement,  stored in water     A smaller maximum size of aggregate significantly improves impact resistance; so does  the use of aggregate with a low modulus of elasticity and a low Poisson's ratio.  Cement content below 400 
kg / m3 (670 lb / yd3) is  advantageous. The influence of fine  aggregate is not well defined but the use of fine sand usually leads to a slightly  lower impact strength. Dahms found a high content of sand advantageous. We could  try to generalize and say that a mix of materials which have a limited variation in  properties is conducive to a good 
impact strength. Extensive tests on the impact  strength of concretes of different properties were made by Hughes and Gregory.   Storage conditions influence the impact strength in a manner different  from  compressive strength. Specifically, the impact strength of water stored concrete is  lower than when the concrete is dry, although the former 
concrete can withstand more  blows before cracking. Thus, as already stated, the compressive strength without  reference to storage conditions, does not give a satisfactory indication of the impact  strength.   There is evidence that under uniformly applied impact loading (a condition difficult  to achieve in practice) the impact strength of concr
ete is significantly greater than  its static compressive strength. This increase in  strength would explain the greater  ability of concrete to absorb strain energy under uniform impact. Fig. 5.67 shows that  strength increases greatly when the rate of application of stress exceeds about 500  GPa / s,  reaching at 4.9 TPa / s more than double the
 value at normal speeds of loading  (about 0 5 MPa / s).    Fig. 5.67. Relation between compressive strength and rate of loading up to impact level     A parallel relation is that between strength and the rate of application of strain  shown in Fig. 5.68. In these tests, the rate of application of stress was lower  than in those shown in Fig. 5.67
;	for instance, the rate of strain of 14 per sec  corresponds to a rate of stress of 400 MPa / s (58 000 psi / s). The tests of Fig. 5.68 give  a lower increase in impact strength than Popp's tests, probably because in the  former the platen friction was greatly reduced by the use of pads    Fig. 5.68. Relation between compressive strength and rat
e	of strain in impact tests for  the data of reference      Quality of Mixing Water    The vital influence of the quantity of water in the mix on the strength of the resulting  concrete has been repeatedly mentioned. The quality of the water also plays its role:  impurities in water may interfere with the setting of the cement, may adversely affec
t	 the strength of the concrete or cause staining of its surface, and may also lead to  corrosion of the reinforcement. For these reasons, the suitability of water for mixing  and curing purposes should be considered. Clear distinction must be made between the  effects of mixing water and the attack on hardened concrete by aggressive waters. Some 
 of the latter type of water may be harmless or even beneficial when used in mixing.   In many specifications, the quality of water is covered by a clause saying that water  should be fit for drinking. Such water very rarely contains dissolved solids in excess  of 2000 parts per million (ppm), and as a rule less than 1000 ppm. For a water / cement
	ratio of 0.5, the latter content corresponds to a quantity of solids representing 0.05  per cent of the weight of cement, and any effect of the common solids would be small.  There is, however, one situation when drinking water is unsuitable as mixing water:  this is when there is a danger of alkali - aggregate reaction and the water has a high 
 concentration of sodium or potassium.   While the use of potable water is generally safe, water not fit for drinking may  often also be satisfactorily used in making concrete. As a rule, water with pH of 6 0  to 8.0 which does not taste saline or brackish is suitable for use, but dark  colour or bad smell do not necessarily mean that deleterious 
substances are present.  A simple way of determining the suitability of such water is to compare the setting  time of cement and the strength of mortar cubes using the water in question with the  corresponding results obtained using known "good" water or distilled water; there is no  appreciable difference between the behaviour of distilled and or
dinary drinking water.  A tolerance of about 10 per cent is usually  permitted to allow for the chance  variations in strength; an appendix to BS 3148:1980 also suggests 10 per cent. Such  tests are recommended when water for which no service record is available contains  dissolved solids in  excess of 2000 ppm or, in the case of alkali carbonate 
or  bicarbonate, in excess of 1000 ppm. When unusual solids are present a test is also  advisable.   Since it is undesirable to introduce large quantities of silt into the concrete, mixing  water with a high content of suspended solids should be allowed to stand in a settling  basin before use; a turbidity limit of 2000 ppm has been suggested. How
ever, water  used to wash out truck mixers is satisfactory as mixing water, provided of course that  it was satisfactory to begin with. ASTM Standard C 94-78a allows the use of wash water.  Clearly, different cements and different admixtures should not be involved.   Brackish water contains chlorides and sulphates. When chloride does not exceed 50
0  ppm, or SO3 does not exceed 1000 ppm, the water is harmless, but water with even higher  salt  contents has been used satisfactorily. The Appendix to BS 3148: 1980 recommends limits  on chloride and on SO3 as above, and also recommends that alkali carbonates and  bicarbonates should not exceed 1000 ppm. Somewhat less severe limitations are  rec
ommended in American literature.   Sea water has a total salinity of about 3.5 per cent (78 per cent of the dissolved  solids being NaCl and 15 per cent MgCl2 and MgSO4), and produces a slightly higher  early strength but a lower long-term strength; the loss of strength is usually no more than  15  per cent and can therefore often be tolerated. So
me tests suggest that sea water  slightly accelerates the setting time of cement, others, a substantial reduction in  the initial setting time but not necessarily in the final set. Generally, the effects on  setting are unimportant if water is acceptable from strength considerations. An appendix  to BS 3148: 1980 suggests a tolerance of 30 minutes
 in the initial setting time.   Water containing large quantities of chlorides (e.g. sea water) tends to cause  persistent dampness and surface efflorescence. Such water should, therefore, not be used  where appearance of the concrete is of importance, or where a plaster finish is to be  applied.   In the case of reinforced concrete, sea water is 
believed to increase the risk of  corrosion of the reinforcement, although there is no experimental evidence that the use  of sea water in mixing leads to attack on the reinforcing steel. The danger is  believed to be greater in tropical countries. Corrosion has been observed in structures  exposed to humid air when the cover to reinforcement is i
nadequate or the concrete is  not sufficiently dense so that the corrosive action of residual salts in the presence of  moisture can take place. On the other hand, when reinforced concrete is permanently in  water, either sea or fresh, the use of sea water in mixing seems to have no ill-effects.  However, in practice it is generally considered ina
dvisable to use sea water  for mixing unless this is unavoidable. In prestressed concrete the use of sea water  is not permitted because the small cross-section of the wire means that the effects of  corrosion are relatively more serious.   In this connection, the use of aggregate won from the sea should be considered. Sand  dried out in sea water
 may contain a large amount of salt, but if sand dredged from the  sea is washed in sea water and is allowed to drain, and fresh water is used as mixing  water, then the salt content represents no more than 1 per cent of the total weight of  water (see p. 154).   Natural waters that are slightly acid are harmless, but water containing humic or oth
er  organic acids may adversely affect the hardening of concrete; such water, as well as  highly alkaline water, should be tested. The effects of different ions vary, as shown by  Steinour.   It may be interesting to note that the presence of algae in mixing water results in air  entrainment with a consequent loss of strength. The use of algae is,
 however, hardly a  practical means of air entrainment. Hardness of water does not affect the efficiency of  air-entraining admixtures.   As far as curing of concrete is concerned, water satisfactory for mixing is also  suitable for curing purposes. However, iron or organic matter can cause staining,  particularly if water flows slowly over concre
te and evaporates rapidly.   Whether or not staining will take place cannot be stated on the basis of chemical  analysis and should be checked by a performance test. U.S. Army Engineers recommend  a preliminary test in which 300 ml of the water to be used for curing is evaporated from a  slight depression, 100 mm (4 in.) diameter, in the surface o
f a specimen of neat white  cement or plaster of Paris. If the resulting colouring is not considered objectionable a  further test is performed. Here, 150 litres (40 U.S. gallons) of water are allowed to flow  lengthwise over a 150 by 150 by 750 mm (6 by 6 by 30 in.) concrete beam with a  channel-shaped top surface, placed at 15 to 20 to the horiz
ontal; the rate of flow is 4 litres  in 3 to 4 hours. Forced circulation of air and heating by electric lamps encourage  evaporation and thus deposition of the residue. The test is again evaluated by  observation only, and if necessary an actual field test in which a 2 m2 (or 20 ft2) slab  is cured may b performed.   In some cases discoloration is
 of no significance, and any water suitable for mixing,  or even slightly inferior in quality, is acceptable for curing.   It is essential that curing water be free from substances that attack hardened  concrete; these are discussed in Chapter 7.  4 Fresh Concrete    The strength of concrete of given mix proportions is very seriously affected  by 
the degree of its compaction; it is vital, therefore, that the consistence of the mix  be such that the concrete can be transported, placed and finished sufficiently easily and  without segregation.      Definition of Workability    A concrete satisfying these conditions is said to be workable, but to say merely that  workability determines the ea
se of placement and the resistance to segregation is too loose  a description of this vital property of concrete. Furthermore, the desired workability in any  particular case would depend on the means of compaction available; likewise, a  workability suitable for mass concrete is not necessarily sufficient for thin, inaccessible, or  heavily reinf
orced sections. For these reasons, workability should be defined as a physical  property of concrete alone without reference to the circumstances of a particular type  of construction.   To obtain such a definition it is necessary to consider what happens when  concrete is being compacted. Whether compaction is achieved by ramming or by  vibration
, the process consists essentially of the elimination of  entrapped air from the concrete until it has achieved as close a configuration as is  possible for a given mix. Thus, the work done is used to overcome the friction between  the individual particles in the concrete and also between the concrete and the surface of  the mould or of the reinfo
rcement. These two can be called internal friction and surface  friction respectively.  In addition, some of the work done is used in vibrating the mould or  in shock and indeed in vibrating those parts of the concrete which have already been fully  consolidated. Thus the work done consists of a "wasted" part and "useful" work, the  latter, as men
tioned before, comprising work done to overcome the internal friction and  the surface friction. Since only the internal friction is an intrinsic property of the mix,  workability can be best defined as the amount of useful internal work necessary to produce  full compaction. This definition has been developed by Glanville, Collins and Matthews  o
f the Road Research Laboratory who have exhaustively examined the field of compaction  and workability.   Another term used to describe the state of fresh concrete is consistence. In ordinary  English usage this word refers to the firmness of form of a substance or to the  ease with which it will flow. In the case of concrete, consistence is somet
imes taken to  mean the degree of wetness; within limits, wet concretes are more workable than dry  concretes, but concretes of the same consistence may vary in workability.      The Need for Sufficient Workability    Workability has so far been discussed merely as a property of fresh concrete: it is,  however, also a vital property as far as the 
finished product is concerned since concrete  must have a workability such that compaction to maximum density is possible with a  reasonable amount of work or with the amount that we are prepared to put in under given  conditions.   The need for compaction becomes apparent from a study of the relation between  the degree of compaction and the resu
lting strength. It is convenient to express  the former as a density ratio, i.e. a ratio of the actual density of the given concrete to the  density of the same mix if fully compacted. Likewise, the ratio of the strength of the  concrete as actually (partially) compacted to the strength of the same mix when fully  compacted can be called the stren
gth ratio. Then the relation between the strength  ratio and the density ratio is of the form shown in Fig. 4.1. The presence of voids in  concrete greatly reduces its strength: 5 per cent of voids can lower strength by as much as  30 per cent and even 2 per cent voids can result in a drop of strength of more than 10 per  cent. This, of course, is
 in agreement with Feret's expression relating strength to the sum  of the volumes of water and air in the hardened paste (see p. 268).    Fig. 4.1. Relation between strength ratio and density ratio     Voids in concrete are in fact either bubbles of entrapped air or spaces left after  excess water has been removed. The volume of the latter depend
s solely on the  water / cement ratio of the mix. The air bubbles, which represent "accidental" air, i.e.  voids within an originally loose granular material, are governed by the grading of the finest  particles in the mix and are more easily expelled from a wetter mix than from a dry one. It  follows, therefore, that for any given method of compa
ction there may be an optimum  water content of the mix at which the sum of the volumes of air bubbles and water space  will be a minimum. At this optimum water content the highest density ratio of the concrete  would be obtained. It can be seen, however, that the optimum water content may vary for  different methods of compaction.      Factors Af
fecting Workability    The main factor is the water content of the mix, expressed in kilogrammes of water per  cubic metre of concrete: it is convenient, though approximate, to assume that for a  given type and grading of aggregate and workability of concrete, the water content is  independent of the aggregate / cement ratio. On the basis of this 
assumption, the mix  proportions of concretes of different richness can be estimated, and Table 4.1 gives  typical values of water content for different slumps and maximum sizes of aggregate.  These values are applicable to non-air-entrained concrete only. When air is entrained  the water content can be reduced in accordance with the data of Fig. 
4.2. This is  indicative of the order of values only, since the effect of entrained air on workability  depends on the mix proportions, as described in detail on page 484.    Table 4.1: Approximate Water Content for Different Slumps and Maximum Sizes of  Aggregate    Fig. 4.2. Reduction in mixing water requirement due to air-entrainment     If the
 water content and the other mix proportions are fixed, workability is governed  by the maximum size of aggregate, its grading, shape and texture. The influence of these  factors was discussed in Chapter 3. However, the grading and the water / cement ratio  have to be considered  together, as a grading producing the most workable concrete for  one
 particular value of water / cement ratio may not be the best for another value of the ratio.  In particular, the higher the water / cement ratio the finer the grading required for the highest  workability. In actual fact, for a given value of water / cement ratio there is one value of  the coarse / fine aggregate ratio (using given materials) tha
t gives the highest workability.  Conversely, for a given workability, there is one value of  coarse / fine aggregate ratio  which needs the lowest water content The influence of these factors was discussed in  Chapter 3.   It should be remembered, however, that, although when discussing gradings of  aggregate required for a satisfactory workabili
ty proportions by weight were laid  down, these apply only to aggregate of a constant specific gravity. In actual fact,  workability is governed by the volumetric proportions of particles of different  sizes, so that when aggregates of varying specific gravity are used (e.g. in the case  of some lightweight aggregates or mixtures of ordinary and l
ightweight aggregates)  the mix proportions should be assessed on the basis of absolute volume of each size  fraction. This applies also in the case of air-entrained concrete since the entrained  air behaves like weightless fine particles. An example of calculation on absolute  volume basis is given on page 683. The influence of the properties of 
aggregate on  workability decreases with an increase in the richness of  the mix, and possibly  disappears altogether when the aggregate / cement ratio is as low as 2 1/2 or 2.   In practice, predicting the influence of mix proportions on workability requires  care since of the three factors, water / cement ratio, aggregate / cement ratio, and  wa
ter content, only two are independent. For instance, if the aggregate / cement ratio  is reduced, but the water / cement ratio is kept constant, the water content increases, and  consequently the workability also increases. If, on the other hand, the water content is  kept constant when the aggregate / cement ratio is reduced, then the water / cem
ent ratio  decreases but workability is not seriously affected.   The last qualification is necessary because of some secondary effects: a lower  aggregate / cement ratio means a higher total surface area of solids (aggregate and  cement) so that the same amount of water results in a somewhat decreased workability.  This could be offset by the use
 of a slightly coarser grading of aggregate. There are  also other minor factors such as fineness of cement, but the influence of this is  still controversial.      Measurement of Workability    Unfortunately, there is no acceptable test which will measure directly the workability  as defined earlier. Numerous attempts have been made, however, to 
correlate workability  with some easily determinable physical measurement, but none of these is fully satisfactory  although they may provide useful information within a range of variation in workability.    Slump Test    This is a test used extensively in site work all over the world. The slump test does not  measure the workability of concrete b
ut is very useful in detecting variations in the  uniformity of a mix of given nominal proportions.   There are some slight differences in the details of procedure used in different countries,  but these are not significant. The prescriptions of ASTM C 143-78 are summarized  below.   The mould for the slump test is a frustum of a cone, 305 mm (12 
in.) high. It is placed  on a smooth surface with the smaller opening at the top, and filled with concrete in  three layers. Each layer is tamped 25 times with a standard 16 mm (5/8 in.) diameter steel  rod, rounded at the end, and the top surface is struck off by means of a screeding and  rolling motion of the tamping rod. The mould must be firml
y held against its base during  the entire operation; this is facilitated by handles or foot-rests brazed to the mould.   Immediately after filling, the cone is slowly lifted, and the unsupported concrete will  now slump - hence the name of the test. The decrease in the height of the centre of the  slumped concrete is called slump, and is measured
 to the nearest 1/4 in. (5 mm). In order  to reduce the influence on slump of the variation in the surface friction, the inside of  the mould and its base should be moistened at the beginning of every test, and prior to  lifting of the mould the area immediately around the base of the cone should be cleaned  from concrete which may have dropped ac
cidentally.   If instead of slumping evenly all round as in a true slump (Fig. 4.3) one half of the  cone slides down an inclined plane, a shear slump is said to have taken place, and the  test should be repeated. If shear slump persists, as may be the case with harsh mixes,  this is an indication of lack of cohesion in the mix.    Fig. 4.3. Slump
: true, shear, and collapse     Mixes of stiff consistence have a zero slump, so that in the rather dry range no  variation can be detected between mixes of different workability. Rich mixes behave  satisfactorily, their slump being sensitive to variations in workability. However, in a  lean mix with a tendency to harshness a true slump can easily
 change to the shear type,  or even to collapse (Fig. 4.3), and widely different values of slump can be obtained in  different samples from the same mix.   The order of magnitude of slump for different workabilities is given in Table 4.2. It  should be  remembered, however, that with different aggregates the same slump can be recorded for differen
t  workabilities, as indeed the slump bears no unique relation to the workability as defined earlier.    Table 4.2: Workability, Slump, and Compacting Factor of concretes with 19 or 38 mm (3/4  or 1 1/2 in.) Maximum Site of Aggregate     Despite these limitations, the slump test is very useful on the site as a check on the  day-to-day or hour-to-h
our variation in the materials being fed into the mixer. An  increase in slump may mean, for instance, that the moisture content of aggregate has  unexpectedly increased; another cause would be a change in the grading of the aggregate,  such as a deficiency of sand. Too high or too low a slump gives immediate warning and  enables the mixer operato
r to remedy the situation. This application of the slump test,  as well as its simplicity, is responsible for its widespread use.    Compacting Factor Test    There is no generally accepted method of directly measuring workability, i.e. the amount  of work necessary to achieve full compaction. Probably the best test yet available uses  the inverse
 approach: the degree of compaction achieved by a standard amount of work is  determined. The work applied includes perforce the work done against the surface friction  but this is reduced to a minimum, although probably the actual friction varies with the  workability of the mix.  The degree of compaction, called the compacting factor, is measure
d by the density ratio,  i.e. the ratio of the density actually achieved in the test to the density of the same concrete  fully compacted.   The test, known as the compacting factor test, was developed at the Road Research Laboratory  and is described in BS 1881 : Part 2 : 1970 and in ACI Standard 211.3-75. The apparatus consists  essentially of t
wo hoppers, each in the shape of a frustum of a cone, and one cylinder, the three  being above one another. The hoppers have hinged doors at the bottom, as shown in Fig. 4.4.  All inside surfaces are polished to reduce friction.    Fig. 4.4. Compacting factor apparatus     The upper hopper is filled with concrete, this being placed gently so that 
at this stage no work is  done on the concrete to produce compaction. The bottom door of the hopper is then released  and the concrete falls into the lower hopper. This is smaller than the upper one and is, therefore,  filled to overflowing and thus always contains approximately the same amount of concrete in a  standard state; this reduces the in
fluence of the personal factor in filling the top hopper. The bottom  door of the lower hopper is released and the concrete falls into the cylinder. Excess concrete is cut by  two floats slid across the top of the mould, and the net weight of concrete in the known volume of the  cylinder is determined.   The density of the concrete in the cylinder
 is now calculated, and this density divided by the  density of the fully compacted concrete is defined as the compacting factor. The latter density  can be obtained by actually filling the cylinder with concrete in four layers, each tamped or  vibrated, or alternatively calculated from the absolute volumes of the mix ingredients.   The compacting
 factor apparatus shown in Fig. 4.4 is about 1.2 m (4 ft) high. For  concretes with a maximum aggregate size of over 20 mm and up to 40 mm (3/4 in. to 1 1/2 in.)  a "large" apparatus should be employed. Its height is 1.8 m (6 ft) and for this reason  the large apparatus is not used in practice. For the same concrete, the large apparatus  yields a 
somewhat higher value of the compacting factor than the small apparatus.  Unfortunately, the compacting factor apparatus is not often used outside precast  concrete works and large sites.   Table 4.2 lists values of the compacting factor for different workabilities,  as given in Road Note No. 4. Unlike the slump test, variations in the workability
 of  dry concrete are reflected in a large change in the compacting factor, i.e. the test is  more sensitive at the low workability end of the scale than at high workability.  However, very dry mixes tend to stick in one or both hoppers and the material has to be  eased gently by poking with a steel rod. Moreover, it seems that for concrete of ver
y low  workability the actual amount of work required for full compaction depends on the  richness of the mix while the compacting factor does not: leaner mixes need more work  than richer ones. This means that the implied assumption that all mixes with the same  compacting factor require the same amount of useful work is not always justified.  Li
kewise, the assumption, mentioned earlier, that the wasted work represents a constant  proportion of the total work done regardless of the properties of the mix is not quite  correct. Nevertheless, the compacting factor test undoubtedly provides a good measure of  workability.   An automatic compacting factor test apparatus has also been developed
.  Here, the cylinder is supported by a spring balance which can be calibrated for a given  mix so as to read workability directly, or even to indicate the excess or deficiency of  water in kilograms per batch.    Flow Test    This laboratory test gives an indication of the consistence of concrete and its  proneness to segregation by measuring the
 spread of a pile of concrete subjected to  jolting. It is with regard to segregation that the flow test is of greatest value, but  it also gives a good assessment of consistence of stiff, rich, and rather cohesive mixes.   The test was covered by ASTM Standard C 124-39 (reapproved in 1966, but withdrawn  in 1974 because it is little used rather t
han because the test is not appropriate). The  apparatus consists essentially of a brass-top table, 762 mm (30 in.) in diameter,  mounted so that it can be jolted by a drop of 13 mm (1/2 in.). A mould in the shape of a  frustum of a cone, much more squat than the slump cone, is placed at the centre of the  table, filled with concrete in two layers
 and compacted in a manner similar to the slump  test. The mould is now removed and the table is jolted 15 times in 15 seconds. This is  done by a wheel operating an actuating cam. As a result, the concrete spreads over the  table, and the average diameter of the spread concrete is measured. The flow of concrete  is defined as the percentage incre
ase in the average diameter of the spread concrete (D in.)  over the original diameter of the base (10 in.) i.e.    (formula).   Values from 0 to 150 per cent can be obtained.   The jolting applied during the test encourages segregation, and if the mix is not cohesive the larger  particles of aggregate will separate out and move toward the edge of
 the table. Another form of  segregation is possible: in a sloppy mix the cement paste tends to run away from the centre of the table  leaving the coarser material behind. It should be noted that the flow test does not measure  workability, as concretes having the same flow may differ considerably in their workability.    Remoulding Test    Use is
 made of the flow table in another test, in which an assessment of workability is  made on the basis of the effort involved in changing the shape of a sample of concrete.  This is the remoulding test, developed by Powers. The apparatus is shown  diagrammatically in Fig. 4.5. A standard slump cone is placed in a cylinder 305 mm (12  in.) in diamete
r	and 203 mm (8 in.) high, the cylinder being mounted rigidly on a flow  table, adjusted to give a 6.3 mm (1/4 in.) drop. Inside the main cylinder there is an inner ring,  210 mm (8 1/4 in.) in diameter and 127 mm (5 in.) high. The distance between the bottom of  the inner ring and the bottom of the main cylinder can be set between 67 mm and 76 mm
	(2 5/8 in. and 3 in.).    Fig. 4.5. Remoulding test apparatus     The slump cone is filled in the standard manner, removed, and a disc-shaped rider  (weighing 1.9 kg (4.3 lb)) is placed on top of the concrete. The table is now jolted at the  rate of one jolt per second until the bottom of the rider is 81 mm (3 3/16 in.) above the base plate.  At
 this stage, the shape of the concrete has changed from a frustum of a cone to a cylinder.  The effort required to achieve this remoulding is expressed as the number of jolts required.  For very dry mixes a considerable effort may be necessary.   The test is purely a laboratory one but is valuable as the remoulding effort appears to be  closely re
lated to workability.     Vebe Test    This is a development of the remoulding test in which the inner ring of Powers'  apparatus is omitted and compaction is achieved by vibration instead of jolting. The  apparatus is shown diagrammatically in Fig. 4.6.    Fig. 4.6. Vebe apparatus    The name Vebe is derived from the initials of V. Bahrner of Swe
den who developed the  test. The test is covered by BS 1881 : Part 2 : 1970, where it is referred to as the "V-B"  consistometer test; it is referred to also in ACI standard 211.3-75, under the name Vebe  test.   The remoulding is assumed to be complete when the glass plate rider is  completely covered with concrete and all cavities in the surface
 of the concrete have  disappeared. This is judged visually, and the difficulty of establishing the end point  of the test may be a source of error. To overcome it an automatically operated device  for recording the movement of the plate against time may be fitted.   Compaction is achieved using a vibrating table with an eccentric weight rotating 
at  3000 rev / min and a maximum acceleration of 3 g to 4 g. It is assumed that the input of  energy required for compaction is a measure of workability of the mix, and this is  expressed in Vebe seconds, i.e. the time required for the remoulding to be complete.  Sometimes, a correction for the change in the volume of concrete from V2 before to V1
	after vibration is applied, the time being multiplied by V2/V1.   Vebe is a good laboratory test, particularly for very dry mixes. This is in contrast to  the compacting factor test where error may be introduced by the tendency of some dry  mixes to stick in the hoppers. The Vebe test also has the additional advantage that the  treatment of conc
rete during the test is comparatively closely related to the method of  placing in practice.    German Flow Table    This test has recently become more widespread in its use and is therefore of interest.  The apparatus consists essentially of a wooden board covered by a steel plate with a  total mass of 16 kg. This board is hinged along one side t
o	a base board, each board  being 700 mm square. The upper board can be lifted up to a stop so that the free edge  rises 40 mm. Appropriate markings indicate the location of the concrete to be deposited  on the table.   The table top is moistened and a frustum of a cone of concrete, lightly tamped by a  wooden tamper in a prescribed manner, is pla
ced using a mould 200 mm high with a  bottom diameter of 200 mm and a top diameter of 130 mm. Excess concrete is removed,  the surrounding table top is cleaned, and after an interval of 15 seconds the table top is  lifted 15 times in 15 seconds, this motion avoiding a significant force against the stop. In  consequence, the concrete spreads and th
e	maximum spread parallel to the two edges of  the table is measured. The average of these two values, given to the nearest millimetre,  represents the flow. A value of 400 indicates a medium workability, and 500  high workability.   Concrete should at this stage appear uniform and cohesive or else the test is  considered inappropriate for the giv
en mix. Thus the test offers an indication of the  cohesiveness of the mix. The test is applicable to flowing concrete (see page 110). For a  given mix, the effect of varying quantities of superplasticizer is such that the  relation between the flow and slump is linear.   Full details of the test are given in the German standard DIN 1048: Part 1. 
In the  same standard, there is prescribed also a compaction test, in which a mould 200 mm  square and 400 mm high is loosely filled with concrete, which is then fully compacted. The  ratio of the initial height to the final height is a measure of consistence; this is related to  the reciprocal of the compacting factor.    Ball Penetration Test   
 This is a simple field test consisting of the determination of the depth to which a 152  mm (6 in.) diameter metal hemisphere, weighing 136 kg (30 lb), will sink under its own  weight into fresh concrete. A sketch of the apparatus, devised by J. W. Kelly and known  as the Kelly ball, is shown in Fig. 4.7.    Fig. 4.7. Kelly ball     The use of th
is test is similar to that of the slump test, that is routine checking of  consistence for control purposes. The test is essentially an American one, covered by  ASTM Standard C 360-63 (reapproved in 1975), and is rarely used in Britain. It is,  however, worth considering the Kelly ball test as an alternative to the slump test, over  which it has 
some advantages. In particular, the ball test is simpler and quicker to  perform and, what is more important, it can be applied to concrete in a wheelbarrow or  actually in the form. In order to avoid the effects of a boundary the depth of the  concrete being tested should be not less than 200 mm (8 in.), and the least lateral  dimension 460 mm (1
8	in.).   As would be expected, there is no simple correlation between penetration and slump,  since neither test measures any basic property of concrete but only the response to  specific conditions. On a site, when a particular mix is used, correlation can be found,  as shown for instance in Fig. 4.8. In practice, the ball test is essentially us
ed to  measure variations in the mix, such as those due to a variation in the moisture content  of the aggregate.    Fig. 4.8. Relation between Kelly ball penetration and slump    Nasser's K-probe    Among the various recent attempts to devise a simple workability test, the probe test of  Nasser deserves mention. Here, a probe is inserted vertical
ly to a certain depth into  fresh concrete in the mould, either before or after compaction, and, following  withdrawal of the probe after one minute, the residual height of the mortar in the tube  is measured. The external diameter of the probe is 19 mm (3/4 in.) and it contains  openings through which mortar enters the tube. Nasser and Rezk claim
 that the test  gives a measure of workability of the concrete because the probe reading is affected by  cohesive, adhesive, and friction forces within the mix. Thus, an over-wet mix, which  exhibits a high slump, would lead to a relatively low level of mortar retained in the probe,  this being the result of segregation. Nevertheless, the probe re
ading appears to be related  to slump, providing this does not exceed 80 mm (or 3 in.).    Two-point Test    Tattersall has repeatedly criticized all the existing workability tests on the  grounds that they measure only one parameter. His argument is that the flow of fresh  concrete can be described by the Bingham model, i.e. by the equation    (f
ormula),    where  t  = shear stress at rate of shear y,    t0 = yield stress,  and  u  = plastic viscosity.    In consequence, two "points" are required to characterize workability, and Tattersall  developed techniques of torque measurement using a modified food mixer. Hence, he  deduced experimentally data related to the shear stress at a given 
rate of shear and to  constants representing the yield stress, t0, and plastic viscosity, u, of the mix. It is the  latter two that, in his view, provide a measure of the fundamental rheological properties of  concrete. Their determination requires the measurement of torque to rotate the mixer at  two speeds.   Apparatus actually exists for concre
tes with a slump of at least 100 mm but the test  has not yet found widespread acceptance. As mentioned earlier, Tattersall's  approach is based on the assumption that the Bingham model describes the flow of  concrete but the actual situation is more complex. Indeed it will probably be some time  before practical alternatives to the existing, rela
tively simple, methods have been developed.  These methods, by and large, give a helpful indication of  "workability" and are certainly  adequately sensitive to variations in the water content of  the mix.    Comparison of Tests    It should be said at the outset that no comparison is really possible as each test  measures the behaviour of concret
e under different conditions. The  particular uses of each test have been mentioned.   The compacting factor or test is closely related to the reciprocal of workability. The  remoulding and Vebe tests are direct functions of workability. The Vebe test measures the  properties of concrete under vibration as compared with the free-fall conditions of
 the  compacting factor test and the jolting in the remoulding test. All three tests are  satisfactory in the laboratory, but the compacting factor apparatus is the  most suitable for site use.   An indication of the relation between the compacting factor and the Vebe time is given  by Fig. 4.9, but this applies only to the mixes used, and the rel
ation must not be  assumed to be generally applicable since it depends on factors  such as the shape and  texture of the aggregate or presence of entrained air, as well as on mix proportions. For  specific mixes, the relation between compacting factor and slump has been obtained, but  such a relation is also a function of the properties of the mix
. The relation between the  number of jolts in Powers' remoulding test and slump (Fig. 4.10) is also only broadly  defined. A general indication of the pattern of the relation between the compacting factor,  Vebe time and slump is shown in Fig. 4.11. The influence of the richness of the mix in two  of these relations is clear. The absence of influ
ence in the case of the relation between slump  and Vebe time is illusory because slump is insensitive at one end of the scale (low  workability) and Vebe time at the other; thus two asymptotic lines with a small  connecting part are present.    Fig. 4.9. Relation between compacting factor and Vebe time    Fig. 4.10 Relation between the number of 
jolts using Powers' remoulding test apparatus  and slump for mixes with fine aggregates of different fineness    Fig. 4.11 General pattern of relations between workability tests for mixes of varying  aggregate / cement ratios     The flow test is valuable in assessing the cohesiveness of a laboratory mix.   The slump and penetration tests are pure
ly comparative, and in that capacity both are  very useful except that the slump test is unreliable with lean mixes, for which good  control is often of considerable importance. As already stated, the ideal test for  workability has yet to be devised. For this reason, it is worth stressing the value of visual  inspection of workability and of asse
ssing it by patting with a trowel in order to see the  ease of finishing. Experience is clearly necessary but, once it has been acquired, the "by  eye" test, particularly for the purpose of checking uniformity, is both rapid and reliable.      Effect of Time and Temperature on Workability    Freshly mixed concrete stiffens with time. This should n
ot be confused with setting of  cement. It is simply that some water from the mix is absorbed by the aggregate, some is  lost by evaporation, particularly if the concrete is exposed to sun or wind, and some is  removed by the initial chemical reactions. The compacting factor decreases by up to  about 0.1 during a period of one hour from mixing; th
e exact value of the loss in  workability varies with the richness of the mix, the type of cement, the temperature of  the concrete, and the initial workability. An example of a slump - time curve, obtained by  Evans, is given in Fig. 4.12. The change in workability with time depends also on the  moisture condition of aggregate (at a given total w
ater content): the loss is greater with  dry aggregate owing to the absorption of water by aggregate, as of course would be  expected.    Fig. 4.12. Relation between slump and time since completion of mixing for a 1:2:4  concrete with a water / cement ratio of 0.775     Because of this change in apparent consistence and also because we are really 
 interested in the workability at the time of placing, i.e. some time after mixing, it is  preferable to delay the appropriate test until, say, 15 minutes after mixing.   The workability of a mix is also affected by the ambient temperature, although, strictly  speaking, we are concerned with the temperature of the concrete itself. Fig. 4.13 gives 
an  example of the effect of temperature on slump of laboratory-mixed concrete: it is apparent  that on a hot day the water content of the mix would have to be increased for a constant  workability to be maintained. Fig. 4.14 shows that as the concrete temperature increases the  percentage increase in water required to effect a 25 mm (1 in.) chang
e in slump also increases.    Fig. 4.13. Influence of temperature on slump of concretes with different maximum  aggregate size    Fig. 4.14. Influence of temperature on amount of water required to change slump     It is interesting that the loss of workability with increase in temperature has not  been confirmed by site tests made by Shalon in a h
ot climate. Up to a temperature of  40 C (104 F) and with a relative humidity within the range 20 to 70 per cent, no  effect of temperature on slump has been observed. Only above 50 C (122 F) or with  humidities lower than 20 per cent does the slump fall off rapidly.   Moreover, it seems that the loss of slump in hot and dry air is greater than th
e  decrease in ease of placing. There is therefore no corresponding large increase in the  water requirement. These findings apply up to 40 C (104 F) and within 20 minutes of  mixing. Over longer periods, there is an unmistakable loss of slump so that, for  instance, with a long haul of ready-mixed concrete, high temperature would increase the  wa
ter requirement for a given workability.   The discrepancy between site tests and the laboratory tests of Fig. 4.13 may possibly  be explained by the use of air-entrained concrete in the latter case: there may be some  form of interaction between the entrained air, the cement, and the climate. The full  answer is not known and it is recommended th
at for any new conditions actual site tests  be made.      Segregation    In discussing workable concrete in general terms it was implied that such concrete  should not easily segregate i.e. it ought to be cohesive. However, strictly speaking the  absence of a tendency to segregate is not included in the definition of a workable mix.  Nevertheless
, the absence of appreciable segregation is essential as full compaction of a  segregated mix is impossible.   Segregation can be defined as separation of the constituents of a heterogeneous  mixture so that their distribution is no longer uniform. In the case of concrete, it is  the differences in the size of particles and in the specific gravity
 of the mix  constituents that are the primary causes of segregation, but its extent can be  controlled by the choice of suitable grading and by care in handling.   There are two forms of segregation. In the first, the coarser particles tend to  separate out since they tend to travel further along a slope or to settle more than  finer particles. T
he second form of segregation, occurring particularly in wet mixes, is  manifested by the separation of grout (cement plus water) from the mix. With some  gradings when a lean mix is used, the first type of segregation may occur if the mix is  too dry; addition of water would improve the cohesion of the mix, but when the mix  becomes too wet the s
econd type of segregation would take place.   The influence of grading on segregation was discussed in detail in Chapter 3, but the actual  extent of segregation depends on the method of handling and placing of concrete. If the concrete  does not have far to travel and is transferred directly from the wheelbarrow to the final  position in the form
, the danger of segregation is small. On the other hand, dropping  concrete from a considerable height, passing along a chute, particularly with changes of  direction, and discharging against an obstacle - all these encourage segregation so that  under such circumstances a particularly cohesive mix should be used. With a correct  method of handlin
g, transporting and placing, the likelihood of segregation can be greatly  reduced: there are many practical rules but these are outside the scope of this book.   It should be stressed, however, that concrete should always be placed direct in the  position in which it is to remain and must not be allowed to flow or be worked along the  form. This 
prohibition includes the use of a vibrator to spread a heap of concrete over  a larger area. Vibration provides a most valuable means of compacting concrete, but,  because a large amount of work is being done on the concrete, the danger of segregation  (in placing as distinct from handling) due to an improper use of a vibrator is  increased. This 
is particularly so when vibration is allowed to continue too long: with  many mixes, separation of coarse aggregate toward the bottom of the form and of the  cement paste toward the top may result. Such concrete would obviously be weak, and the  laitance (scum) on its surface would be too rich and too wet so that a crazed surface  with a tendency 
to dusting might result.   It may be noted that entrained air reduces the danger of segregation. On the other  hand, the use of coarse aggregate whose specific gravity differs appreciably from that  of the fine aggregate would lead to increased segregation.   Segregation is difficult to measure quantitatively, but is easily detected when  concrete
 is handled on a site in any of the ways listed earlier as undesirable. A good  picture of cohesion of the mix is obtained by the flow test. As far as proneness to  segregation on over-vibration is concerned, a good test is to vibrate a concrete cube  for about 10 min and then to strip it and observe the distribution of coarse aggregate:  any segr
egation will be easily seen.    Bleeding    Bleeding, known also as water gain, is a form of segregation in which some of the water  in the mix tends to rise to the surface of freshly placed concrete. This is caused by  the inability of the solid constituents of the mix to hold all of the mixing water when  they settle downwards. We are thus deali
ng with subsidence, and Powers treats  bleeding as a special case of sedimentation. Bleeding can be expressed quantitatively as  the total settlement per unit height of concrete. The bleeding capacity as well as the  rate of bleeding can be determined experimentally using the test of ASTM Standard C  232-71 (reapproved 1977).   As a result of blee
ding the top of every lift may become too wet and if the water is  trapped by superimposed concrete, porous, weak, and non-durable concrete will result. If  the bleeding water is remixed during finishing of the top surface a weak wearing surface  will be formed. This can be avoided by delaying the finishing operations until the  bleeding water has
 evaporated, and also by the use of wood floats and avoidance of  overworking the surface. On the other hand, if evaporation of water from the surface of  the concrete is faster than the bleeding rate plastic shrinkage cracking may result (see  p. 371).   Some of the rising water becomes trapped on the underside of coarse aggregate particles  or o
f reinforcement, thus creating zones of poor bond. This water leaves behind  capillaries, and since all the voids are oriented in the same direction, the  permeability of the concrete in a horizontal plane may be increased. A small amount of  voids of this type is nearly always present, but appreciable bleeding must be avoided as  the danger of fr
ost damage may be increased. Bleeding is often pronounced in thin slabs,  such as road slabs, and it is in these that frost generally constitutes a considerable  danger.   Bleeding need not necessarily be harmful. If it is undisturbed (and the water  evaporates) the effective water / cement ratio may be lowered with a resulting increase in  streng
th. On the other hand, if the rising water carries with it a considerable amount  of the finer cement particles a layer of laitance will be formed. If this is at the top  of a slab a porous surface will result, with a permanently "dusty" surface. At the top  of a lift a plane of weakness would form and the bond with the next lift would be  inadequ
ate. For this reason, laitance should always be removed by brushing and washing.   The tendency to bleeding depends largely on the properties of cement. Bleeding is  decreased by increasing the fineness of cement and is also affected by certain chemical  factors: there is less bleeding when the cement has a high alkali content, a high C3A  content
, or when calcium chloride is added. A higher temperature, within the normal  range, increases the rate of bleeding, but the total bleeding capacity is probably  unaffected. The physical properties of fine aggregate, especially that smaller than a  150 um (No. 100) sieve, may also affect bleeding. Rich mixes are less prone to  bleeding than lean o
nes. Reduction in bleeding is obtained by the addition of pozzolanas  or of aluminium powder. Air entrainment effectively reduces bleeding so that finishing  can follow casting without delay.   Bleeding of concrete continues until the cement paste has stiffened sufficiently to put  an end to the process of sedimentation.      The Mixing of Concret
e	   Mixing concrete by hand is expensive in labour and it is, therefore, not surprising that  mechanical mixers have been in general use for a great many years.    Concrete Mixers    The object of mixing is to coat the surface of all aggregate particles with cement  paste, and to blend all the ingredients of concrete into a uniform mass; this uni
formity  must furthermore not be disturbed by the process of discharging from the mixer. In fact,  the method of discharging is one of the bases of classification of concrete mixers.  Several types exist. In the tilting mixer, the mixing chamber, known as the drum, is  tilted for discharging. In the non-tilting type the axis of the mixer is always
	horizontal, and discharge is obtained either by inserting a chute into the drum or by  reversing the direction of rotation of the drum (when the mixer is known as a reversing  drum mixer), or rarely by splitting of the drum. There are also pan-type mixers, rather  similar in operation to an electric cake-mixer; these are called forced action mix
ers,  as distinct from the tilting and non-tilting mixers which rely on the free fall of  concrete in the drum.   Tilting mixers usually have a conical or bowl-shaped drum with vanes inside. The  efficiency of the mixing operation depends on the details of design, but the discharge  action is always good as all the concrete can be tipped out rapid
ly and in an  unsegregated mass as soon as the drum is tilted. for this reason, tilting-drum mixers  are preferable for mixes of low workability and for those containing large-size  aggregate.   On the other hand, because of a rather slow rate of discharge from a non-tilting drum  mixer, concrete is sometimes susceptible to segregation. In particu
lar, the largest size  of aggregate may tend to stay in the mixer so that the discharge starts as mortar and  ends as a collection of coated stones. It is not suggested, of course, that this is a  universal fault of non-tilting drum mixers but it may be advisable to check the  performance of an untried type. This can be done by comparing the conte
nt of coarse  aggregate in the second tenth of the batch with that in the ninth tenth. Limits for the  difference between the two contents are discussed on p. 229.   Non-tilting mixers are always charged by means of a loading skip, which is also used  with the larger tilting drum mixers. It is important that the whole charge from the skip  be tran
sferred into the mixer every time, i.e. no sticking must occur. Sometimes, a  shaker mounted on the skip assists in emptying it.   The pan mixer is generally not mobile and is therefore used either at a central mixing  plant on a large concrete project, at a precast factory, or in a small version in the  concrete laboratory. The mixer consists ess
entially of a circular pan rotating about its axis,  with one or two stars of paddles  rotating about a vertical axis not coincident with the axis  of the pan.  Sometimes, the pan is static and the axis of the star travels along a circular  path about the axis of the pan. In either case, the relative movement between the paddles  and the concrete 
is the same, and concrete in every  part of the pan is thoroughly mixed.  Scraper blades prevent mortar sticking  to the sides of the pan, and the height of the  paddles can be adjusted so as  to prevent a permanent coating of mortar forming on the  bottom of the  pan.   Pan mixers are particularly efficient with stiff and cohesive mixes and are, 
 therefore, often used in the manufacture of precast concrete. They are also suitable,  because of the scraping arrangements, for mixing very small quantities of concrete - hence  their use in the laboratory. A bowl-and-stirrer of the cake-mixer type, working on  the same principle as the pan mixer, is sometimes used for mixing of mortar.   It may
 be relevant to mention that in drum-type mixers no scraping of the sides takes  place during mixing so that a certain amount of mortar adheres to the sides of the drum  and stays there until the mixer has been cleaned at the end of the day's work. It  follows that at the beginning of concreting the first mix would leave a large  proportion of its
 mortar behind, and the discharge would consist largely of coated  coarse particles. This initial batch should be discarded. As an alternative, a certain  amount of mortar may be introduced into the mixer prior to the commencement of  concreting, a procedure known as buttering the mixer. A convenient and  simple way is to  charge the mixer with th
e	usual quantities of cement, water and fine aggregate, simply  omitting the coarse material. The mix in excess of that stuck in the mixer can be used  in construction and may in fact be particularly suitable for placing at a cold joint.  The necessity of  buttering should not be forgotten in laboratory work.   The size of a mixer is described by 
the volume of concrete after compaction (BS 1305  : 1974). Previously, the volume of the unmixed ingredients in a loose state was also  included in the description; the latter  may be up to 50 per cent greater than the  compacted volume. Mixers are made in a variety of sizes from 0.04 m3 (1 1/2 ft3) for  laboratory use up to 13 m3 (17 yd3). If the
 quantity mixed represents only a small  fraction of the capacity of the mixer the resulting mix may not be uniform, and the  operation would, of course, be uneconomical. Overload not exceeding 10 per cent is  generally harmless.   One other type of mixer is of interest: the dual drum mixer used in road construction.  In this, there are two drums 
in series, concrete being mixed part of the time in one and  then transferred to the other for the remainder of the mixing time, and finally  discharged. In the meantime, the first drum is recharged and initial mixing takes place.  The operations are synchronized so that there is no intermixing of batches. In this  manner, the yield of concrete ca
n	be doubled compared with an ordinary mixer with the  same batching equipment, and in road construction, where space and access are often limited,  this is a considerable advantage. Triple-drum mixers are also used.   All the mixers considered so far are batch mixers, since one batch of concrete is mixed and  discharged before any more materials 
are added. As opposed to this, a continuous mixer discharges  mixed concrete steadily and is fed by a continuous weigh batching system. The mixer itself may  be of a drum type or may rely on a screw moving in a stationary housing.   Specialized mixers are used in shotcreting and for mortar for preplaced aggregate concrete. In  the "colloid" mixer 
used for the latter, cement and water are formed into colloidal grout by passage,  at a speed of 2000 rev / min, through a narrow gap, and sand is subsequently added to the grout.  This type of mixer may become of interest in concrete-making because the pre-mixing of cement  and water allows better hydration and may lead to a higher strength at a 
given water / cement  ratio than conventional mixing. For instance, at water / cement ratios of 0.45 to 0.50, a gain in  strength of 10 per cent has been observed. However, two-stage mixing undoubtedly represents  a higher cost and is likely to be justifiable only in special cases.    Uniformity of Mixing    In any mixer, it is essential that suff
icient interchange of materials between  different parts of the chamber takes place, so that uniform concrete is  produced. The efficiency of the mixer can be measured by the variability of  the mix discharged into a number of receptacles without interrupting the  flow of concrete. For instance, in Belgium the mix is discharged in eight  parts and
 these are compared for homogeneity. The classification of the  efficiency of mixing is then based on the data of Table 4.3. As is the case in  some European countries, the Table uses mean deviation (arithmetic mean  of the absolute values of deviation from the mean) and maximum deviation  (the absolute value of deviation from the mean) instead of
 the standard  deviation which is common in Britain and the U.S..    Table 4.3: Variability of Concrete in a "Satisfactory" Mixer     The values given in Table 4.3 would be typical of a satisfactory mixer, a  smaller deviation denoting a very good mixer, and a greater deviation a  poor one. It should be stressed, however, that the values listed gi
ve only  the order of variation as the performance of a mixer depends on the  consistence of the mix and on the maximum size of the aggregate used.   A rather rigid test of ASTM Standard C 94-78a (formally applicable  only to mixers for ready-mixed concrete) lays down that samples of  concrete should be taken from about 1/6 and 5/6 points of a bat
ch, and the  differences in the properties of the two samples should not exceed any of  the following:   density of concrete  16 kg / m3 (1 lb / ft3)  air content   1 per cent  slump    25 mm (1 in.) when the average is      under 100 mm (4 in.), and 40 mm      (1.5 in.) when the average is 100 to      150 mm (4 to 6 in.)  percentage of aggregate 
 retained on a 4.75 mm  (3/16 in.) sieve   6 per cent  density of air-free mortar  1.6 per cent  compressive strength (average  7-day strength of 3 cylinders) 7.5 per cent.    The U.S. Bureau of Reclamation lays down similar requirements.   In the United Kingdom, BS 3963: 1974 lays down a test of performance  of mixers using a specified concrete m
ix. Tests are made on two samples  from each quarter of a batch. Each sample is subjected to wet analysis  (see page 261) in accordance with British Standard BS 1881: Part 2: 1970  and the following are determined:    water content as percentage of solids to 0.1 per cent,  fine aggregate content as percentage of total aggregate to 0.5 per cent,  c
ement as percentage of total aggregate to 0.01 per cent,  water / cement ratio to 0.01.   The sampling accuracy is assured by a limit on the average range of  pairs. If two samples in a pair differ unduly, i.e. their range is an outlier,  that pair of results can be discarded.   The mixer performance is judged by the difference between the highest
	and the lowest average of pairs for each batch using three separate test  batches; thus one bad mixing operation does not condemn a mixer. The  maximum acceptable variabilities of the percentages listed earlier are  prescribed by British Standard BS 1305 : 1974 for different maximum  aggregate sizes and in BS 4251 : 1974 for truck-type mixers. T
ypical values  in the latter case are:    for percentage water content 0.8 to 0.9 per cent,  for percentage cement content 8 per cent of the nominal value,  for percentage of fine aggregate 8 to 9 per cent of the nominal value.    Mixing Time    On a site, there is often a tendency to mix concrete as rapidly as possible,  and it is, therefore, imp
ortant to know what is the minimum mixing time  necessary to produce a concrete uniform in composition and, as a result, of  satisfactory strength. This time varies with the type of mixer, and, strictly  speaking, it is not the mixing time but the number of revolutions of the  mixer that is the criterion of adequate mixing. Generally, about 20  re
volutions are sufficient. Since, however, there is an optimum speed of  rotation recommended by the manufacturer of the mixer, the number of  revolutions and the time of mixing are interdependent.   For a given mixer, there exists a relation between mixing time and  uniformity of the mix. Typical data are shown in Fig. 4.15, based on  Shalon's tes
ts, the variability being represented as the range of strengths  of specimens made from the given mix after a specified mixing time. Fig.  4.16 shows the results of the same tests plotted as a coefficient of variation  against mixing time. It is apparent that mixing for less than 1 to 1 1/4 min  produces an appreciably more variable concrete, but 
prolonging the mixing  time beyond these values results in no significant improvement in uniformity.    Fig. 4.15. Relation between compressive strength and mixing time    Fig. 4.16. Relation between coefficient of variation of strength and mixing time     The average strength of concrete also increases with an increase in  mixing time, as shown f
or instance by Abrams' tests (Fig. 4.17). The  rate of increase falls rapidly beyond about one minute and is not significant  beyond two minutes; sometimes, even a slight decrease in strength has  been observed. Within the first minute, however, the influence of  mixing time on strength is of considerable importance. For instance,  Shalon calculat
ed that, for a given required strength, increasing the  mixing time from 30 sec to 1 min permits a saving in the cement content of  as much as 30 kilograms per cubic metre (50 lb /yd3).    Fig. 4.17. Effect of mixing time on strength of concrete     As mentioned before, the exact value of the minimum mixing time varies  with the type of mixer and 
depends also on its size. Table 4.4 gives typical  values. The mixing time is reckoned from the time when all the solid  materials have been put in the mixer, and it is usual to specify that all the  water has to be added not later than after one quarter of the mixing time.  The figures quoted refer to the usual mixers but there are many modern  l
arge mixers which perform satisfactorily with a mixing of 1 to 1 1/2 min. In  high-speed pan mixers, the mixing time can be as short as 35 sec. On the  other hand, when lightweight aggregate is used, the mixing time should be  not less than 5 min, sometimes divided into 2 min of mixing the aggregate  with water, followed by 3 min with cement added
.	In general, the length of  mixing time required for sufficient uniformity of the mix depends on the  quality of blending of materials during charging of the mixer: simultaneous  feed is beneficial.    Table 4.4: Recommended Minimum Mixing Times     No general rules on the order of feeding the ingredients into the mixer  can be given as they depe
nd on the properties of the mix and of the mixer.  Generally, a small amount of water should be fed first, followed by all the  solid materials, preferably fed uniformly and simultaneously in to the  mixer. If possible, the greater part of the water should also be fed during  the same time, the remainder of the water being added after the solids. 
 With some drum mixers, however, when a very dry mix is used, it is  necessary to feed first some water and the coarse aggregate, as otherwise  its surface does not become sufficiently wetted. Moreover, if coarse  aggregate is totally absent to begin with, sand or sand and cement become  lodged in the head of the mixer and do not become incorporat
ed in the  mix; this is known as head pack. If water or cement are fed too fast or are  too hot there is a danger of formation of cement balls, sometimes up to  70 mm (or 3 in.) in diameter (see page 250). With small laboratory pan  mixers and very stiff mixes, it has found convenient to feed first  sand, a part of the coarse aggregate, and cement
, then the water, and  finally the remainder of the coarse aggregate so as to break up any nodules  of mortar.   Let us consider now the other extreme - mixing over a long period.  Generally, evaporation of water from the mix takes place, with a consequent  decrease in workability and increase in strength. A secondary effect  is that of grinding o
f the aggregate, particularly if soft: the grading of the  aggregate thus become finer, and the workability lower. The friction  effect also produces an increase in the temperature of the mix.   In the case of air-entrained concrete, prolonged mixing reduces the air  content by about 1/6 per hour (depending on the type of air-entraining  agent), w
hile a delay in placing without continuous mixing causes a drop in  air content by only about 1/10 per hour. On the other hand, a decrease in  mixing time below 2 or 3 minutes may lead to inadequate entrainment of  air.   Intermittent remixing up to about 3 hours, and in some cases up to 6  hours, is harmless as far as strength and durability are 
concerned, but the  workability falls off with time unless loss of moisture from the mixer is  prevented. Adding water to restore workability, known as re-tempering,  will lower the strength of the concrete.   Some investigators have reported that this loss of strength is  smaller than would be expected from the consideration of the total  water /
 cement ratio; others have found the loss to be the same as if the  re-tempering water were added during initial mixing. The explanation of  this discrepancy probably lies in how the water was lost: water lost by  evaporation should not be included in the effective water / cement ratio but  if re-tempering replaces the water used up in hydration t
hen the added  water forms part of the effective water which governs the strength (see Fig.  4.19).   Re-tempering has been reported slightly to increase the shrinkage but  this probably occurs only if the effective water / cement ratio is increase by  the added water.    Hand Mixing    There may be occasions when concrete has to be mixed by hand 
and,  because in this case uniformity is more difficult to achieve, particular care  and effort are necessary. The aggregate should be spread in a uniform layer  on a hard, clean and non-porous base; cement is then spread over the  aggregate, and the dry materials are mixed by turning over from one end of  the tray to another and "cutting" with a 
shovel until the mix appears  uniform. Turning three times is usually required. Water is then gradually  added so that neither water by itself nor with cement can escape. The mix  is turned over again, usually three times, until it appears uniform in colour  and consistence.   It is obvious that during hand mixing no soil or other extraneous mater
ial  must be allowed to become included in the concrete.      Ready-mixed Concrete    If instead of being batched and mixed on the site, concrete is delivered  ready for placing from a central plant it is referred to as ready-mixed or  pre-mixed concrete. This type of concrete is used extensively as it offers  numerous advantages in comparison wit
h the orthodox method of manufacture.  In many countries, including the United Kingdom, more than half  the concrete used in in situ construction is ready-mixed.   Ready-mixed concrete is particularly useful on congested sites or in road  construction where little space for the mixing plant and for extensive  aggregate stockpiles is available, but
 perhaps the greatest single advantage   of ready-mixed concrete is that it may be made under better conditions of  control than are normally possible on any but large construction sites.  Control has to be enforced but, since the central mixing plant operates  under near-factory conditions, a really close control of all operations of  manufacture
 of fresh concrete is possible. In a modern batching and mixing  plant, interlocking prevents incorrect batching quantities, and sometimes a  printed record of weights of ingredients of every batch is made. Proper  care during transportation of the concrete is also ensured by the use of  agitator trucks, but the placing and compaction remain, of c
ourse, the  responsibility of the personnel on the site. Ready-mixed concrete can be  considered to be more in the nature of a factory-mixed product, almost  comparable with steel, so that a great deal of uncertainty and variability  associated with the production of concrete on many a site is removed.   The used of ready-mixed concrete is also ad
vantageous when only small  quantities of concrete are required or when concrete is placed only at  intervals. Usually, the price of ready-mixed concrete is somewhat higher  than of site-mixed concrete, but this may often be offset by savings in the  cement content, site organization, and supervisory staff.   There are two principal categories of 
ready-mixed concrete. In the first,  the mixing is done at a central plant and the mixed concrete is then  transported, usually in an agitator truck which revolves slowly so as to  prevent segregation and undue stiffening of the mix. Such concrete is  known as central-mixed concrete. Here, the materials are batched at a  central plant but are mixe
d in a mixer truck either in transit to the site or  immediately prior to the concrete being discharged. Transit-mixing permits  a longer haul and is less vulnerable in case of delay, but the capacity of a  truck used as a mixer is only about three-quarters of the same truck used  solely to agitate pre-mixed concrete. Sometimes, the concrete is pa
rtially  mixed at a central plant in order to increase the capacity of the agitator  truck. The mixing is completed en route. Such concrete is known as  shrink-mixed concrete. Truck mixers usually have a capacity of 6 m3  (8 yd3) but 7.5 m3 (10 yd3) trucks also exist. (For a definition of mixer size  see page 227.)   It should be explained that ag
itating differs from mixing solely by the  speed of rotation of the mixer: the agitating speed is between 2 and  6 rev / min compared with the mixing speed of 4 to about 16 rev / min; there  is thus some overlap in the definitions.  BS 1926 : 1962 specifies a minimum  mixing speed of 7 rev / min. It may be noted that the speed of mixing affects  t
he rate of stiffening, while the total number of revolutions controls the  uniformity of mixing. Usually, 70 revolutions at mixing speed are required  as a minimum, and 100 revolutions at mixing speed is the maximum  allowed. An overriding limit of 300 revolutions in toto is laid down by  ASTM Standard C 94-78a.   The details of batching plants ar
e outside the scope of the present book  and it will suffice to say that they are generally highly automated. As many  as eight different ingredients can be weigh-batched into large capacity  mixers, the quantities being accurately dispensed and automatically recorded  for control purposes. For all this no more than one operator is  required.   Th
e main problem in the production of ready-mixed concrete is maintaining  the workability of the mix right up to the time of placing. Concrete  stiffens with time (Fig. 4.18) and handling ready-mixed concrete often  takes quite a long while. The stiffening may also be aggravated by  prolonged mixing and by a high temperature. In the case of transit
-mixing,  water need not be added till nearer the commencement of mixing, but the  time during which the cement and moist aggregate are allowed to remain in  contact should be limited to about 90 minutes, although an amendment to  BS 5328 : 1976 allows 2 hours. The 90-minute limit is probably not unduly  conservative.    Fig. 4.18. Loss in slump d
uring agitating at 4 rev / min     Many specifications impose the same limit on the time of haul of  central-mixed concrete, and it is usual also to limit the total number of  revolutions during both mixing and agitating to approximately 300.  However, agitating up to 6 hours need not adversely affect the strength of  concrete provided the mix rem
ains sufficiently workable for full compaction.  Unless, however, the initial workability is high, the stiffening caused  by prolonged agitation would result in a concrete of very low workability,  especially in hot weather, when a high loss of water by evaporation takes  place in addition to the loss of free water by hydration of cement. For this
	reason, concrete is sometimes re-tempered (see page 233) by the addition  of water immediately before discharge; the workability is thus restored but  it must be realized that the resultant compressive strength will be affected  by the amount of water added to the mix (see Fig. 4.19).    Fig. 4.19. Effect of re-tempering water on the strength of
 concrete     There is no doubt that an increase in the use of ready-mixed concrete  will continue, but in the United Kingdom the situation has not yet settled  down to as well-defined a system as, for instance, in Sweden. There, plants  are classified, and a higher classification of production control allows the  use of a lower cement content tha
n	generally required and permits less  frequent sampling on site. Thus there is strong encouragement to use  ready-mixed concrete from a plant carefully monitored by an independent  national body (see p. 665). In Britain, on the other hand, information  about ready-mixed plants sometimes has to be accepted, "taking into  account the personal integ
rity of the person giving it and that of his  company".      Pumped Concrete    Since this book deals primarily with the properties of concrete, the details  of the means of transporting and placing are not considered. An exception  should, however, be made in the case of pumping of concrete since this  means of transportation requires the use of 
a	mix having special properties.   The system consists essentially of a hopper into which concrete is  discharged from the mixer, a concrete pump of the type shown in Fig. 4.20  or 4.21, and pipes through which the concrete is pumped.    Fig. 4.20. Direct-acting concrete pump    Fig. 4.21. Squeeze-type concrete pump     Many pumps are of the direc
t-acting, horizontal piston type with semi-rotary  valves set so as to permit always the passage of the largest particles  of aggregate being used: there is thus no full closure. Concrete is fed into  the pump by gravity and is also partially sucked in during the suction  stroke. The valves open and close with definite pauses so that concrete  mov
es in a series of impulses but the pipe always remains full.   There exist also small portable pumps, sometimes called squeeze pumps,  for use with small (up to 75 or 100 mm (3 or 4 in.)) pipes; Fig. 4.21 shows  such a pump. Concrete placed in a collecting hopper is fed by rotating  blades into a pliable pipe located in the pumping chamber. The va
cuum  inside the chamber is about 660 mm (26 in.) of mercury. This ensures that,  except when actually squeezed by a roller, the pipe has a normal  (cylindrical) shape so that a continuous flow of concrete is ensured. Two  rotating rollers progressively squeeze the tube and thus pump the concrete  in the suction pipe toward the delivery pipe. Sque
eze pumps are often  lorry-mounted and may deliver concrete through a folding boom.   Squeeze pumps move concrete for distances up to 90 m (300 ft) horizontally  or 30 m (100 ft) vertically. However, using piston pumps, concrete  can be moved up to about 450 m (1500 ft) horizontally or 40 m (140 ft)  vertically or to proportionate combinations of 
distance and lift. We should  note that the ratio of equivalent horizontal and vertical distances varies  with the consistence of the mix and with the velocity of the concrete in the  pipe: the greater the velocity the smaller the ratio: at 0.1 m / s it is 24 but  at 0.7 m / s only 4.5. Relay pumping is possible for greater distances. When  bends 
are used, and these must never be sharp, the loss of head should be  allowed for in the calculation of the range of delivery: roughly, each 10  bend is equivalent to 1 m of pipe.   Pumps of different sizes are available and likewise pipes of various  diameters are used but the pipe diameter must be at least three times the maximum aggregate size. 
Using squeeze pumps, an output of 20 m3  (25 yd3) of concrete per hour can be obtained with 75 mm (3 in.) pipes, but  piston pumps with 220 mm (9 in.) pipes can deliver up to 60 m3 (78 yd3)  per hour.   Pumping is economical only if it can be used over long uninterrupted  period as at the beginning of each period of pumping the pipes have to be  l
ubricated by mortar (at the rate of about 0.25 m3 per 100 m (1 yd3 per  1000 ft) of 150 mm (6 in.) diameter pipe) and also because at the end of  the operation a considerable effort is required in cleaning the pipes.  However, alterations to the pipeline system can be made very quickly as  special couplings are used. A short length of flexible hos
e	near the  discharge end facilitates placing but increases the friction loss. Aluminium  pipes must not be used because aluminium reacts with the alkalis in cement  and generates hydrogen. This gas introduces voids in the hardened  concrete with a consequent loss of strength, unless the concrete is placed in  a confined space.   The main advantag
es of pumping concrete are that it can be delivered to  points over a wide area otherwise not easily accessible, with the mixing  plant clear of the site; this is especially valuable on congested sites or in  special applications such as tunnel linings, etc. Pumping delivers the  concrete direct from the mixer to the form and so avoids double hand
ling.  Placing can proceed at the rate of the output of the mixer and is not held  back by the limitations of the transporting and placing equipment. A  significant proportion of ready-mixed concrete is nowadays pumped.   Furthermore, pumped concrete is unsegregated but of course in order to  be able to be pumped the mix must satisfy certain requi
rements. It might  be added that unsatisfactory concrete cannot be pumped so that any  pumped concrete is satisfactory as far as its properties in the fresh state are  concerned. Control of the mix is afforded by the force required to stir it in  the hopper and by the pressure required to pump it.   Concrete which is to be pumped must be well mixe
d	before feeding into  the pump and sometimes remixing in the hopper by means of a stirrer is  carried out. Broadly speaking, the mix must not be harsh or sticky, too dry  or too wet, i.e. its consistence is critical. A slump of between 40 and  100 mm (1 1/2 and 4 in.) or a compacting factor of approximately 0.90 to 0.95  or Vebe time of 3 to 5 se
c	is generally recommended, but pumping  produces a partial compaction so that at the point of delivery the slump  may be decreased by 10 mm to 25 mm (1/2 in. to 1 in.). With a lower water  content, the solid particles, instead of moving longitudinally in a coherent  mass in suspension, would exert pressure on the walls of the pipe. When  the wate
r	content is at the correct, or critical, value friction develops only  at the surface of the pipe and in a thin, 1 to 2.5 mm (0.04 to 0.1 in.), layer  of the lubricating mortar. Thus nearly all the concrete moves at the same  velocity, i.e. by way of plug flow. It is possible that the formation of the  lubricating film is aided by the fact that t
he dynamic action of the piston is  transmitted to the pipe, but such a film is also caused by steel trowelling of  a concrete surface. To allow for the film in the pipe a cement content  slightly higher than otherwise would be used is desirable. The magnitude of  the friction developed depends on the consistence of the mix, but there  must be no 
excess water as segregation would result.   It may be useful to consider the problems of friction and segregation in  more general terms. In a pipe through which a material is pumped, there is  a pressure gradient in the direction of flow due to two effects: head of the  material and friction. This is another way of saying that the material must  
be capable of transmitting a sufficient pressure to overcome all resistances  in the pipeline. Of all the components of concrete, it is only water that is  pumpable in its natural state, and it is the water, therefore, that transmits  the pressure to the other mix components.   Two types of blockage can occur. In one, water escapes through the mix
	so that pressure is not transmitted to the solids, which therefore do not  move. This occurs when the voids in the concrete are not small enough or  intricate enough to provide sufficient internal friction within the mix to  overcome the resistance of the pipeline. Therefore, an adequate amount of  closely packed fines is essential to create a "
blocked filter" effect, which  allows the water phase to transmit the pressure but not to escape from the  mix. In other words, the pressure needed to pump the concrete. It should be  remembered of course that more fines mean a higher surface area of the  solids and therefore a higher frictional resistance in the pipe.   We can see thus how the se
cond type of blockage can occur. If the fines  content is too high, the friction resistance of the mix can be so large that  the pressure exerted by the piston through the water phase is not sufficient  to move the mass of concrete, which becomes stuck. This type of failure is  more common in high strength mixes or in mixes containing a high  prop
ortion of very fine material such as crusher dust or fly ash, while the  segregation failure is more apt to occur in medium or low strength mixes  with irregular or gap grading .  The optimum situation therefore is to produce maximum frictional  resistance within the mix with minimum void sizes, and minimum frictional  resistance against the pipe 
walls with a low surface area of the aggregate.  This means that the coarse aggregate content should be high, but the  grading should be such that there is a low void content so that only little of  the very fine material is required to produce the "blocked filter" effect.   Experiments have shown that the size fractions which have the largest  in
fluence on the void content of practical mixes are: 2.36 mm to 5 mm  (No. 8 ASTM to 3/16 in.), 300 um to 600 um (No. 50 to No. 30 ASTM), and  150 um to 300 um (No. 100 to No. 50 ASTM). Of these, the first one is the  most significant as inadequate amount of material between 2.36 mm and  5 mm (No. 8 ASTM and 3/16 in.) results in a high void content
 of the  aggregate and hence leads to difficulties in pumping.   For concretes with maximum aggregate size of 20 mm (3/4 in.), the  optimum fine aggregate content lies between 35 and 40 per cent, and the  material finer than 300 um (No. 50 ASTM) should represent 15 to 20 per  cent of the weight of fine aggregate.   In mixes with low cement content
s, an adequate amount of material  passing the 150 um (No. 100 ASTM) sieve is necessary to achieve the fine  sieve effect. On the other hand, when the cement content of the mix is  high, a high content of fines is also necessary in order to increase the  surface area and to increase friction. Generally, the proportion of fine  aggregate which pass
es the 150 um (No. 100 ASTM) sieve should be about  3 per cent. This material may be the finer fraction of sand or a suitable  additive, such as tuff or tass. This fine material gives continuity in grading  right down to the cement fraction but still avoids a very high pipe  friction   Tests at the Building Research Establishment have shown that g
enerally the  volumetric cement content (at an assumed density of 1450 kg / m3) has to be at least equal  to the void content of the aggregate but very fine material other than cement can be  included with the latter. The pattern of the effect of the relation between the cement  content and void content on pumpability is shown in Fig. 4.22. Howeve
r, it is only fair  to add that theoretical calculations are not very helpful because the shape of the  aggregate particles influences their void content. Some experimental data are shown in  Fig. 4.23: they indicate that the upper limit of pumpability can be exceeded by very  rich concrete.    Fig. 4.22. Pumpability of concrete in relation to cem
ent content and void content of aggregate	Fig. 4.23. Limits on cement content for aggregates with various void contents with  respect to pumpability of concrete   It may be noted that a sudden rise in pressure caused by a restriction or by a  reduction in the diameter of the pipe may result in segregation of the aggregate which  is left behind 
as the cement paste moves past the obstacle.   The shape of the aggregate influences the suitability of a mix for pumping. Natural  sands are often particularly suitable for pumping because of their rounded shape and  also because the true grading is more continuous than with crushed aggregate where  within each size fraction there is less variety
 in size. For both these reasons, the void  content is low. On the other hand, using combinations of size fractions of crushed aggregate,  a suitable void content can be achieved. However, care is required as many crushed fines are  deficient in the size fraction 300 um to 600 um (No. 50 to No. 30 ASTM) but have  excess of material smaller than 15
0	um (No. 100). When using crushed coarse aggregate,  it should be remembered that crusher dust may be present and this should be taken into  account in considering the grading of the fine aggregate. Generally, with crushed coarse  aggregate, the fine aggregate content should be increased by about 2 per cent.   Recently, pumping of lightweight agg
regate concrete has been introduced. If the  aggregate surface is sealed, then it pumps as well as normal aggregate. But if the  aggregate surface is porous then the internal voids may not become fully saturated even  on thorough wetting. As a result, when pressure is applied in the pipeline, the air in  these voids contracts, and water is forced 
into the pores with the result that the mix  becomes too dry and too stiff. If pumping is stopped and pressure removed, water is  discharged from the aggregate; this water may carry with it the fine material so that a  plug forms on resumption of pumping. Some of the aggregate may also become crushed  by pumping. Nevertheless, with special admixtu
res, which have a thickening as well as  dispersing effect, lightweight aggregate concrete can be successfully pumped over  moderate distances. Pumping aids are considered by Valore.   Concrete with a high content of entrained air behaves in a similar manner. Under a high  pumping pressure, the air becomes compressed and no longer aids the mix by 
its  "ball-bearing" effect. The friction rises and so does the pressure: the air becomes compressed  further so that the workability drops even more. If the pipeline is long enough, the  reduction in volume of the air under pressure can absorb the entire movement of the  piston so that no concrete will come out at the delivery end. For this reason
,	air-entrained  concrete is usually pumped only over short distances: about 45 m (150 ft).      Vibration of Concrete    The process of compacting the concrete consists essentially of the elimination  of entrapped air. The oldest means of achieving this is by ramming or punning the  surface of the concrete in order to dislodge the air and force t
he particles into a  closer configuration. The more modern means is by vibration, by which the particles are  momentarily separated, thus allowing them to be drawn into a compact mass.   The use of vibration as a means of compaction makes it possible to use drier mixes than  can be compacted by hand (a compacting factor below about 0.75 or 0.80, d
own to 0.60  when pressure may also be required). In fact, extremely dry and stiff mixes can be vibrated  satisfactorily so that for a given desired strength concrete can be made with a lower cement  content. This means a saving in cost, but against that we have to offset the cost of the  vibrating equipment, and of heavier and more sturdy formwor
k. In any case,  the cost of labour would probably be the deciding factor if the choice is to  be made on the basis of cost alone. As far as the quality of the concrete is concerned,  both vibration and compaction by hand can, with the right mix and good workmanship, produce  excellent concrete. Likewise, both means of compaction can produce poor 
concrete: in the  case of hand-rammed concrete, inadequate compaction is the most common fault; when  vibration is used it is possible that this has not been applied uniformly to the entire  concrete mass so that some parts of it are not fully compacted while others are segregated  owing to over-vibration. However, with a sufficiently stiff and we
ll-graded mix the ill effects of  over-vibration can be largely eliminated.   It has been mentioned that the two basic means of compaction require mixes of different  workabilities: too dry a mix cannot be sufficiently worked by hand; and, conversely, too  wet a mix should not be vibrated as segregation may result. This point has to be watched as,
	for instance, some mixes suitable for pumping may have too wet a consistence for  vibration. Furthermore, different vibrators require different consistence of concrete  for most efficient compaction so that the consistence of the concrete and the  characteristics of the available vibrator have to be matched.    Internal Vibrators    Of the sever
al types of vibrators this is perhaps the most common one. It consists  essentially of a poker, housing an eccentric shaft driven through a flexible drive from  a motor. The poker is immersed in concrete and thus applies approximately harmonic  forces to it; hence, the alternative names of poker - or immersion - vibrator.   The frequency of vibrat
ion varies up to 12 000 cycles of vibration per minute: between  3500 and 5000 has been suggested as a desirable minimum with an acceleration of not less  than 4 g, but, more recently, vibration at 4000 to 7000 cycles has found favour.   The poker is easily moved from place to place, and is applied at 0.5 to 1 m (or 2 to 3  ft) centres for 5 to 30
 sec, depending on the consistence of the mix, but with some mixes  up to 2 min may be required. The actual completion of compaction can be judged by the  appearance of the surface of the concrete, which should be neither honeycombed nor  contain an excess of mortar. Gradual withdrawal of the poker at the rate of about 80  mm per sec (3 in. / sec)
 is recommended so that the hole left by the vibrator closes  fully without any air being trapped. The vibrator should be immersed through the  entire depth of the freshly deposited concrete and into the layer below if this is still  plastic or can be brought again to a plastic condition. In this manner a plane of  weakness at the junction of the 
two layers can be avoided and monolithic concrete is  obtained. With a lift greater than about 0.5 m (2 ft) the vibrator may not be fully  effective in expelling air from the lower part of the layer.   Internal vibrators are comparatively efficient since all the work is done directly  on the concrete, unlike other types of vibrators. Pokers are ma
de in sizes down to 20 mm  (3/4 in.) diameter so that they can be used even with heavily reinforced and relatively  inaccessible sections. An immersion vibrator will not expel air from the form boundary so  that "slicing" along the form by means of a flat plate on edge is necessary. The use of  absorptive linings to the form is helpful in this res
pect but expensive.    External Vibrators    This type of vibrator is rigidly clamped to the formwork resting on an  elastic support, so that both the form and the concrete are vibrated. As a  result, a considerable proportion of the work done is used in vibrating the formwork,  which also has to be strong and tight so as to prevent distortion and
 leakage of grout.   The principle of an external vibrator is the same as that of an internal one, but the  frequency is usually between 3000 and 6000 cycles of vibration  per minute, although  some vibrators reach 9000 cycles per minute. The  Bureau of Reclamation recommends  at least 8000 cycles. The power  output varies between 80 and 1100 W.  
 External vibrators are used for precast or thin in situ sections of such  shape or thickness that an internal vibrator cannot be used.   When an external vibrator is used, concrete has to be placed in layers of suitable depth as  air cannot be expelled through too great a thickness of concrete. The position of the  vibrator may have to be changed
 as concreting progresses.   Portable, non-clamped external vibrators may be used at sections not  otherwise accessible, but the range of compaction of this type of vibrator is very  limited. One such vibrator is an electric hammer, sometimes used for compaction of  concrete test specimens.    Vibrating Tables    This can be considered as a case o
f	formwork clamped to the vibrator instead of  the other way round, but the principle of vibrating the concrete and formwork together  is unaltered.   The source of vibration, too, is similar. Generally a rapidly-rotating eccentric  weight makes the table vibrate with a circular motion. With two shafts rotating in  opposite directions the horizont
al component of vibration can be neutralized so that  the table is subjected to a simple harmonic motion in the vertical direction only.  There exist also some small good quality vibrating tables operated by an electro-magnet  fed with alternating current. The range of frequencies used varies between 25 and about  120 Hz. An acceleration of about 
4	g to 7 g is desirable. About 1.5 g  and an  amplitude of 40 um (0.0015 in.) are believed to be the minima necessary for  compaction, but with these values a long period of vibration may be necessary. For  simple harmonic motion the amplitude, a, and the frequency, f, are related by the  equation    (formula).     When concrete sections of differ
ent sizes are to be vibrated, and in laboratory use,  a table with a variable amplitude should be used. Variable frequency of vibration is an  added advantage.   In practice, the frequency may rarely be varied during the actual compaction but, at  least theoretically, there are considerable advantages in increasing the frequency and  decreasing am
plitude as consolidation progresses. The reason for this lies in the fact  that initially the particles in the mix are far apart and the movement induced has to be  of corresponding magnitude. On the other hand, once partial compaction has taken place  the use of a higher frequency permits a greater number of adjusting movements in a given  time; 
a reduced amplitude means that the movement is not too large for the space  available. Vibration at too large an amplitude relative to the inter-particle space  results in the mix being in a constant state of flow so that full compaction is never  achieved. Bresson and Brusin found that there is an optimum amount of energy of  vibration for every 
mix, and various combinations of frequency and acceleration will be  satisfactory. However, a prediction of the optimum in terms of mix parameters is not  possible.   A vibrating table provides a reliable means of compaction of precast concrete and has the  advantage of offering uniform treatment.   A variant of the vibrating table is a shock tabl
e used in some precast works in Britain  and much more commonly in the Netherlands and in Denmark. The principle  of this process of compaction is rather different from the high frequency vibration  discussed earlier: in a shock table violent vertical shocks are  imparted at the rate of about 1 to 3 per second. The shocks are produced by a vertica
l  drop of up to 13 mm (1/2 in.), this being achieved by means of cams. Concrete is placed in  the form in shallow layers while the shock treatment progresses: extremely good results  have been reported but the process is rather specialized and, although 50 years old, not  widely used.    Other Vibrators    Various types have been developed for sp
ecial purposes but only a very brief mention of  these will be made.   A surface vibrator applies vibration through a flat plate direct to the top surface of  the concrete. In this manner the concrete is restrained in all directions so that the  tendency to segregate is limited; for this reason a more intense vibration can be used.   An electric h
ammer can be used as a surface vibrator when fitted with a bit having a  large flat area, say 100 mm by 100 mm (4 in. by 4 in.); one of the main applications is  compacting test cubes.   A vibrating roller is used for consolidating thin slabs. For road construction various  vibrating screeds and finishers are available, but the details of these ma
chines are  outside the scope of this book. A power float is used mainly for granolithic floors in  order to bind the granolithic layer to the main body of the concrete, and is more an aid  in finishing than a means of compaction.      Revibration    It is usual to vibrate concrete immediately after placing so that consolidation is  generally comp
leted before the concrete has stiffened. All the preceding sections refer  to this type of vibration.   It has been mentioned, however, that in order to ensure good bond between lifts the  upper part of the underlying lift should be revibrated, provided the lower lift can still  regain a plastic state; settlement cracks and the internal effects of
 bleeding can  thus be eliminated.  This successful application of revibration raises the question whether revibration can  be more generally used. On the basis of experimental results it appears that concrete  can be successfully revibrated up to about 4 hours from the time of mixing. Revibration at 1 to 2  hours after placing was found to result
 in an increase in the 28-day compressive  strength of the form shown in Fig. 4.24. The comparison is on the basis of the same  total period of vibration, applied either immediately after placing or in part then and  in part at a specified time later. An increase in strength of approximately 14 per cent has  been reported, but actual values would 
depend on the workability of the mix and on details  of the procedure: other workers have found increases of 3 to 9 per cent. In general the  improvement in strength is more pronounced at earlier ages, and is greatest in concretes liable  to high bleeding since the trapped water is expelled on revibration. For the same reason,  revibration greatly
 improves bond between concrete and reinforcement. It is possible  also that some of the improvement in strength is due to a relief of the plastic shrinkage  stresses around aggregate particles.    Fig. 4.24. Relation between 28-day compressive strength and the time of revibration     Despite these advantages revibration is not widely used as it i
nvolves an  additional step in the production of concrete, and hence increased cost; also, if applied  too late, revibration can damage the concrete.      Concreting in Hot Weather    There are some special problems involved in concreting in hot weather, arising both from  a higher temperature of the concrete, and, in many cases, from an increased
 rate of  evaporation from the fresh mix. These problems concern the mixing, placing and curing  of the concrete.   A higher temperature of fresh concrete results in a more rapid hydration and leads  therefore to accelerated setting and to a lower strength of hardened concrete since a less uniform  framework of gel is established (see p. 318). Fur
thermore, rapid evaporation may cause plastic  shrinkage and crazing, and subsequent cooling of the hardened concrete would introduce tensile  stresses. It is generally believed that plastic shrinkage is likely to occur when the rate of  evaporation exceeds the rate at which the bleeding water rises to the surface, but it  has been observed that c
racks also form under a layer of water and merely become  apparent on drying.   It appears useful to distinguish those cracks which are associated with evaporation  from those caused by differential settlement of fresh concrete due to some obstruction  to settlement, such as the coarse aggregate or the reinforcement. The latter can be avoided  by 
the use of a dry mix, good compaction, and not allowing too fast a rate of build-up of concrete.  Cracks caused by evaporation can be very deep, range in width between 0.1 mm and 3 mm  (0.004 and 0.12 in.), and can be quite short or as long as 1 m (or 3 ft). A drop in the ambient relative  humidity encourages this type of cracking so that in fact 
the causes of it appear to be rather complex.  According to the American Concrete Institute, the risk of plastic cracking is the  same at the following combinations of temperature and relative humidity:  41 C (106 F) and 90 per cent  35 C (90 F) and 70 per cent  24 C (75 F) and 30 per cent.   One practical conclusion is that thin sections, such as
 shell roofs, should  not be cast under hot and dry conditions   Plastic shrinkage appears to be related to some physical characteristics of  cement but the full extent of the problem is yet to be studied (see p. 371).  There are  some further complications in hot weather concreting: air-entraining  is more difficult, although this can be remedied
 by using larger  quantities of the entraining agent. A related problem is that, if relatively cool  concrete is allowed to expand when placed at a higher temperature,  the air voids expand and the strength is reduced. This would occur, for  instance, with horizontal panels but not with vertical ones in steel moulds where  expansion is prevented. 
Curing also presents additional problems  as the curing water tends to evaporate rapidly. The use of curing compounds is not  entirely satisfactory since it leads to lower compressive strengths than when continuous  water curing is applied; some experimental values are given in Table 4.5. What must be  remembered above all, however, is the overrid
ing importance of preventing evaporation  from the concrete.    Table 4.5: Effect of Curing on the 56-day Compressive Strength of Concrete Cast in the Desert     There are a number of remedial measures that can be taken. In the first instance, the  cement content should be kept as low as possible so that the heat of hydration  does not unduly aggr
avate the effects of high ambient  temperature. The temperature  of the fresh concrete can be lowered by pre-cooling one or more of the ingredients of the  mix. For instance, ice can used instead of some of the mixing water, but it is essential that the  ice melts completely before the mixing has been completed. The cooling of the  aggregate is mo
re difficult and, because of the low specific heat of stone, less  effective. All materials used should be protected from direct insolation.   The temperature of concrete delivered at site in hot weather should be as low as possible; an  upper limit of 29 C (85 F) is often specified.   The temperature T of the freshly mixed concrete  can be easily
 calculated from that of the ingredients, using the expression    (formula),    where T denotes temperature in  C or  F, W weight of ingredient per unit volume of  concrete, and the suffixes a, c, w refer to aggregate, cement, and water (both added and  in aggregate) respectively. The figure 0.2 is the approximate ratio of the specific heat  of th
e dry ingredients to that of water, and is applicable to both the SI and Imperial  systems of units.   The actual temperature of the concrete will be somewhat higher than indicated by the  above expression owing to the mechanical work done in mixing, and will further rise due  to the development of the heat of wetting and hydration of cement. To o
btain a better  picture we can say that if the water / cement ratio of a mix is 0.5 and the aggregate / cement  ratio is 5.6, then a drop of 1 C (or 1 F) in the temperature of fresh concrete can be obtained  by lowering the temperature either of the cement by 9 C (9 F) or of the water by 3.6 C (3.6 F)  or of the aggregate by 1.6 C (1.6 F). It can 
be seen that because of its relatively small quantity  in the mix the temperature of the cement is not important.   The use of hot cement per se is not detrimental to strength but it is preferable not to use  cement at temperatures above about 75 C (170 F). This statement is of interest since hot  cement is sometimes viewed with suspicion and vari
ous ill effects have at times been  ascribed to its use. However, if hot cement is dampened by a small amount of water before  it is well dispersed with other solids it may set quickly and form cement balls.   The influence of the temperature during setting on the strength at later ages is discussed  on p. 318, here it suffices to say that a tempe
rature not higher than about 16 C (60 F)  should be aimed at, and, if possible, 32 C (90 F) should not be exceeded. After placing,  concrete should be protected from the sun; otherwise, if a cold night follows, cracking is  likely to occur, the extent of cracking being directly related to the temperature difference.  In dry weather, wetting concre
te and allowing evaporation to take place results in  effective cooling; there is no cooling by this means when membrane curing is used so that  a higher temperature may be reached. Large exposed areas such as roads and airfields are  particularly vulnerable. Details of good practice for hot weather concreting have been  published by the American 
Concrete Institute.      Placement of Large Concrete Masses    There are two distinct situations possible. One is the placing of mass, unreinforced,  concrete, as for instance in a gravity dam. This is mass concrete, and its main feature,  indeed, used as a definition by the American Concrete Institute, is the need to deal  with the generation of 
heat and attendant volume changes in order to minimize cracking.  What has to be avoided is too large a differential in temperature between the interior of  the mass, where the heat of hydration of cement produces a large rise in temperature, and  the exterior, where the heat generated is lost to the surrounding atmosphere. Shrinkage  can also con
tribute to cracking. The role of creep under a temperature cycle is considered  on page 424.   To minimize the temperature rise several measures can be taken. First, the ingredients  of the fresh concrete can be cooled so as to reduce its temperature down to 7 C (45 F).  Cooling of the hardened concrete mass can be continued by circulating refrige
rated water  through embedded pipes. The mix ingredients and proportions have to be chosen very  carefully: low heat cement, low cement content, pozzolana replacement, rounded  aggregate, water-reducing admixtures - all help; Fig. 4.25 shows some of the influences on  the temperature rise.    Fig. 4.25. Temperature rise of mass concrete made with 
different cements     Mather has suggested that a Portland cement content as low as 36 kg / m3 (60 lb / yd3)  is possible, but the content of pozzolanas would be up to twice as high. Such a mix might  have a water content down to 48 kg / m3 (80 lb / yd3). The slump required is about 40  mm (1 1/2 in.) and the resulting 28-day cylinder strength wou
ld be in excess of 14 MPa  (2000 psi), which is generally adequate in a gravity dam. We can note that using a very  low cement content is not only economical per se but leads also to economy in other  measures used to overcome the undesirable effects of the heat of hydration of the  cement.   The placing programme is also a factor in heat generati
on as cooling before further  concrete has been placed reduces the temperature rise.   The fundamental problem is of course the presence of restraint to volume changes. This is  a topic outside the scope of this book but the American Concrete Institute has published  some helpful reports.   A somewhat different problem arises when we are dealing w
ith a large mass of reinforced  concrete. Here, many of the techniques used with mass concrete are inapplicable because a  medium or high-strength mix is required, early strength may be necessary, and embedding  pipework may not be permitted. The essential problem is, nevertheless, the same, i.e. the  interior of the mass will heat up more than th
e exterior, if a loss of heat at the surface is  large. If the difference in temperature between the interior and the exterior is large enough,  cracking will develop. This may occur either during the period of rise in temperature, in which  case the cracks will be in the interior (Fig. 4.26), or in the course of cooling, when cracking  would deve
lop on the surface (Fig. 4.27). The figures shown (Fig. 4.26 and Fig. 4.27)  imply that cracking will occur when the temperature difference exceeds 20 C (36 F).  There is also a possibility of cracking if concrete is placed against a cold surface; this may  occur even if the pour cannot be classified as large.    Fig. 4.26. An example of the patte
rn of temperature change which causes internal cracking  of a large concrete mass. The critical 20 C temperature difference occurs during heating,  but the cracks open only when the interior has cooled through a greater temperature range  than the exterior    Fig. 4.27. An example of the pattern of temperature change which causes external  crackin
g of a large concrete mass. The critical 20 C temperature difference occurs during  cooling   The solution to the problem is not by limiting the temperature rise in the interior but  rather by preventing the heat loss at the surface. Thus the entire concrete mass is allowed  to heat, more or less to the same degree, and expand without restraint; w
ith time, cooling,  again more or less uniform throughout, takes place, and the structure reaches its final  dimensions, again without restraint. To prevent a large heat loss the formwork and the top  surface of the structure must be adequately insulated.   According to FitzGibbon, who largely developed the large-pour technique for reinforced conc
rete,  the limit on temperature differential is about 20 C (36 F). Taking the coefficient of thermal expansion  of concrete as 10 x 10'-6 per C (5.5 x 10'-6 per F) (see Table 7.18), the differential strain is 200 x 10'-6.  This is a realistic estimate of tensile strain at cracking (see p. 282). FitzGibbon estimates  that the temperature rise under
 adiabatic conditions is 12 C per 100 kg of cement per  cubic metre of concrete (13 F per 100 lb / yd3), regardless of the type of cement used, for  cement contents between 300 and 600 kg / m3 (500 and 1000 lb / yd3). The reason for  disregarding the type of Portland cement is that, at high temperatures and later ages, the  total heat generated by
 low-heat cement is not necessarily lower than the total heat of  hydration of ordinary Portland cement. However, using Cemsave (see p. 74), the  temperature rise is only about one-half that for Portland cement.   In practice, the temperature at various points should be monitored by thermocouples, and  insulation should be adjusted accordingly. Th
e insulation must control loss of heat by  evaporation, conduction, and radiation. To achieve the first, a plastic membrane or a  curing compound should be used, but not spraying or pending as these have a cooling  effect. Plastic-coated quilts are useful in all respects but soft board can also be used. The  insulation must be maintained until the
 temperature differential has been reduced to 10 C  (18 F).   The technique is highly specialized and attention has to be paid to many factors, e.g. the  initial temperature of reinforcement, but very successful continuous pours of hundreds and  thousands of cubic metres of concrete have been made.      Preplaced Aggregate Concrete    This type of
 concrete is produced in two stages. In the first operation, coarse aggregate is  placed and compacted in the forms. The voids between the particles, forming some 30 to  35 per cent of the overall volume to be concreted, are filled with mortar in the second  stage. Preplaced aggregate concrete is also known as prepacked concrete, intrusion  concre
te or grouted concrete.   It is clear that the aggregate in the resulting concrete is of the gap-graded type. Typical  coarse and fine aggregate gradings are shown in Tables 4.6 and 4.7 respectively. Optimum  packing of the aggregate particles leads to great theoretical advantages but is not  necessarily achieved in practice.    Table 4.6: Typical
 Gradings of Coarse Aggregate for preplaced Aggregate Concrete	Table 4.7: Typical Grading of Fine Aggregate for Preplaced Aggregate Concrete     The coarse aggregate must be free from dirt and dust because, since these are not  removed in mixing, they would impair bond. The coarse aggregate has to be thoroughly  wetted or inundated before the m
ortar is intruded. However, water should not be allowed to  stand too long as algae can grow on the aggregate. The mortar is pumped under pressure  through slotted pipes, typically 35 mm (or 1 1/2 in.) in diameter and spaced at 2 m (7 ft) centres,  starting from the bottom of the mass, the pipes being gradually withdrawn. Pumping over  long distan
ces is possible.   A typical mortar consists (by weight) of two parts of Portland cement, one part of a very  finely divided and highly active pozzolana (for instance, fly ash), and three to four parts of  fine sand, with sufficient water to form a fluid mixture. The pozzolana reduces bleeding  and segregation and improves fluidity of the mortar. 
An intrusion aid (representing about 1  per cent of the weight of the cement plus the pozzolana) is added in order to improve the  fluidity of the mortar and to hold the solid constituents in suspension. The intrusion aid  also delays somewhat the stiffening of the mortar and contains a small amount of  aluminium powder, which causes a slight expa
nsion before setting takes place.   As an alternative, a mortar consisting of cement and fine sand can be mixed in a special  "colloid" mixer (see p. 228) which disperses the cement to such a degree that it remains in  suspension until the pumping has been completed. This type of preplaced aggregate  concrete is sometimes called colloidal concrete
.   The consistence of the mortar (about that of a thick cream) is expressed  as the time taken by a fixed quantity of mortar to discharge from a special cone; this is  known as the flow factor. Alternatively, a flowmeter is used; this simply shows how far a  given quantity of material discharged from a funnel will travel along a horizontal channe
l.   Preplaced aggregate concrete is economical in cement, as little as 120 to 150 kg per cubic  metre (200 to 250 lb / yd3) of concrete being used, but the strength of the resultant  concrete is limited by the high water / cement ratio necessary for a sufficient plasticity of  mortar; typical strength is 20 MPa (2900 psi). However, for the usual 
applications of  preplaced aggregate concrete, this strength is generally adequate and a concrete of more  uniform properties is obtained than is the case with conventional methods of placing, as  segregation is practically eliminated. As a result, a dense, impermeable and durable  concrete is produced. No internal vibration is used but external v
ibration at the level of the  grout surface may improve the exposed surfaces.   Preplaced aggregate concrete can be placed in locations not easily accessible by ordinary  concreting techniques; it can also be placed in sections containing a large number of  embedded items that have to be precisely located: this arises, for instance, in nuclear  sh
ields. Likewise, because the coarse and fine aggregate are placed separately, the danger  of segregation of heavy coarse aggregate, especially of steel aggregate used in nuclear  shields, is eliminated. In this case, pozzolana should not be used because it reduces the  density of the concrete and fixes less water. Because of the reduced segregatio
n,  preplaced aggregate concrete is also suitable for underwater construction, and indeed the  technique there differs little from placing under normal conditions. It may be relevant to  mention here that for concreting in depths of water over 30 m (100 ft) a tremie should be used.  For large concrete placement under water, there exist specialized
 systems.   The drying shrinkage of preplaced aggregate concrete is lower than that of ordinary  concrete, usually 200 x 10'-6 to 400 x 10'-6. The reduced shrinkage is due to the point-to-point  contact of the coarse aggregate particles, without a clearance for the cement paste  necessary in ordinary concrete. This contact restrains the amount of 
shrinkage that can  actually be realized, but occasionally shrinkage cracking can develop. Because of the  reduced shrinkage, preplaced aggregate concrete is suitable for the construction of  water-retaining and large monolithic structures and for repair work. The low permeability of  preplaced aggregate concrete gives it a high resistance to free
zing and thawing.   Preplaced aggregate concrete may be used in mass construction where the temperature  rise has to be controlled: cooling can be achieved by circulating refrigerated water round  the aggregate and thus chilling it; the water is later displaced by the rising mortar. At the  other extreme, in cold weather when frost damage is feare
d, steam can be circulated in  order to pre-heat the aggregate.	Preplaced aggregate concrete is used also to provide an exposed aggregate finish: special  aggregates are placed against the surfaces and become subsequently exposed by  sandblasting or by acid wash.   Preplaced aggregate concrete appears thus to have many useful features, but becau
se of  numerous practical difficulties considerable skill and experience in application of the  process are necessary for good results to be obtained.      Vacuum-processed Concrete    One solution to the problem of combining a sufficiently high workability with a minimum  water / cement ratio is offered by vacuum-processing of freshly placed conc
rete. Such  concrete is usually referred to as vacuum concrete - a rather misleading term but one in  common usage.   The procedure is briefly as follows. A mix with a medium workability is placed in the  forms in the usual manner. Since fresh concrete contains a continuous system of water-filled  channels, the application of a vacuum to the surfa
ce of the concrete results in a large  amount of water being extracted from a certain depth of the concrete. In other words,  what might be termed "water of workability" is removed when no longer needed. It may  be noted that air bubbles are removed only from the surface since they do not form a  continuous system.  The final water / cement ratio 
before setting is thus reduced, and, as this ratio largely  controls the strength, vacuum-processed concrete has a higher strength and also a higher  density, a lower permeability and a greater durability than would otherwise be obtained.  The magnitude of the decrease in the water / cement ratio due to vacuum-processing is  given in Table 4.8. Ho
wever, some of the water extracted leaves behind voids, so  that the full theoretical advantage of water removal may not be achieved in practice. In  fact, the increase in strength on vacuum treatment is proportional to the amount of water  removed up to a critical value beyond which there is no significant increases, so that  prolonged vacuum tre
atment is not useful. The critical value depends on the thickness of  concrete and on the mix proportions. Nevertheless, the strength of vacuum-processed  concrete broadly follows the usual dependence on the final water / cement ratio, as shown  in Fig. 4.28.    Table 4.8: Water / Cement Ratio and Strength of Vacuum-Processed Concrete    Fig. 4.28
. Relation between the strength of concrete and the calculated water / cement ratio  after vacuum treatment    The vacuum is applied through porous mats connected to a vacuum pump. The mats  consist of an airtight cover, usually made of plywood, with a vacuum chamber formed by  expanded metal. This is faced with a fine wire gauze covered by muslin
 which prevents the  removal of cement together with the water. A diagrammatic representation of a mat is  shown in Fig. 4.29. The mats can be placed on top of the concrete immediately after  screeding, and can also be incorporated in the inside faces of forms.    Fig. 4.29. Cross-section of a vacuum mat     Vacuum is created by a vacuum pump; its
 capacity is governed by the perimeter of the  mat and not its area. The magnitude of the applied vacuum is usually in the range of 400  to 650 mm (15 to 25 in.) of mercury. This vacuum reduces the water content by up to 20  per cent over a depth of 150 to 300 mm (6 to 12 in.). The reduction is greater nearer to  the mat and it is usual to assume 
the suction to be fully effective over a depth of 150 mm  (6 in.) only. Thus a concrete section 300 mm (12 in.) thick should have a vacuum applied  from two opposite faces. The withdrawal of water produces settlement of the concrete to  the extent of about 3 per cent of the depth over which the suction acts. The rate of  withdrawal of water falls 
off with time, and it has been found that processing during 15 to  25 minutes is usually most economical. Little reduction in water content occurs beyond 30  minutes.   It has been said that, strictly speaking, no suction of water takes place during vacuum-processing  but merely a fall of pressure below atmospheric is communicated to the interstit
ial fluid  of the fresh concrete. This would mean that compaction by atmospheric  pressure is taking place. Thus the amount of water removed would be equal to the  contraction in the total volume of concrete and no voids would be produced. However, in  practice some voids are formed and for the same final water / cement ratio ordinary  concrete ha
s been found to have a somewhat higher strength than vacuum-processed  concrete. This is discernible in Fig. 4.28.   The formation of voids can be prevented if in addition to vacuum-processing intermittent  vibration is applied; under those circumstances a higher degree of consolidation is  achieved and the amount of water withdrawn can be nearly 
doubled. In tests by  Garnett good results were obtained with vacuum-processing for 20 minutes  accompanied by vibration between the 4th and 8th minutes, and again between the 14th  and 18th minutes.   Vacuum-processing can be used over a fairly wide range of aggregate / cement ratios and  aggregate gradings, but a coarser grading yields more wate
r than a finer one. Furthermore,  some of the finest material is removed by the processing, and fine additions, such as pozzolanas,  should  not be incorporated in the mix.   Vacuum-processed concrete stiffens very rapidly so that formwork can be removed within  about 30 minutes of casting, even on columns 4.5 m (15 ft) high. This is of considerab
le  economic value, particularly in a precast factory, as the forms can be re-used at frequent  intervals.   The surface of vacuum-processed concrete is entirely free from pitting and the uppermost  1 mm (0.04 in.) is highly resistent to abrasion. These characteristics are of special  importance in the construction of concrete which is to be in co
ntact with water flowing at  a high velocity. Another useful characteristic of vacuum-processed concrete is that it  bonds well to old concrete and can, therefore, be used for re-surfacing road slabs and in  other repair work. Vacuum treatment thus appears to be a valuable process; it is, however,  rather an expensive one and is now less used in t
he U.S. but more extensively in  Scandinavian countries especially for slabs and floors.      Shotcrete    This is the name given to mortar or concrete conveyed through a hose and pneumatically  projected at high velocity onto a backup surface. The force of the jet impacting on the  surface compacts the material so that it can support itself witho
ut sagging or sloughing  even on a vertical face or overhead.   Since the essence of the process is that the material is projected pneumatically, shotcrete  is more formally called pneumatically applied mortar or concrete; it is also known as  gunite, although, in the U.S., this name applies only to shotcrete placed by the dry mix  process. In Bri
tain the general term sprayed concrete has recently been introduced.   The properties of shotcrete are no different from the properties of conventionally placed  mortar or concrete of similar proportions: it is the method of placing that bestows on  shotcrete significant advantages in many applications. At the same time, considerable skill  and ex
perience are required in the application of shotcrete so that its quality depends to a  large extent on the performance of the operators involved, especially in control of the actual  placing by the nozzle.   Since shotcrete is pneumatically projected on a backup surface and then gradually built-up,  only one side of formwork is needed. This repre
sents economy, especially when account is  taken of the absence of form ties, etc. On the other hand, the cement content of shotcrete  is high. Also, the necessary equipment and mode of placing are more expensive than in the  case of conventional concrete. For these reasons, shotcrete is used primarily in certain types  of construction: thin, ligh
tly reinforced sections, such as roofs, especially shell or folded plate,  tunnel linings, and prestressed tanks. Shotcrete is also used in repair of deteriorated concrete, in  stabilizing rock slopes, in encasing steel for fireproofing, and as a thin overlay on concrete, masonry  or steel. If shotcrete is applied to a surface covered by running w
ater, an accelerator producing  flash set, such as washing soda, is used. This adversely affects strength but makes repair work  possible. Generally, shotcrete is applied in a thickness up to 100 mm (4 in.).   There are two basic processes by which shotcrete is applied. In the dry mix process  (which is the more common of the two) cement and damp 
aggregate are intimately mixed  and fed into a mechanical feeder or gun. The mixture is then transferred by a feed wheel or  distributor (at a known rate) into a stream of compressed air in a hose, and carried up to  the delivery nozzle. The nozzle is fitted inside with a perforated manifold through which  water is introduced under pressure and in
timately mixed with the other ingredients. The  mixture is then projected at high velocity onto the surface to be shotcreted.   The fundamental feature of the wet mix process is that all the ingredients, including the  mixing water, are mixed together to begin with. The mixture is then introduced into the  chamber of the delivery equipment and fro
m there conveyed pneumatically or by positive  displacement. A pump similar to that of Fig. 4.21 can be used. Compressed air (or in the  case of pneumatically conveyed mix, additional air) is injected at the nozzle, and the  material is projected at high velocity onto the surface to be shotcreted.   Either process can produce excellent shotcrete, 
but the dry mix process is better suited  for use with porous lightweight aggregate and with flash set accelerators, and is also  capable of greater delivery lengths. On the other hand, the wet mix process gives a  better control of the quantity of mixing water (which is metered, as opposed to judgement  by the nozzle operator) and of any admixtur
e used. Also, the wet mix process leads to less  dust being produced.   Not all the shotcrete projected on a surface remains in position. Because of the high  velocity of the impacting jet, some material rebounds. This consists of the coarsest  particles in the mix, so that the shotcrete in situ is richer than would be expected from the  mix propo
rtions as batched. This may lead to increased shrinkage. The rebound is greatest  in the initial layers and becomes smaller as a plastic cushion of shotcrete is built-up.  Typical percentages of material rebounded are:    in floors and slabs   5 to 15  on sloping or vertical surfaces  15 to 30  on soffits    25 to 50.     The significance of rebou
nd is not so much in the waste of the material as in the danger  from accumulation of rebounded sand in a position where it will become incorporated in  the subsequent layers of shotcrete. This can occur if the rebound collects in inside corners,  at the base of walls, behind reinforcement or embedded pipes, or on horizontal surfaces.  Great care 
in placing of shotcrete is therefore necessary, and the use of large  reinforcement is undesirable. The latter also leads to the risk of unfilled pockets behind the  obstacle to the jet.   The projected shotcrete has to have a relatively dry consistence so that the material can  support itself in any position; at the same time, the mix has to be w
et enough to obtain  compaction without excessive rebound. The usual range of water / cement ratios is 0.35 to  0.50; there is generally very little bleeding. In the case of mortar, the usual mix is 1:3.5 to  1:4.5, and 28-day strengths ranging between 20 and 50 MPa (3000 and 7000 psi) are  obtained. Sand with the same grading as for use in conven
tional mortar is satisfactory.   Concrete can also be shotcreted, the maximum aggregate size being 25 mm (1 in.) but the  need for this material is small and its advantages are limited. Also, the larger the aggregate  the larger the rebound. The coarse aggregate content is lower than in conventionally  placed concrete.   Curing of shotcrete is par
ticularly important because the large surface - volume ratio can  lead to rapid drying. Recommended practices are given in ACI Standards 506.2-77  and 506-66.      Analysis of Fresh Concrete    In considering the ingredients of a concrete mix we have so far assumed that the actual  proportions correspond to those specified. If this were invariably
 so there would be little  need for testing the strength of hardened concrete. However, in practice, mistakes, errors  and even deliberate actions can lead to incorrect mix proportions, and it is sometimes  useful to determine the composition of concrete at an early stage; the two values of  greatest interest are the cement content and water / cem
ent ratio.   The existing standardized test is that prescribed by BS 1881 : Part 2 : 1970. The test has  to be commenced virtually as soon as the concrete has been discharged from the mixer  because loss of water can occur; even if evaporation is prevented, an unknown amount of  hydration will take place during any period of delay. The analysis of
 the concrete has to be  supplemented by tests on the fine and coarse fractions of the aggregate: specific gravity,  absorption, and grading. This information is essential in the calculation of the quantity of  the solid constituents of the mix. The weights of fine and coarse aggregate are determined  by weighing in water, and the weight of cement
 is obtained as a difference between the  weight in water of the concrete and of the aggregates. The water content is  determined by drying the sample over a heater. Hence, the mix proportions can be  calculated. The air content is determined independently but the analysis of air-entrained  concrete is less accurate than of non-air-entrained concr
ete.   A slightly different method of determination of the water content of fresh concrete was  suggested by Nischer. A sample of such concrete is mixed with 10 per cent of its  weight of alcohol and the mixture is burnt while being stirred. The process is repeated till  the concrete has lost all cohesion and has a uniform light colour. The loss i
n weight gives  the water content.   The water content of aggregate and of fresh concrete can be measured also by estimating  the degree of scattering of thermal neutrons emitted by a source placed within the bulk of  the aggregate or within a sample of the mix. Hydrogen is the most important element  influencing scattering and retardation of ther
mal neutrons, and, since hydrogen is almost  exclusively bound in water, the nuclear method can provide a value of the water content  with an accuracy of +/-0.3 per cent. The technique also requires the dry density of the  aggregate to be taken into account, and this is calculated from the back-scattering of  gamma radiation from a second source. 
The complete apparatus comprises gamma and  thermal neutron sources, neutron and scintillation detectors, and associated counters.  Calibration is carried out in situ and is a time-consuming process.   An accurate determination of the cement content is of major importance and several  approaches have been proposed. Turton suggested a modification 
of the BS 1881 :  Part 2: 1970 procedure: he collects the washed very fine material into a container, and  obtains the weight of cement direct rather than as a difference. The calculation of the  weight of cement has to include a correction for the silt and dust in the aggregate as  determined by a sieve analysis on the aggregate samples. In anoth
er variant, the material  smaller than 150 um (No. 100 sieve) is separated out by filtering and pressing dry;  the weight of cement is taken as the weight of this fraction corrected for aggregate finer  than  150 um (No. 100 sieve) in the material as batched. Separation of cement by flotation has  recently been developed.   The U.S. Army uses a te
st which relies on chloride titration for determining the water  content and on calcium titration for cement content. The test can be performed in the field  and takes no more than a quarter of an hour. However, the fine part (smaller than 150 um  (No. 100 sieve)) of calcareous aggregate cannot be distinguished from the cement.   Another method of
 determining the cement content of fresh concrete consists of the  measurement of electrical conductivity of an aqueous suspension of the material. The test  is based on the fact that hydration of cement produces salts whose concentration is  proportional to the quantity of cement so that there is a definite relation between the  cement content an
d	electrical conductivity. The test must be performed within an hour of  mixing but the limitations of the test conditions (e.g. temperature) are such that the  determination cannot be made on site.   A totally different approach in the determination of cement content of fresh concrete is  based on the separation of cement using a heavy liquid and
 a centrifuge. This has not  been very successful, especially when the finest aggregate particles have a specific gravity  not significantly lower than that of cement.   The so-called Rapid Analysis Machine developed by the Cement and Concrete  Association  can also be used. An 8 +/- 1 kg (18 +/- 1 lb) sample of concrete is put into an elutriation
	column and material smaller than 600 um (No. 30 ASTM sieve) is lifted. A part of this  slurry is vibrated on a 150 um (No. 100) sieve, then flocculated and transferred into a  constant volume vessel. This is weighed and, using a calibration chart, the cement content  of the sample is determined. A correction for aggregate particles smaller than 
150 um (No.  100) has to be made. The calibration has to be performed for each set of materials used.   There is some argument about the variability of the test. Cooper and Barber found  that the standard deviation of the cement content is equivalent to about  22 kg / m3 (37 lb / yd3)  of concrete, of which some 13 kg / m3 (22 lb / yd3) is due to 
variations in the sample  and 15 kg / m3 (25 lb / yd3) to variations in the machine. Thus, for a mix with the cement  content of 370 kg / m3 (620 lb / yd3), the 95 per cent confidence limits would be +/-43 kg  / m3 (+/-72 lb / yd3). The use of duplicate sub-samples would reduce the variability but,  according to Cooper and Barber, the accuracy of 
the test is unsatisfactory in that the  cement content is underestimated, typically by 26 kg / m3 (44 lb / yd3).   No doubt, improvements in the Rapid Analysis Machine test are possible, and of course  the psychological effects of the knowledge that the test is being used on those responsible  for mix control should not be ignored. Nevertheless, i
t	seems that no test for the  composition of fresh concrete that is convenient and reliable enough to be used as a  preplacement acceptance test is as yet available.  13    Properties of Fresh Concrete    Fresh concrete is a mixture of water, cement, aggregate and admixture (if any).  After mixing, operations such as transporting, placing, compact
ing and finishing of fresh  concrete can all considerably affect the properties of hardened concrete. It is important  that the constituent materials remain uniformly distributed within the concrete mass  during the various stages of its handling and that full compaction is achieved. When either  of these conditions is not satisfied the properties
 of the resulting hardened concrete, for  example, strength and durability, are adversely affected.   The characteristics of fresh concrete which affect full compaction are its  consistency, mobility and compactability. In concrete practice these are often  collectively known as workability. The ability of concrete to maintain its uniformity is  g
overned by its stability which depends on its consistency and its cohesiveness. Since the  methods employed for conveying, placing and consolidating a concrete mix, as well as the  nature of the section to be cast, may vary from job to job it follows that the corresponding  workability and stability requirements will also vary. The assessment of t
he suitability of a  fresh concrete for a particular job will always to some extent remain a matter of personal  judgement.   In spite of its importance, the behaviour of plastic concrete often tends to be  overlooked. It is recommended that students should learn to appreciate the significance of  the various characteristics of concrete in its pla
stic state and know how these may alter  during operations involved in casting a concrete structure.      13.1 Workability    Workability of concrete has never been precisely defined. For practical purposes  it generally implies the ease with which a concrete mix can be handled from the mixer to its  finally compacted shape. The three main charact
eristics of the property are consistency,  mobility and compactability. Consistency is a measure of wetness or fluidity. Mobility  defines the ease with which a mix can flow into and completely fill the formwork or mould.  Compactability is the ease with which a given mix can be fully compacted, all the trapped  air being removed. In this context 
the required workability of a mix depends not only on  the characteristics and relative proportions of the constituent materials but also  on (1) the methods employed for conveyance and compaction, (2) the size, shape and  surface roughness of formwork or moulds and (3) the quantity and spacing of  reinforcement.   Another commonly accepted defini
tion of workability is related to the amount  of useful internal work necessary to produce full compaction. It should be appreciated that  the necessary work again depends on the nature of the section being cast. Measurement of  internal work presents many difficulties and several methods have been developed for this  purpose but none gives an abs
olute measure of workability.   The tests commonly used for measuring workability do not measure the individual  characteristics (consistency, mobility and compactability) of workability. However, they do  provide useful and practical guidance on the workability of a mix. Workability affects the  quality of concrete and has a direct bearing on cos
t	so that, for example, an unworkable  concrete mix requires more time and labour for full compaction. It is most important that a  realistic assessment is made of the workability required for given site conditions before any  decision is taken regarding suitable concrete mix proportions.      13.2 Measurement of Workability    Three tests widely 
used for measuring workability are the slump, compacting  factor and Vebe time tests (figure 13.1). These are standard tests in the United Kingdom  and are described in detail in BS 1881: Parts 102, 103 and 104 respectively. Their use is  also recommended in BS 5328, which also requires the measured workability of concrete  to be within certain li
mits of the required value as given in table 13.1. For flowing  concrete, normal workability tests are not sensitive enough and a flow test based on the  German Standard DIN 1048 has recently been adopted as the British Standard test BS  1881: Part 105. It is important to note that there is no single relationship between the  slump, compacting fac
tor, Vebe time and flow results for different concretes. In the  following sections the salient features of these tests together with their merits and  limitations are discussed.     TABLE 13.1 Permissible variations in concrete workability    Figure 13.1 Apparatus for workability measurement: (a) slump cone, (b) compacting  factor and (c) Vebe co
nsistometer    Slump Test    This test was developed by Chapman in the United States in 1913. A 300 mm high  concrete cone, prepared under standard conditions (BS 1881: Part 102) is allowed to  subside and the slump or reduction in height of the cone is taken to be a measure of  workability. The apparatus is inexpensive, portable and robust and is
 the simplest of all  the methods employed for measuring workability. It is not surprising that, in spite of  its several limitations, the slump test has retained its popularity.   The test primarily measures the consistency of plastic concrete and although it  is difficult to see any significant relationship between slump and workability as defin
ed  previously, it is suitable for detecting changes in workability. For example, an increase in  the water content or deficiency in the proportion of fine aggregate results in an increase  in slump. Although the test is suitable for quality-control purposes it should be  remembered that it is generally considered to be unsuitable for mix design s
ince concretes  requiring varying amounts of work for compaction can have  similar numerical values of slump. The sensitivity and reliability of the test for detecting  variation in mixes of different workabilities is largely dependent on its sensitivity to  consistency. The test is not suitable for very dry or wet mixes. For very dry mixes, with 
 zero or near-zero slump, moderate variations in workability do not result in measurable  changes in slump. For wet mixes, complete collapse of the concrete produces unreliable  values of slump.   The three types of slump usually observed are true slump, shear slump and  collapse slump, as illustrated in figure 13.2. A true slump is observed with 
cohesive  and rich mixes for which the slump is generally sensitive to variations in workability. A  collapse slump is usually associated with very wet mixes and is generally indicative of  poor quality concrete and most frequently results from segregation of its constituent  materials. Shear slump occurs more often in leaner mixes than in rich on
es and indicates a  lack of cohesion which is generally associated with harsh mixes (low mortar content).  Whenever a shear slump is obtained the test should be repeated and, if persistent, this fact  should be recorded together with test results, because widely different values of slump can  be obtained depending on whether the slump is of true o
r shear form.  The standard slump apparatus is only suitable for concretes in which the  maximum aggregate size does not exceed 40 mm. It should be noted that the value of  slump changes with time after mixing owing to normal hydration processes and  evaporation of some of the free water, and it is desirable  therefore that tests are performed wit
hin a fixed period of time. Figure 13.2 Three main types of slump    Compacting Factor Test    This test, developed in the United Kingdom by Glanville et al. (1947), measures the degree of  compaction for a standard amount of work and thus offers a direct and reasonably reliable assessment  of the workability of concrete as previously defined. The
 apparatus is a relatively simple mechanical  contrivance (figure 13.1) and is fully described in BS 1881: Part 103. The test requires measurement  of the weights of the partially and fully compacted concrete and the ratio of the partially compacted  weight to the fully compacted weight, which is always less than 1, is known as the compacting fact
or.  For the normal range of concretes the compacting factor lies between 0.80 and 0.92. The test is  particularly useful for drier mixes for which the slump test is not satisfactory. The sensitivity of the  compacting factor is reduced outside the normal range of workabilities. However, BS 1881: Part 103  only regards the test unsuitable for conc
rete having the measured compacting factor below 0.70 or  above 0.98.  It should also be appreciated that, strictly speaking, some of the basic assumptions of the test are not  correct. The work done to overcome surface friction of the measuring cylinder probably varies with  the characteristics of the mix. It has been shown by Cusens (1956) that 
for concretes with very low  workability the actual work required to obtain full compaction depends on the richness of a mix while  the compacting factor remains sensibly unaffected. Thus it follows that the generally held belief that  concretes with the same compacting factor require the same amount of work for full compaction  cannot always be j
ustified. One further point to note is that the procedure for placing concrete in the  measuring cylinder bears no resemblance to methods commonly employed on the site. As in the slump  test, the measurement of compacting factor must be made within a certain specified period. The  standard apparatus is suitable for concrete with a maximum aggregat
e size of up to 40 mm.    Vebe Test    This test was developed in Sweden by Bahrner (1940) (see figure 13.1). Although generally regarded  as a test primarily used in research its potential is now more widely acknowledged in industry and the  test is gradually being accepted. In this test (BS 1881: Part 104) the time taken to transform, by means  
of vibration, a standard cone of concrete to a compacted flat cylindrical mass is recorded in seconds,  to the nearest 0.5 s. Unlike the two previous tests, the treatment of concrete in this test is comparable  to the method of compacting concrete in practice. Moreover, the test is sensitive to change in  consistency, mobility and compactability, 
and therefore a reasonable correlation between the test  results and site assessment of workability can be expected.   The test is suitable for a wide range of mixes and, unlike the slump and compacting factor tests, it is  sensitive to variations in workability of very dry and also air-entrained concretes. It is also more  sensitive to variation 
in aggregate characteristics such as shape and surface texture. The  reproducibility of results is good. As for other tests, its accuracy tends to decrease with increasing  maximum size of aggregate; above 20 mm the test results become somewhat unreliable. However, BS  1881: Part 104 permits its use for concrete having aggregate of maximum size up
 to 40 mm. For  concretes requiring very little vibration for compaction the Vebe time is only about 3 s. Such results  are likely to be less reliable than for larger Vebe times because of the difficulty in estimating the time  of the end point (concrete in contact with the whole of the underside of the plastic disc). At the other  end of the work
ability range, such as with very dry mixes, the recorded Vebe times are likely to be in  excess of their true workability since prolonged vibration is required to remove the entrapped air  bubbles under the transparent disc. To overcome this difficulty an automatic device which records the  vertical settlement of the disc with respect to time can 
be attached to the apparatus. This recording  device can also assist in eliminating human error in judging the end point. The apparatus for the Vebe  test is more expensive than that for the slump and compacting factor tests, requiring an electric power  supply and greater experience in handling; all these factors make it more suitable for the pre
cast  concrete industry and ready-mixed concrete plants than for general site use.    Flow Test    The test, a slightly modified form of the German Standard DIN 1048 test, is used to measure the  ability of concrete to flow (BS 1881: Part 105). The apparatus consists essentially of a 700 mm square  wooden table (covered by a flat steel plate) whic
h is hinged along one side to permit a vertical  movement of 40 mm, a slump cone (200 mm high, internal diameters 200 mm bottom and 130 mm  top) and a wooden tamper (40 mm square base). The cone is placed at the centre of the table and filled  with concrete in two equal layers, each layer being compacted by 10 light strokes of the tamper. The  exc
ess concrete is struck off the cone and the table top cleaned immediately; 30 seconds later the cone  is removed gently. The free side of the table is then raised 40 mm (upper stop) and allowed to fall  freely 15 times in about 45 to 75 seconds. The resulting spread of concrete is measured along its two  diameters parallel to the sides of the tabl
e and the mean diameter, expressed to the nearest 5 mm, is  taken as the measurement of flow. The flow test is valid only if the concrete is cohesive and uniform  at the conclusion of the test. BS 1881: Part 105 limits the use of the flow test to concrete having  maximum aggregate size up to 20 mm and a measured flow diameter between 500 and 650 m
m.      13.3 Factors affecting Workability    Various factors known to influence the workability of a freshly mixed concrete are shown in figure  13.3. From the following discussion it will be apparent that a change in workability associated with  the constituent materials is mainly affected by water content and specific surface of cement and  agg
regate.    Figure 13.3 Factors affecting workability of fresh concrete    Cement and Water    Typical relationships between the cement - water ratio (by volume) and the volume fraction of cement  for different workabilities are shown in figure 15.5. Hughes (1971) has shown that similar linear  relationships exist irrespective of the properties of 
the constituent materials. Workability is relatively  insensitive to changes in only the cement content and, for practical purposes, may be considered  dependent on only the water content for variations in cement content of up to 10 to 20 per cent  depending on the cement content, the effect of cement content being greater for the richer mixes.  F
or a given mix, the workability of the concrete decreases as the fineness of the  cement increases as a result of the increased specific surface, this effect being more marked in rich  mixtures. It should also be noted that the finer cements improve the cohesiveness of a mix. With the  exception of gypsum, the composition of cement has no apparent
 effect on workability. Unstable  gypsum is responsible for false set, which can impair workability unless prolonged mixing or  remixing of the fresh concrete is carried out. Variations in quality of water suitable for making  concrete have no significant effect on workability.    Admixtures    The principal admixtures effecting improvement in the
 workability of concrete are water-reducing and  air-entraining agents. The extent of the increase in workability is dependent on the type and amount  of admixture used and the general characteristics of the fresh concrete.  Water-reducing admixtures are used to increase workability while the mix proportions are kept  constant or to reduce the wat
er content while maintaining constant workability. The former may result  in a slight reduction in concrete strength. The water-reducing admixtures, including superplasticizers,  are now more widely accepted as a workability aid. It should also be noted that the use of  pulverized-fuel ash can also result in a significant improvement in concrete w
orkability.  Air-entraining agents are by far the most commonly used workability admixtures because they also  improve both the cohesiveness of the plastic concrete and the frost resistance of the resulting hardened  concrete. Two points of practical importance concerning air-entrained concrete are that for a given  amount of entrained air, the in
crease in workability tends to be smaller for concretes containing  rounded aggregates or low cement - water ratios (by volume) and, in general, the rate of increase in  workability tends to decrease with increasing air content. However, as a guide it may be assumed that  every 1 per cent increase in air content will increase the compacting factor
 by 0.01 and reduce the  Vebe time by 10 per cent.    Figure 13.4 Effect of aggregate shape on aggregate - cement ratio of concretes for different  workabilities, based on Cornelius (1970)    Aggregate    For given cement, water and aggregate contents, the workability of concrete is mainly influenced by  the total surface area of the aggregate. Th
e surface area is governed by the maximum size, grading  and shape of the aggregate. Workability decreases as the specific surface increases, since this requires  a greater proportion of cement paste to wet the aggregate particles, thus leaving a smaller amount of  paste for lubrication. It follows that, all other conditions being equal, the worka
bility will be increased  when the maximum size of aggregate increases, the aggregate particles become rounded or the overall  grading becomes coarser. However, the magnitude of this change in workability depends on the mix  proportions, the effect of the aggregate being negligible for very rich mixes (aggregate - cement ratios  approaching 2). Th
e practical significance of this is that for a given workability and cement - water  ratio the amount of aggregate which can be used in a mix varies depending on the shape, maximum  size and grading of the aggregate, as shown in figure 13.4 and tables 13.2 and 13.3 . The influence  of air-entrainment (4.5 per cent) on workability is shown also in 
figure 13.4.  Several methods have been developed for evaluating the shape of aggregate, a  subject discussed in chapter 12. Angularity factors together with grading modulus and equivalent  mean diameter provide a means of considering the respective effects of shape, size and grading of  aggregate (see chapter 15). Since the strength of a fully co
mpacted concrete, for given materials and  cement - water ratio, is not dependent on the ratio of coarse to fine aggregate, maximum economy can  be obtained by using the coarse aggregate content producing the maximum workability for a given  cement content (Hughes, 1960) (see figure 13.5). The use of optimum coarse aggregate content in  concrete m
ix design is described in chapter 15. It should be noted that it is the volume fraction of an  aggregate, rather than its weight, which is important.  The effect of surface texture on workability is shown in figure 13.6. It can be  seen that aggregates with a smooth texture result in higher workabilities than aggregates with a rough  texture. Abso
rption characteristics of aggregate also affect workability where dry or partially dry  aggregates are used. In such a case workability drops, the extent of the reduction being dependent on  the aggregate content and its absorption capacity.   TABLE 13.2 Effect of maximum size of aggregate of similar grading zone on aggregate - cement  ratio of co
ncrete having water - cement ratio of 0.55 by weight, based on McIntosh (1964)    TABLE 13.3 Effect of aggregate grading (maximum size 19. 0 mm) on aggregate - cement ratio of  concrete having medium workability and water - cement ratio of 0.55 by weight, based on McIntosh  (1964)    Figure 13.5 A typical relationship between workability and coars
e	aggregate content of concrete,  based on Hughes (1960)    Figure 13.6 Effect of aggregate surface texture on aggregate - cement ratio of concretes for  different workabilities, based on Cornelius (1970)   Ambient Conditions    Environmental factors that may cause a reduction in workability are temperature, humidity and wind  velocity. For a give
n	concrete, changes in workability are governed by the rate of hydration of the  cement and the rate of evaporation of water. Therefore both the time interval from the commencement  of mixing to compaction and the conditions of exposure influence the reduction in workability. An  increase in the temperature speeds up the rate at which water is use
d	for hydration as well as its loss  through evaporation. Likewise wind velocity and humidity influence the workability as they affect the  rate of evaporation. It is worth remembering that in practice these factors depend on weather  conditions and cannot be controlled. Time   The time that elapses between mixing of concrete and its final compact
ion depends on the general  conditions of work such as the distance between the mixer and the point of placing, site procedures  and general management. The associated reduction in workability is a direct result of loss of free  water with time through evaporation, aggregate absorption and initial hydration of the cement. The  rate of loss of work
ability is affected by certain characteristics of the constituent materials, for  example, hydration and heat development characteristics of the cement, initial moisture content and  porosity of the aggregate, as well as the ambient conditions.    Figure 13.7 Loss of workability of concrete with time: (A) no agitation and (B) continuously  agitate
d	after mixing    For a given concrete and set of ambient conditions, the rate of loss of workability with time depends  on the conditions of handling. Where concrete remains undisturbed after mixing until it is placed, the  loss of workability during the first hour can be substantial, the rate of loss of workability decreasing  with time as illus
trated by curve A in figure 13.7. On the other hand, if it is continuously agitated, as  in the case of ready-mixed concrete, the loss of workability is reduced, particularly during the first  hour or so (see curve B in figure 13.7). However, prolonged agitation during transportation may  increase the fineness of the solid particles through abrasi
on and produce a further reduction in  workability. For concretes continuously agitated and undisturbed during transportation, the time  intervals permitted (BS 5328) between the commencement of mixing and delivery on site are 2 hours  and 1 hour respectively.  For practical purposes, loss of workability assumes importance when concrete  becomes s
o	unworkable that it cannot be effectively compacted, with the result that its strength and  other properties become adversely affected. Corrective measures frequently taken to ensure that  concrete at the time of placing has the desired workability are either an initial increase in the water  content or an increase in the water content with furth
er mixing shortly before the concrete is  discharged. When this results in a water content greater than that originally intended, some reduction  in strength and durability of the hardened concrete is to be expected unless the cement content is  increased accordingly. This important fact is frequently overlooked on site. It should be recalled that
	the loss of workability varies with the mix, the ambient conditions, the handling conditions and the  delivery time. No restriction on delivery time is given in BS 8110: Part 1 but the concrete must be  capable of being placed and effectively compacted without the addition of further water. For detailed  information on the use of ready-mixed con
crete the reader is advised to consult the work of Dewar  (1973).   13.4 Stability    Apart from being sufficiently workable, fresh concrete should have a composition such that  its constituent materials remain uniformly distributed in the concrete during both the period  between mixing and compaction and the period following compaction before the
 concrete  stiffens. Because of differences in the particle size and specific gravities of the constituent  materials there exists a natural tendency for them to separate. This tendency is generally  greater in high workability mixes. Flowing concrete is particularly prone to it and should  be more carefully designed; as a guide a minimum of 450 k
g	m-3 combined fines (cement  plus sand fractions less than 0.3 mm) should be used. Concrete capable of maintaining the  required uniformity is said to be stable and most cohesive mixes belong to this category.  For an unstable mix the extent to which the constituent materials will separate depends  on the methods of transportation, placing and co
mpaction. The two most common features  of an unstable concrete are segregation and bleeding.    Segregation    When there is a significant tendency for the large and fine particles in a mix to become  separated segregation is said to have occurred. In general, the less cohesive the mix the  greater the tendency for segregation to occur. Segregati
on is governed by the total specific  surface of the solid particles including cement and the quantity of mortar in the mix. Harsh,  extremely wet and dry mixes as well as those deficient in sand, particularly the finer  particles, are prone to segregation. As far as possible, conditions conducive to segregation  such as jolting of concrete during
 transportation, dropping from  excessive heights during placing and over-vibration during compaction should be avoided.  Blemishes, sand streaks, porous layers and honeycombing are a direct result of  segregation. These features are not only unsightly but also adversely affect strength,  durability and other properties of the hardened concrete. I
t	is important to realise that the  effects of segregation may not be indicated by the routine strength tests on control  specimens since the conditions of placing and compaction of the specimens differ from  those in the actual structure. There are no specific rules for suspecting possible segregation  but after some experience of mixing and hand
ling concrete it is not difficult to recognise  mixes where this is likely to occur. For example if a handful of concrete is  squeezed in the hand and then released so that it lies in the palm, a cohesive concrete will  be seen to retain its shape. A concrete which does not retain its shape under these  conditions may well be prone to segregation 
and this is particularly so for wet mixes.    Bleeding    During compaction and until the cement paste has hardened there is a natural tendency for  the solid particles, depending on size and specific gravity, to exhibit a downward  movement. Where the consistency of a mix is such that it is unable to hold all its water  some of it is gradually di
splaced and rises to the surface, and some may also leak through  the joints of the formwork. Separation of water from a mix in this manner is known as  bleeding. While some of the water reaches the top surface some may become trapped  under the larger particles and under the reinforcing bars. The resulting  variations in the effective water conte
nt within a concrete mass produce corresponding  changes in its properties. For example, the strength of the concrete immediately  underneath the reinforcing bars and coarse aggregate particles may be much less than the  average strength and the resistance to percolation of water in these areas is reduced. In  general, the concrete strength tends 
to increase with depth below the top surface. The  water which reaches the top surface presents the most serious  practical problems. If it is not removed, the concrete at and near the top surface will be  much weaker and less durable than the remainder of the concrete. This can be particularly  troublesome in slabs which have a large surface area
. On the other hand, removal of the  surface water will unduly delay the finishing operation on the site.   The risk of bleeding increases when concrete is compacted by vibration although  this may be minimised by using a correctly designed mix and ensuring that the concrete is  not over-vibrated. Rich mixes tend to bleed less than lean mixes. The
 type of cement  employed is also important, the tendency for bleeding to occur decreasing as the fineness  of the cement or its alkaline and tricalcium aluminate (C3A) content increases.  Air-entrainment provides another very effective means of controlling  bleeding in, for example, wet lean mixes where both segregation and bleeding are  frequent
ly troublesome.  1 Portland Cement      Cement, in the general sense of the word, can be described as a material with adhesive and  cohesive properties which make it capable of bonding mineral fragments into a compact  whole. This definition embraces a large variety of cementing materials.    For constructional purposes the meaning of the term cem
ent is restricted to the  bonding materials used with stones, sand, bricks, building blocks, etc. The principal  constituents of this type of cement are compounds of lime, so that in building and civil  engineering we are concerned with calcareous cement. The cements of interest in the  making of concrete have the property of setting and hardening
 under water by virtue of a  chemical reaction with it and are, therefore, called hydraulic cements.    Hydraulic cements consist mainly of silicates and aluminates of lime, and can be  classified broadly as natural cements, Portland cements and high-alumina cements. The  present chapter deals with the manufacture of Portland cement and its struct
ure and  properties both when unhydrated and in a hardened state. The different types of Portland  and other cements are described in Chapter 2.      Historical Note    The use of cementing materials is very old. The ancient Egyptians used calcined impure  gypsum. The Greeks and the Romans used calcined limestone and later learned to add to  lime 
and water, sand and crushed stone or brick and broken tiles. This was the first  concrete in history. Lime mortar does not harden under water, and for construction under  water the Romans ground together lime and a volcanic ash or finely ground burnt clay  tiles. The active silica and alumina in the ash and the tiles combined - with the lime to  p
roduce what became known as pozzolanic cement from the name of the village of  Pozzuoli, near Vesuvius, where the volcanic ash was first found. The name pozzolanic  cement is used to this day to describe cements obtained simply by the grinding of natural materials  at normal temperature. Some of the Roman structures in which masonry was bonded by 
 mortar, such as the Coliseum in Rome and the Pont-du-Gard near Nimes have survived to this  day, with the cementitious material still hard and firm. In the ruins at Pompeii, the mortar is  often less weathered than the rather soft stone.   The Middle Ages brought a general decline in the quality and use of cement, and  it is only in the 18th cent
ury that an advance in the knowledge of cements can be reported.  John Smeaton, commissioned in 1756 to rebuild the Eddystone Lighthouse, off the  Cornish coast, found that the best mortar was produced when pozzolana was mixed with  limestone containing a high proportion of clayey matter. By recognizing the role of the  clay, hitherto considered u
ndesirable, Smeaton was the first to understand the chemical  properties of hydraulic lime.   There followed a development of other hydraulic cements, such as the "Roman  cement" obtained by James Parker by calcining nodules of argillaceous limestone,  culminating in the patent for "Portland cement" taken out by Joseph Aspdin, a Leeds  builder, in
 1824. This cement was prepared by heating a mixture of finely-divided clay and  hard limestone in a furnace until CO2 had been driven off; this temperature was much  lower than that necessary for clinkering. The prototype of modern cement was made in  1845 by Isaac Johnson, who burnt a mixture of clay and chalk until clinkering, so that the  reac
tions necessary for the formation of strongly cementitious compounds took place.   The name Portland cement, given originally due to the resemblance of the colour  and quality of the set cement to Portland stone - a limestone quarried in Dorset - has  remained throughout the world to this day to describe a cement obtained by intimately  mixing tog
ether calcareous and argillaceous, or other silica-, alumina-, and iron oxide-bearing  materials, burning them at a clinkering temperature, and grinding the resulting  clinker. The definition of the current British Standard (BS 12: 1978) is on these lines; the  standard stipulates also that no materials other than gypsum and water may be added aft
er  burning (see also p. 83).      Manufacture of Portland Cement    From the definition of Portland cement given above it can be seen that it is made primarily  from a calcareous material, such as limestone or chalk, and from alumina and silica found  as clay or shale. Marl, a mixture of calcareous and argillaceous materials, is also used. In  Gr
eat Britain, chalk is found in the South-East, and for this reason the largest  concentration of cement works is near the mouth of the Thames and on the banks of the  Medway. Limestone occurs in many parts of the South-West, the Midlands, Northern  England and Wales, and argillaceous deposits are found throughout Britain. Raw materials  for the ma
nufacture of Portland cement are found in nearly all countries.   The process of manufacture of cement consists essentially of grinding the raw  materials, mixing them intimately in certain proportions and burning in a large rotary kiln  at a temperature of approximately 1400 C when the material sinters and partially fuses  into balls known as cli
nker. The clinker is cooled and ground to a fine powder, with some  gypsum added, and the resulting product is the commercial Portland cement so widely  used throughout the world.   Some details of the manufacture of cement will now be given, and these can best  be followed with reference to the diagrammatic representation of the process given in 
Fig.  1.1.   The mixing and grinding of the raw materials can be done either in water or in a  dry condition, hence the names "wet" and "dry" processes. The actual methods of  manufacture depend also on the hardness of the raw materials used and on their moisture  content.   Let us consider first the wet process. When chalk is used it is finely br
oken up  and dispersed in water in a washmill; this is a circular pit with revolving radial arms carrying  rakes which break up the lumps of solid matter. The clay is also broken up and mixed with  water, usually in a similar washmill. The two mixtures are now pumped so as to mix in  predetermined proportions and pass through a series of screens. 
The resulting cement  slurry flows into storage tanks.   When limestone is used it has to be blasted, then crushed, usually in two  progressively smaller crushers, and then fed into a ball mill with the clay dispersed in  water. There, the comminution of the limestone (to the fineness of flour) is completed, and  the resultant cement slurry is pum
ped into storage tanks. From here onwards the process  is the same regardless of the original nature of the raw materials.   The slurry is a liquid of creamy consistence, with a water content of between 35  and 50 per cent, and only a small fraction of material - about 2 per cent - larger than a 90  um (No. 170 ASTM) sieve size. There are usually 
a number of storage tanks in which the  slurry is kept, the sedimentation of the suspended solids being prevented by mechanical  stirrers or bubbling by compressed air. The lime content of the slurry is governed by the  proportioning of the original calcareous and argillaceous materials, as mentioned earlier.  Final adjustment in order to achieve 
the required chemical composition can be made by  blending slurries from different storage tanks, sometimes using an elaborate system of  blending tanks. Occasionally, for example in the world's northernmost plant in Norway,  the raw material is a rock of such composition that it alone is crushed and no blending is  required.   Finally, the slurry
 with the desired lime content passes into the rotary kiln. This is  a large, refractory-lined steel cylinder, up to 7.5 m (or 25 ft) in diameter, sometimes as long  as 230 m (or 760 ft), slowly rotating about its axis, which is slightly inclined to the  horizontal. The slurry is fed in at the upper end while pulverized coal is blown in by an air 
 blast at the lower end of the kiln, where the temperature reaches 1400 to 1500 C. The  coal, which must not have too high an ash content, deserves a special mention because  some 190 to 350 kg (420 to 770 1b) of coal is used to make one tonne of cement. This is  worth bearing in mind when considering the price of cement. Oil (of the order of 150 
litres  (40 U.S. gallons) per tonne of  cement) or natural gas is frequently used instead of coal,  but recently some  oil-fired plants have been converted to coal.    Fig. 1.1. Diagrammatic representation of the wet process and the dry process of manufacture of cement     The slurry, in its movement down the kiln, encounters a progressively highe
r  temperature. At first, the water is driven off and CO, is liberated; further on, the dry  material undergoes a series of chemical reactions until finally, in the hottest part of the  kiln, some 20 to 30 per cent of the material becomes liquid, and lime, silica and alumina  recombine. The mass then fuses into balls, 3 to 25 mm (1/8 to 1 in.) in 
diameter, known as  clinker. The clinker drops into coolers, which are of various types and often provide  means for an exchange of heat with the air subsequently used for the combustion of the  pulverized coal. The largest existing kiln in a wet-process plant produces 3600 tonnes of  clinker a day.   The cool clinker, which is characteristically 
black, glistening, and hard, is  interground with gypsum in order to prevent flash-setting of the cement. The grinding is  done in a ball mill consisting of several compartments with progressively smaller steel  balls. In some plants a closed-circuit grinding system is used: the cement discharged by the  mill is passed through a separator, fine pa
rticles being removed to the storage silo by an air  current, while the coarser particles are passed through the mill once again. Closed-circuit  grinding avoids the production of a large amount of excessively fine material or of a small  amount of too coarse material, faults often encountered with open-circuit grinding.   Once the cement has been
 satisfactorily ground, when it will have as many as 1.1 x 10'2  particles per kg (5 x 10'11 per lb), it is ready for packing in the familiar paper bags or in  drums, or for transport in bulk.   In the dry and semi-dry processes, the raw materials are crushed and fed in the  correct proportions into a grinding mill, where they are dried and reduce
d in size to a fine  powder. The dry powder, called raw meal, is then pumped to a blending silo, and final  adjustment is now made in the proportions of the materials required for the manufacture  of cement. To obtain a uniform and intimate mixture the raw meal is blended, usually by  means of compressed air inducing an upward movement of the powd
er and decreasing its  apparent density. The air is pumped over one quadrant of the silo at a time and this  permits the apparently heavier material from the non-aerated quadrants to move laterally  into the aerated quadrant. Thus the aerated material tends to behave almost like a liquid,  and by aerating all quadrants in turn for a total period o
f about one hour a uniform mixture  is obtained. In some cement works continuous blending is used.   In the semi-dry process, the blended meal is now sieved and fed into a rotating  dish called a granulator, water weighing about 12 per cent of the meal being added at the same  time. In this manner hard pellets about 15 mm (1/2 in.) in diameter are
 formed. This is  necessary, as cold powder fed direct into a kiln would not permit the air flow and  exchange of heat necessary for the chemical reactions of formation of cement clinker.   The pellets are baked hard in a pre-heating grate by means of hot gases from the  kiln. The pellets then enter the kiln, and subsequent operations are the same
 as in the wet  process of manufacture. Since, however, the moisture content of the pellets is only 12 per  cent as compared with the 40 per cent moisture content of the slurry used in the wet  process, the dry-process kiln is considerably smaller. The amount of heat required is also  very much lower as only some 12 per cent of moisture has to be 
driven off, but additional  heat has already been used in removing the original moisture content of the raw materials  (usually 6 to 10 per cent). The process is thus quite economical, but only when the raw  materials are comparatively dry. In such a case the total coal consumption can be as little  as 100 kg (220 lb) per tonne of cement.   In the
 dry process, the raw meal, which has a moisture content of about 0.2 per  cent, is passed through a pre-heater, usually of a suspension type. Here it is heated to  about 800  C before being fed into the kiln. Because the raw meal contains no moisture to  be driven off and because it is already pre-heated, the size of the kiln can be much smaller 
 than in the wet process. The major part of the raw meal can be passed through a fluidized  calciner (using a separate heat source) introduced between the pre-heater and the kiln.  This increases the decarbonation (dissociation of CaCO3) of the raw meal prior to entry  into the kiln and greatly increases the kiln throughput. What is probably the l
argest dry-process  plant in the world produces 6200 tonnes of clinker a day using a kiln 6.2 m (20 ft)  in diameter and 105 m (345 ft) long. This output of one kiln is equal to approximately  one-seventh of the cement consumption of the whole of the United Kingdom.   Except when the raw materials necessitate the use of the wet process, the dry  p
rocess is used nowadays in order to minimize the energy required for burning. Typically,  the burning process represents 40 to 60 per cent of the production cost, while the  extraction of raw materials for the manufacture of cement represents only 10 per cent of  the total cost of cement.   The average energy consumption in the United States for t
he production of 1  tonne of cement by the dry process was 1.8 MWh in 1973. Electricity consumption which  accounts for some 6 to 8 per cent of total energy used is typically of the order: 10 kWh for  crushing the raw materials, 28 kWh in the raw mill, 24 kWh in burning, and 41 kWh in  grinding.   There are also other processes of manufacture of c
ement, of which one, using  gypsum instead of lime, perhaps deserves mention. Gypsum, clay and coke with sand and  iron oxide are burnt in a rotary kiln, the end products being Portland cement and sulphur  dioxide which is further converted into sulphuric acid.   In areas where only a small cement production is required, a vertical kiln of the  Go
ttlieb type can be used. This fires nodules of raw meal and fine coal powder combined  and produces agglomerated clinker which is then broken up. A single kiln, 8.5 m (28 ft)  high, produces 150 tonnes of cement a day.   It should be stressed that all processes require an intimate mixture of the raw  materials because a part of the reactions in th
e	kiln must take place by diffusion in solid  materials, and a uniform distribution of materials is essential to ensure a uniform product.      Chemical Composition of Portland Cement    We have seen that the raw materials used in the manufacture of Portland cement consist  mainly of lime, silica, alumina and iron oxide. These compounds interact w
ith one another  in the kiln to form a series of more complex products, and, apart from a small residue of  uncombined lime which has not had sufficient time to react, a state of chemical equilibrium  is reached. However, equilibrium is not maintained during cooling, and the rate of cooling  will affect the degree of crystallization and the amount
 of amorphous material present in  the cooled clinker. The properties of this amorphous material, known as glass, differ  considerably from those of crystalline compounds of a nominally similar chemical  composition. Another complication arises from the interaction of the liquid part of the  clinker with the crystalline compounds already present. 
	Nevertheless, cement can be considered as being in frozen equilibrium, i.e. the  cooled products are assumed to reproduce the equilibrium existing at the clinkering  temperature. This assumption is, in fact, made in the calculation of the compound  composition of commercial cements: the "potential" composition is calculated from the  measured qu
antities of oxides present in the clinker as if full crystallization of equilibrium  products had taken place.   Four compounds are usually regarded as the major constituents of cement: they  are listed in Table 1.1 together with their abbreviated symbols. This shortened notation, used  by cement chemists, describes each oxide by one letter, viz.:
 CaO = C; SiO2, = S; A12O3 = A; and  Fe2O3 = F. Likewise, H2O in hydrated cement is denoted by H.    Table 1.1: Main Compounds of Portland Cement     In reality, the silicates in cement are not pure compounds, but contain minor oxides in solid  solution. These oxides have significant effects on the atomic arrangements, crystal form and hydraulic  
properties of the silicates.   The calculation of the potential composition of Portland cement is based on the  work of R. H. Bogue and others, and is often referred to as "Bogue composition". There  are also other methods of calculating the composition but the subject is not considered to  be within the scope of the present work. We should note, 
however, that the Bogue  composition underestimates the C3S content (and overestimates C2S) because other  oxides replace some of the CaO in C3S.   Bogue's equations for the percentages of main compounds in cement are given below.  The terms in brackets represent the percentage of the given oxide in the total weight of cement.    (equations)     I
n	addition to the main compounds listed in Table 1.1, there exist minor  compounds, such as MgO, TiO2, Mn2O3, K2O and Na2O; they usually  amount to not  more than a few per cent of the weight of cement. Two of  the minor compounds are of  interest: the oxides of sodium and potassium,  Na2O and K2O, known as the alkalis  (although other alkalis als
o	exist in  cement). They have been found to react with some  aggregates, the  products of the reaction causing distintegration of the concrete, and have  also been observed to affect the rate of the gain of strength of cement. It  should,  therefore, be pointed out that the term "minor compounds" refers  primarily to their  quantity and not neces
sarily to their importance. The quantity of the alkalis and of Mn2O3  can be rapidly determined using a  spectrophotometer.   The compound composition of cement has been established largely on the basis of  studies of phase equilibria of the ternary systems C-A-S and C-A-F, and the quarternary  system C-C2S-C5A3-C4AF, and others. The course of mel
ting or crystallization was  traced and the compositions of liquid and solid phases at any temperature were computed.  In addition to the methods of chemical analysis, the actual composition of clinker can be  determined by a microscope examination of powder preparations and their identification  by the measurement of the refractive index. The amo
unts of the two silicates can be  measured by means of a Shands micrometer using a section (similar to those used in  petrographic studies) in transmitted light. Polished and etched sections can also be used both in  reflected and transmitted light. Other methods include the use of X-ray powder diffraction  to identify the crystalline phases and a
lso to study the crystal structure of some of the  phases, and of differential thermal analysis; quantitative analysis is also possible, but  complicated calibrations are involved.  Another development is the use of the electron  microscope, which produces a high magnification by employing an electron beam instead  of light waves.   Estimating the
 composition of cement is aided by new and more rapid methods  of determining the elemental composition, such as X-ray fluorescence and electron probe  micro-analysis. X-ray diffractometry is useful in the determination of free lime, i.e. CaO as  distinct from Ca(OH)2, and this is convenient in controlling the kiln performance.   C3S, which is nor
mally present in the largest amounts, occurs as small,  equidimensional colourless grains. On cooling below 1250 C it decomposes slowly, but, if  cooling is not too slow, C3S remains unchanged and is relatively stable at ordinary temperatures.   C2S is known to have three or possibly even four forms. a-C2S, which exists at  high temperatures, inve
rts to the B-form at about 1450 C. B-C2S undergoes further inversion to  Y-C2S at about 670 C, but at the rates of cooling of commercial cements B-C2S is preserved in  the clinker. B-C2S forms rounded grains, usually showing twinning.   C3A forms rectangular crystals, but C3A in frozen glass forms an amorphous  interstitial phase.   C4AF is really
 a solid solution ranging from C2F to C6A2F, but the description  C4AF is a convenient simplification.   The actual quantities of the various compounds vary considerably from cement to  cement, and indeed different types of cement are obtained by  suitable proportioning of the  materials. In the United States, an attempt was at one time made to co
ntrol the properties  of cements required for different purposes by specifying the limits of the four major  compounds, as  calculated from the oxide analysis. This procedure would cut out  numerous  physical tests normally performed, but unfortunately the calculated compound  composition is not sufficiently accurate, nor does it take into account
 all the relevant  properties of cement, and cannot therefore serve as a substitute for direct testing of the  required properties.   A general idea of the composition of cement can be obtained from Table  1.2,  which gives the oxide composition limits of Portland cements. Table  1.3 gives the oxide  composition of a typical cement and the calcula
ted  compound composition, obtained by  means of Bogue's equations on  page 9.   Two terms used in Table 1.3 require explanation. The insoluble residue,  determined by treating with hydrochloric acid, is a measure of adulteration of cement,  largely arising from impurities in gypsum. BS 12: 1978 limits the insoluble residue to 1.5  per cent of the
 weight of cement. The loss on ignition shows the extent of carbonation and  hydration of free lime and free magnesia due to the exposure of cement to the atmosphere.  The maximum loss on ignition (at 1000 C) permitted by BS 12:1978 is 3 per cent for  cements in a temperate climate and 4 per cent for cements in the tropics. Since hydrated  free li
me is innocuous (see p. 53), for a given free lime content of cement, a greater loss  on ignition is really advantageous.   It is interesting to observe the large influence of a change in the oxide composition  on the compound composition of cement. Some data of Czernin's are given in Table 1.4;  column (1) shows the composition of a fairly typica
l	rapid-hardening cement. If the lime  content is decreased by 3 per cent, with corresponding increases in the other oxides  [column (2)], a considerable change in the C3S : C2S ratio results. Column (3) shows a  change of 1 1/2 per cent in the alumina and iron contents compared with the cement of  column (1). The lime and silica contents are unal
tered and yet the ratio of the silicates, as  well as the contents of C3A and C4AF, is greatly affected. It is apparent that the  significance of the control of the oxide composition of cement cannot be over-emphasized.  Within the usual range of ordinary and rapid-hardening Portland cements the sum of the  contents of the two silicates varies onl
y	within narrow limits, so that the variation in  composition depends largely on the ratio of CaO to SiO2 in the raw materials.    Table 1.2: Approximate Composition Limits of Portland Cement    Table 1.3: Oxide and Compound Compositions of a Typical Portland Cement    Table 1.4: Influence of Change in Oxide Composition on the Compound Composition
		It may be convenient at this stage to summarize the pattern of formation and  hydration of cement; this is shown schemetically in Fig. 1.2.    Fig. 1.2. Schematic representation of the formation and hydration of Portland cement      Hydration of Cement    The reactions by virtue of which Portland cement becomes a bonding agent take place in a
	water-cement paste. In other words, in the presence of water, the silicates and aluminates  listed in Table 1.1 form products of hydration, which in time produce a firm and hard mass  - the hardened cement paste.   There are two ways in which compounds of the type present in cement may react  with water. In the first, a direct addition of some m
olecules of water takes place, this being  a true reaction of hydration. The second type of reaction with water is hydrolysis. It is  convenient and usual, however, to apply the term hydration to all reactions of cement with  water, i.e. to both true hydration and hydrolysis.   Le Chatelier was the first to observe, some hundred years ago, that th
e	products  of  hydration of cement are chemically the same as the products of hydration of the individual  compounds under similar conditions. This was later confirmed by Steinour and by Bogue  and Lerch, with the proviso that the products of reaction may influence one another or  may themselves interact with the other compounds in the system. Th
e	two calcium  silicates are the main cementitious compounds in cement, and the physical behaviour of  cement during hydration is similar to that of these two compounds alone. The hydration of  the individual compounds will be described in more detail in the succeeding sections.   The products of hydration of cement have a low solubility in water 
as shown by  the stability of the hardened cement paste in contact with water. The hydrated cement  bonds firmly to the unreacted cement, but the exact way in which this is achieved is not  certain. It is possible that the newly produced hydrate forms an envelope which grows  from within by the action of water that has penetrated the surrounding f
ilm of hydrate.  Alternatively, the dissolved silicates may pass through the envelope and precipitate as an  outer layer. A third possibility is for the colloidal solution to be precipitated throughout  the mass after the condition of saturation has been reached, the further hydration  continuing within this structure.   Whatever the mode of preci
pitation of the products of hydration, the rate of  hydration decreases continuously, so that even after a long time there remains an  appreciable amount of unhydrated cement. For instance, after 28 days in contact with  water, grains of cement have been found to have hydrated to a depth of only 4 um, and 8  um after a year. Powers calculated that
 complete hydration under normal conditions is  possible only for cement particles smaller than 50 um, but full hydration has been obtained  by grinding cement in water continuously for five days.   Microscopic examination of hydrated cement shows no evidence of channelling of  water into the grains of cement to hydrate selectively the more reacti
ve compounds (e.g.  C3S) which may lie in the centre of the particle. It would seem, then, that hydration  proceeds by a gradual reduction in the size of the cement particle. In fact, unhydrated  grains of coarse cement were found to contain C3S as well as C2S at the age of several  months, and it is probable that small grains of C2S hydrate befor
e	the hydration of large  grains of C3S has been completed. The various compounds in cement are generally  intermixed in all grains, and some investigations have suggested that the residue of a grain  after a given period of hydration has the same percentage composition as the whole of the  original grain. However, the composition of the residue d
oes change throughout the  period of cement hydration, and especially during the first 24 hours selective hydration  may take place.   The main hydrates can be broadly classified as calcium silicate hydrates and  tricalcium aluminate hydrate. C4AF is believed to hydrate into tricalcium aluminate  hydrate and an amorphous phase, probably CaO.Fe2O3.
aq. It is possible also that some  Fe2O3 is present in solid solution in the tricalcium aluminate hydrate.   The progress of hydration of cement can be determined by different means, such  as the measurement of: (a) the amount of Ca(OH)2 in the paste; (b) the heat evolved by  hydration; (c) the specific gravity of the paste; (d) the amount of chem
ically combined  water; (e) the amount of unhydrated cement present (using X-ray quantitative analysis);  and (f) also indirectly from the strength of the hydrated paste. Recently, thermogravimetric  techniques and continuous X-ray diffraction scanning of wet pastes undergoing hydration  have been successfully used in studying early reactions.    
Calcium Silicate Hydrates    The rates of hydration of C3S and C2S in a pure state differ considerably, as shown in Fig. 1.3.    Fig. 1.3. Rate of hydration of pure compounds     In commercial cements, the calcium silicates contain small impurities of some of  the oxides present in the clinker. The "impure" C3S is known as alite and the "impure"  
C2S as belite. These impurities have a strong effect on the properties of the calcium  silicate hydrates (see p. 43).   When hydration takes place in a limited amount of water, as is the case in cement  paste and in concrete, C3S is believed to undergo hydrolysis producing a calcium silicate  of lower basicity, ultimately C3S2H3, with the released
 lime separating out as Ca(OH)2.  There exists, however, a considerable uncertainty as to whether C3S and C2S result  ultimately in the same hydrate. It would appear to be so from considerations of the heat of  hydration and of the surface area of the products of hydration, but physical observations  indicate that there may be more than one - poss
ibly several - distinct calcium silicate  hydrates. The C : S ratio would be affected if some of the lime were absorbed or were held  in solid solution, and there is strong evidence that the ultimate product of hydration of  C2S has a lime / silica ratio of 1.65. This may be due to the fact that the hydration of C3S  is controlled by the rate of d
iffusion of ions through the overlying hydrate films while the  hydration Of C2S is controlled by its slow intrinsic rate of reaction. Furthermore,  temperature may affect the products of hydration of the two silicates as the permeability of  the gel is affected by temperature.   The C : S ratio cannot be reliably determined, different methods yie
lding different  results. Nowadays, the calcium silicate hydrates are generally described as C-S-H, and the  C : S ratio is believed to be probably near 2. Since the crystals formed by hydration are  imperfect and extremely small the mole ratio of water to silica need not be a whole  number. C-S-H usually contains small amounts of Al, Fe, Mg and o
ther ions. At one time,  C-S-H was referred to as tobermorite gel because of a structural similarity to a mineral of  this name, but this may not be correct and this description is now rarely used.   Making the approximate assumption that C3S2H3, is the final product of  hydration of both C3S and C2S, the reactions of hydration can be written (as 
a guide,  although not as exact as stoichiometric equations) as follows:    (reactions).     Thus on a weight basis both silicates require approximately the same amount of  water for their hydration, but C3S produces more than twice as much Ca(OH)2 as is  formed by the hydration of C2S.   The physical properties of the calcium silicate hydrates ar
e of interest in  connection with the setting and hardening properties of cement. These hydrates  appear amorphous but electron microscopy shows their crystalline character. It is  interesting to note that one of the hydrates believed to exist, denoted by Taylor, as  CSH(I), has a layer structure similar to that of some clay minerals, e.g. montmor
illonite  and halloysite. The individual layers in the plane of the a and b axes are well crystallized  while the distances between them are less rigidly defined. Such a lattice would be able to  accommodate varying amounts of lime without fundamental change - a point relevant to  the varying lime / silica ratios mentioned above. In fact, powder d
iagrams have shown that  lime in excess of one molecule per molecule of silica is held in a random manner. Steinour  described this as a merger of solid solution and adsorption.      The use of Calcium 45 tracers has shown that the calcium silicates do not hydrate in the  solid state but the anhydrous silicates probably first pass into solution an
d then react to  form less soluble hydrated silicates which separate out of the supersaturated solution. This  is the type of mechanism of hydration first envisaged by Le Chatelier in 1881.   Recent studies by Diamond indicate that the calcium silicate hydrates  exist in a  variety of forms: fibrous particles, flattened particles, a reticular  net
work, all rather  difficult to define. However, the predominant form is  that of fibrous particles, possibly  solid, possibly hollow, sometimes flattened, sometimes branching at the ends. Typically,  they are 0.5 um to 2 um  long and less than 0.2 um across. This is not a precise picture,  but the  structure of calcium silicate hydrates is too dis
ordered to be established by   the existing techniques, including a combination of the scanning electron  microscope and  energy dispersive X-ray spectrometer.     The hydration of C3S does not proceed at a constant rate. The initial rapid  release of calcium hydroxide into the solution leaves an outer layer of  calcium silicate hydrate,  perhaps 
100 A thick. This is followed by a  so-called dormant period during which very  little hydration takes place, probably due to the deposition of coating on unhydrated  cement grains. Eventually, the coating ruptures because of pressure of the products of  hydration beneath it, and hydration speeds up again. Finally, there is a  further slowing  dow
n when diffusion through the pores in the products of  hydration becomes the  controlling factor.   It is interesting to observe that calcium silicate hydrates show a strength  development similar to that of Portland cement. A considerable strength is possessed long  before the reaction of hydration is complete and it would thus seem that a small 
amount  of the hydrate binds together the unhydrated remainder; further hydration results in little  or no increase in strength.   Ca(OH)2 liberated by the hydrolysis of the calcium silicates forms thin hexagonal  plates, often tens of um across, but later they merge into a massive deposit.    Tricalcium Aluminate Hydrate and the Action of Gypsum 
	The amount of C3A present in most cements is comparatively small but its behaviour and  structural relationship with the other phases in cement make it of interest. The tricalcium  aluminate hydrate forms a prismatic dark interstitial material, possibly with other  substances in solid solution, and is often in the form of flat plates individual
ly surrounded  by the calcium silicate hydrates.   The reaction of pure C3A with water is very violent and leads to immediate  stiffening of the paste, known as flash set. To prevent this from happening, gypsum  (CaSO4.2H2O) is added to cement clinker. Gypsum and C3A react to form insoluble  calcium sulphoaluminate (3CaO.Al2O3.3CaSO4.31H2O), but e
ventually a tricalcium  aluminate hydrate is formed, although this is preceded by a metastable  3CaO.Al2O3.CaSO4.12H2O, produced at the expense of the original high-sulphate  calcium sulphoaluminate. As more C3A comes into solution, the composition changes, the  sulphate content decreasing continuously. The rate of reaction of the aluminate is hig
h	and, if this readjustment in composition is not rapid enough, direct hydration of C3A is  likely. In particular, a peak in the rate of heat development normally observed within five  minutes of adding water to cement means that some calcium aluminate hydrate is formed  directly during that period, the conditions for the retardation by gypsum no
t yet having  been established.   There is some evidence that the hydration of C3A is retarded by Ca(OH)2, liberated  by the hydrolysis of C3S. This occurs due to the fact CA(OH)2 reacts with C3A and water  to form C4AH19 which forms a protective coating on the surface of unhydrated grains of  C3A. It is also possible that Ca(OH)2 decreases the co
ncentration of aluminate ions in the  solution, thus slowing down the rate of hydration of C3A.   The stable form of the calcium aluminate hydrate ultimately existing in the  hydrated cement paste is probably the cubic crystal C4AH6, but it is possible that  hexagonal C4AH12 crystallizes out first and later changes to the cubic form. Thus the fina
l	form of the reaction can be written -    (reaction).    (This again is an approximation and not a stoichiometric equation.)   The molecular weights show that 100 parts of C3A react with 40 parts of water  by weight, which is a much higher proportion of water than that required by the silicates.   The presence of C3A in cement is undesirable: it
 contributes little or nothing to  the strength of cement except at early ages, and when hardened cement paste is attacked by  sulphates, expansion due to the formation of calcium sulphoaluminate from C3A may  result in a disruption of the hardened paste. However, C3A acts as a flux and thus reduces  the temperature of burning of clinker and facil
itates the combination of lime and silica; for  these reasons C3A is useful in the manufacture of cement. C4AF also acts as a flux. It may  be noted that if some liquid were not formed during burning, the reactions in the kiln  would progress much more slowly and would probably be incomplete.   Gypsum reacts not only with C3A: with C4AF it forms c
alcium sulphoferrite as  well as calcium sulphoaluminate, and its presence may accelerate the hydration of the  silicates.   The amount of gypsum added to the cement clinker has to be very carefully  watched; in particular, an excess of gypsum leads to an expansion and consequent  disruption of the set cement paste. The optimum gypsum content is d
etermined by  observation of the generation of the heat of hydration. Generally, the immediate peak in  the rate of heat evolution is followed by a second peak some 4 to 8 hours after the water  has been added to cement, and with the correct amount of gypsum there should be little  C3A available for reaction after all the gypsum has combined, and 
no further peak in the  heat liberation should occur. Thus, an optimum gypsum content leads to a desirable rate of  early reaction and prevents local high concentration of products of hydration (see p. 46).   The amount of gypsum required increases with the C3A content and also with the  alkali content of the cement. Increasing the fineness of cem
ent has the effect of increasing  the quantity of C3A available at early ages, and this raises the gypsum requirement.   The amount of gypsum added to cement clinker is usually expressed as the weight  of SO3 present; this is limited by BS 12: 1978 to a maximum of 2.5 per cent when the C3A  content is not more than 5 per cent, and to 3.0 per cent 
when the amount of C3A exceeds  5 per cent.      Setting    This is the term used to describe the stiffening of the cement paste, although the definition  of the stiffness of the paste which is considered set is somewhat arbitrary. Broadly  speaking, setting refers to a change from a fluid to a rigid state. Although during setting  the paste acqui
res some strength, for practical purposes it is convenient to distinguish  setting from hardening, which refers to the gain of strength of a set cement paste.   In practice, the terms initial set and final set are used to describe arbitrarily chosen  stages of setting. The method of measurement of these setting times is described on page  50.   It
 seems that setting is caused by a selective hydration of cement compounds: the  two first to react are C3A and C3S. The flash-setting properties of the former were  mentioned in the preceding section but the addition of gypsum delays the formation of  calcium aluminate hydrate and it is thus C3S that sets first. Neat C3S mixed with water  also ex
hibits an initial set but C2S stiffens in a more gradual manner.   In a properly retarded cement the framework of the hydrated cement paste is  established by the calcium silicate hydrate, while if C3A were allowed to set first a rather  porous calcium aluminate hydrate would form. The remaining cement compounds would  then hydrate within this por
ous framework and the strength characteristics of the cement  paste would be adversely affected.   Apart from the rapidity of formation of crystalline products, the development of  films around cement grains and a mutual coagulation of components of the paste have also  been suggested as factors in the development of set.   The setting process is 
accompanied by temperature changes in the cement paste:  initial set corresponds to a rapid rise in temperature and final set to the peak temperature.  At this time a sharp drop in electrical conductivity also takes place, and attempts have  been made to measure setting by electrical means.   The setting time of cement decreases with a rise in tem
perature, but above about  30 C (85 F) a reverse effect may be observed. At low temperatures setting is retarded. The  influence of temperature on setting time is indicated in Fig. 7.7.    False Set    False set is the name give to the abnormal premature stiffening of cement within a few  minutes of mixing with water. It differs from flash set in 
that no appreciable heat is  evolved, and remixing the cement paste without addition of water restores plasticity of the  paste until it sets in the normal manner and without a loss of strength.   Some of the causes of false set are to be found in the dehydration of gypsum  when interground with too hot a clinker: hemihydrate (CaSO4.1/2H2O) or any
hdrite (CaSO4)  are formed, and when the cement is mixed with water these hydrate to form gypsum. Thus  plaster set takes place with a resulting stiffening of the paste.   Another cause of false set may be associated with the alkalis in the cement.  During storage they may carbonate, and alkali carbonates react with Ca(OH)2 liberated by the  hydro
lysis of C3S to form CaCO3. This precipitates and induces a rigidity of the paste.   It has also been suggested that false set can be due to the activation of C3S by  aeration at moderately high humidities. Water is adsorbed on the grains of cement, and  these freshly activated surfaces can combine very rapidly with more water during mixing:  this
 rapid hydration would produce false set.   Laboratory tests at cement works generally ensure that cement is free from false  set. If, however, false set is encountered, it can be dealt with by remixing the concrete  without adding any water. Although this is not easy, workability will be improved and the  concrete can be placed in the normal mann
er.      Fineness of Cement    It may be recalled that one of the last steps in the manufacture of cement is the grinding of  clinker mixed with gypsum. Since the hydration starts at the surface of the cement  particles, it is the total surface area of cement that represents the material available for  hydration. Thus the rate of hydration depends
 in the fineness of the cement particles, and  for a rapid development of strength high fineness is necessary (see Fig. 1.4).    Fig. 1.4. Relation between strength of concrete at different ages and fineness of cement     On the other hand, the cost of grinding to a higher fineness is considerable, and  also the finer the cement the more rapidly i
t deteriorates on exposure to the atmosphere.  Finer cement leads to a stronger reaction with alkali-reactive aggregate, and makes a  paste, though not necessarily concrete, exhibiting a higher shrinkage and a greater  proneness to cracking. However, fine cement bleeds less than a coarser one.   An increase in fineness increases the amount of gyps
um required for proper  retardation as in a finer cement more C3A is available for early hydration. The water  content of a paste of standard consistence is greater the finer the cement, but conversely  an increase in fineness of cement slightly improves the workability of a concrete mix. This  anomaly may be due partly to the fact that the tests 
for consistence of cement paste and  workability measure different properties of fresh paste; also, accidental air affects the  workability of cement paste, and cements of different fineness may contain different  amounts of air.   We can see then that fineness is a vital property of cement and has to be carefully  controlled. In the past, the fra
ction of cement retained on a 90 um (No. 170 ASTM) test  sieve was determined, and the maximum residue was limited to 10 per cent by weight for  ordinary and 5 per cent for rapid hardening Portland cement. A cement satisfying these  conditions would not contain an excess of large grains which, because of their  comparatively small surface are per 
unit weight, would play only a small role in the  process of hydration and development of strength.   However, the sieve test yields no information on the size of grains smaller than 90  um (No. 170 ASTM) sieve, and it is the finer particles that play the greatest part in the  early hydration. Attempts to use smaller sieves, down to 53 um (No. 270
 ASTM), have  generally been unsuccessful because of clogging of such extremely fine mesh.   For these reasons, BS 4550: Part 3: Section 3.3: 1978 prescribes a test for  fineness by determination of the specific surface of cement expressed as the total surface  area in square metres per kilogram. A direct approach is to measure the particle size d
istribution  by sedimentation or elutriation: these methods are based on the dependence of the rate of  free fall of particles on their diameter. Stoke's law gives the terminal velocity of fall under  gravity of a spherical particle in a fluid medium. This medium must of course be  chemically inert with respect to cement. It is also important to a
chieve a satisfactory  dispersion of cement particles as partial flocculation would produce a decrease in the  apparent specific surface.   A development of these methods is the Wagner turbidimeter used in the United  States (ASTM Standard C 115-79a). In this test, the concentration of particles in  suspension at a given level in kerosene is deter
mined using a beam of light, the percentage  of light transmitted being measured by a photocell. The turbidimeter gives generally  consistent results, but an error is introduced by assuming a uniform size distribution of  particles smaller than 75 um. It is precisely these finest particles that contribute most to  the specific surface of cement an
d	the error is especially significant with the finer cements  used nowadays. However, an improvement on the standard method is possible if the  concentration of particles 5 um in size is determined and a modification of calculations is  made. A typical curve of particle size distribution is shown in Fig. 1.5, which gives also  the corresponding co
ntribution of these particles to the total surface area of the sample.  As mentioned on page 6, the particle size distribution depends on the method of grinding  and varies, therefore, from plant to plant.    Fig. 1.5. Example of particle size distribution and cumulative surface area contributed by  particles up to any size for 1 gram of cement   
	It must be admitted, however, that it is not quite clear what is a "good" grading  of cement: should all the particles be of the same size or should their distribution be such  that they are able to pack densely? Data on the influence of particle size on the rate of  penetration of water are confusing.   A more recent method of determination of 
the specific surface of cement  is the air permeability method, using an apparatus developed by Lea and Nurse. This is the  method of measurement prescribed by BS 4550: Part 3: Section 3.3: 1978. It is based on  the relation between the flow of a fluid through a granular bed and the surface area of the  particles comprising the bed. From this the 
surface area per unit weight of the bed material  can be related to the permeability of a bed of a given porosity, i.e. containing a fixed  volume of pores in the total volume of the bed.    Fig. 1.6. Lea and Nurse permeability apparatus     The permeability apparatus is shown diagrammatically in Fig. 1.6. Knowing the  density of cement, the weigh
t	required to make a bed of porosity of 0.475 and 1 cm thick  can be calculated. This amount of cement is placed in a cylindrical container, a stream of  dry air is passed through the cement bed at a constant velocity and the resulting pressure  drop is measured by a manometer connected to the top and bottom of the bed. The rate of  airflow is mea
sured by a flowmeter consisting of a capillary placed in the circuit and a  manometer across its ends.   An equation developed by Carman gives the specific surface in square centimetres  per gram as    (formula)    where Q  = density of cement (g / cm3)   E = porosity of cement bed (0.475 in the BS test)   A = cross-sectional area of the bed (5.06
6	cm2)   L  = height of the bed (1 cm)   h1 = pressure drop across the bed   h2 = pressure drop across the flowmeter capillary (between 25 and     55 cm of kerosene)  and K = the flowmeter constant.    For a given apparatus and porosity the expression simplifies to    (formula)  where K1 is a constant.   In the United States and in Germany, a modi
fication of the Lea and Nurse  method, developed by Blaine, is used. Here, the air does not pass through the bed at a  constant rate but a known volume of air passes at a  prescribed average pressure, the rate  of flow diminishing steadily. The time t for the flow to take place is measured, and for a  given apparatus and a standard porosity of 0.5
00, the specific surface is given by -    (formula)  where K2 is a constant.   The Lea and Nurse, and Blaine methods give values of specific surface in close  agreement with one another but very much higher than the value obtained by the Wagner  method. This is due to Wagner's assumptions about the size distribution of particles below  7.5 um, men
tioned earlier. The actual distribution in this range is such that the average  value of 3.75 um, assumed by Wagner, underestimates the surface area of these particles.  In the air permeability method, the surface area of all particles is measured directly, and the  resulting value of the specific surface is about 1.8 times higher than the value c
alculated by  the Wagner method. The actual range of the conversion factor varies between 1.6 and 2.2,  depending on the fineness of the cement and its gypsum content.   Either method gives a good picture of the relative variation in the fineness of  cement, and for practical purposes this is sufficient. The Wagner method is somewhat  more informa
tive in that it gives an indication of the particle size distribution. An absolute  measurement of the specific surface can be obtained by the nitrogen adsorption method,  based on the work of Brunauer, Emmett and Teller.  While in the air permeability methods  only the continuous paths through the bed of cement contribute to the measured area, in
	the nitrogen adsorption method the "internal" area is also accessible to the nitrogen  molecules. For this reason, the measured value of the specific surface is considerably  higher than that determined by the air permeability methods. Some typical values are given  in Table 1.5.    Table 1.5: Specific Surface of Cement Measured by Different Met
hods     BS 12: 1978 lays down the minimum specific surface (determined by the Lea and  Nurse method) as 225 m2 / kg for ordinary Portland cement, and 325 m2 / kg for rapid  hardening Portland cement. Other minimum requirements are 225 m2 / kg for Portland  blast-furnace cement (BS 146: 1973) and 275 m2 / kg for low-heat Portland cement (BS  1370:
 1979).   During the last twenty years, there has been a tendency to grind cement more  finely, and the commercial ordinary Portland cement produced in Britain is generally finer  than the minimum laid down by BS 12: 1978: about 300 m2 / kg is a typical value.        High-alumina cement is normally coarser than Portland cements. A minimum of  225 
m2 / kg is specified by BS 915: 1972, but slightly higher values are generally  encountered in practice.   In some countries, e.g. Switzerland, fineness of cement is not prescribed in  standards but is indirectly controlled by tests on early strength.      Structure of Hydrated Cement    Many of the mechanical properties of hardened cement and con
crete appear to depend not  so much on the chemical composition of the hydrated cement as on the physical structure  of the products of hydration, viewed at the level of colloidal dimensions. For this reason it  is important to have a good picture of the physical properties of the cement gel.   Fresh cement paste is a plastic network of particles 
of cement in water but, once  the paste has set, its apparent or gross volume remains approximately constant. At any  stage of hydration, the hardened paste consists of hydrates of the various compounds,  referred to collectively as gel, of crystals of Ca(OH)2, some minor components,  unhydrated cement, and the residue of the water-filled spaces i
n the fresh paste. These  voids are called capillary pores, but within the gel itself there exist interstitial voids, called  gel pores. There are thus, in hydrated paste, two distinct classes of pores represented  diagrammatically in Fig. 1.7.    Fig. 1.7. Simplified model of paste structure     Since most of the products of hydration are colloid
al (the ratio of calcium silicate  hydrates to CA(OH)2 being 7:2 by weight) during hydration the surface area of the solid  phase increases enormously, and a large amount of free water becomes adsorbed on this  surface. If no water movement to or from the cement paste is permitted, the reactions of  hydration use up the water until too little is l
eft to saturate the solid surfaces, and the  relative humidity within the paste decreases. This is known as self-desiccation. Since gel  can form only in water-filled space, self-desiccation leads to a lower hydration compared  with a moist-cured paste. However, in self-desiccated pastes with water / cement ratios in  excess of 0.5, the amount of 
mixing water is sufficient for hydration to proceed at the  same rate as when moist-cured.      Volume of Products of Hydration    The gross space available for the products of hydration consists of the absolute volume of  the dry cement together with the volume of water added to the mix. The small loss of  water due to bleeding and the contractio
n of the paste while still plastic will be ignored at  this stage. The water bound chemically by C3S and C2S was shown to be very  approximately 24 and 21 per cent of the weight of the two silicates, respectively. The  corresponding figures for C3A and C4AF are 40 and 37 per cent, respectively.   As mentioned earlier, these figures are not accurat
e as our knowledge of  stoichiometry of the products of hydration of cement is inadequate to state the amounts of  water combined chemically. It is preferable, therefore, to consider non-evaporable water  as determined by a given method (see p. 37). This water, determined under specified  conditions, is taken as 23 per cent of the weight of anhydr
ous cement (although in Type  II cement the value may be as low as 18 per cent).   The specific gravity of the products of hydration of cement is such that they  occupy a greater volume than the absolute volume of unhydrated cement but smaller than  the sum of volumes of the dry cement and the non-evaporable water by approximately  0.254 of the vo
lume of the latter. An average value of specific gravity of the products of  hydration (including pores in the densest structure possible) in a saturated state is 2.16.   As an example, let us consider the hydration of 100 g of cement. Taking the  specific gravity of dry cement as 3.15, the absolute volume of unhydrated cement is  100/3.15 = 31.8 
ml. The non-evaporable water is, as we have said, about 23 per cent of  the weight of cement, i.e. 23 ml. The solid products of hydration occupy a volume equal  to the sum of volumes of anhydrous cement and water less 0.254 of the volume of  non-evaporable water, i.e. -    (formula).    Since the paste in this condition has a characteristic porosi
ty of about 28 per cent, the  volume of gel water, wg, is given by    (formula)  whence wg = 19.0 ml,  and the volume of hydrated cement is (formula).    Summarizing, we have -  Weight of dry cement      =  100.0 g  Absolute volume of dry cement    =  31.8 ml  Weight of combined water     =  23.0 g  Volume of gel water     =  19.0 ml  Total water 
in the mix      =  42.0 ml  Water / cement ratio by weight    =  0.42  Water / cement ratio by volume     =  1.32  Volume of hydrated cement    =  67.9 ml  Original volume of cement and water    =  73.8 m1  Decrease in volume due to hydration    =  5.9 ml  Volume of products of hydration of 1 ml of dry cement =  2.l ml.     It should be noted that
 the hydration was assumed to take place in a sealed test  tube with no water movement to or from the system. The volumetric changes are shown in  Fig. 1.8. The "decrease in volume" of 5.9 ml represents the empty capillary space  distributed throughout the hydrated paste.    Fig. 1.8. Diagrammatic representation of volume changes due to hydration 
of paste with a  water / cement ratio of 0.42     The figures given above are only approximate but had the total amount of water  been lower than about 42 ml, it would have been inadequate for full hydration as gel can  form only when sufficient water is available both for the chemical reactions and for the  filling of the gel pores being formed. 
The gel water, because it is held firmly, cannot move  into the capillaries so that it is not available for hydration of the still unhydrated cement.   Thus, when hydration in a sealed specimen has progressed to a stage when the  combined water has become about one-half of the original water content, no further  hydration will take place. It follo
ws also that full hydration in a sealed specimen is possible  only when the mixing water is at least twice the water required for chemical reaction, i.e.  the mix has a water / cement ratio of about 0.5 by weight. In practice, in the example given  above, the hydration would not in fact have progressed to completion as hydration stops  even before
 the capillaries have become empty. It has been found that hydration becomes  very slow when the water vapour pressure falls below about 0.8 of the saturation  pressure.   Let us know consider the hydration of a paste cured under water so that water  can be imbibed as some of the capillaries become emptied by hydration. As shown before,  200 g of 
cement (31.8 ml) will on full hydration occupy 67.9 ml. Thus, for no unhydrated  cement to be left and no capillary pores to be present, the original mixing water should be  approximately (67.9 - 31.8) = 36.1 ml. This corresponds to a water / cement ratio of 1.14  by volume or 0.36 by weight. From other work, values of about 1.2 and 0.38 respectiv
ely  have been suggested.	If the actual water / cement ratio of the mix, allowing for bleeding, is less than  about 0.38 by weight, complete hydration is not possible as the volume available is insufficient to  accommodate all the products of hydration. It will be recalled that hydration can take  place only in water within the capillaries. For 
instance, if we have a mix of 100 g of cement  (31.8 ml) and 30 g of water, the water would suffice to hydrate x g of cement, given by the  following:    Contraction in volume on hydration =  (formula).  Volume occupied by solid products of hydration = (formula).  Porosity = (formula)  and total water = (formula).  Hence x = 71.5  g = 22.7 ml  and
	wg = 13.5 g.  Thus, the volume of hydrated cement = (formula).    The volume of unhydrated cement = (formula).  Therefore, the volume of empty capillaries = (formula).     If water is available from outside some further cement can hydrate, its quantity  being such that the products of hydration occupy 4.2 ml more than the volume of dry  cement. 
We found that 22.7 ml of cement hydrates to occupy 48.5 ml, i.e. the products of  hydration of 1 ml of cement occupy 48.5/22.7 = 2.13 ml. Thus 4.2 ml would be filled by  the hydration of y ml of cement such that (4.2 + y) / y = 2.13; hence y = 3.7 ml. Thus the  volume of still unhydrated cement is 31.8 - (22.7 + 3.7) = 5.4 ml, which weighs 17 g. I
n	 other words, 19 per cent of the original weight of cement has remained unhydrated and  can never hydrate since the gel already occupies all the space available, i.e. the gel / space  ratio (see p. 275) of the hydrated paste is 1.0.   It may be added that unhydrated cement is not detrimental to strength and, in  fact, among pastes all with a gel
 / space ratio of 1.0 those with a higher proportion of unhydrated  cement (i.e. a lower water / cement ratio) have a higher strength, possibly because in such pastes the layers of hydrated paste surrounding the unhydrated grains are thinner. Abrams  obtained strengths of the order of 280 MPa (40 000 psi) using mixes with a water / cement  ratio o
f	0.08 by weight, but clearly considerable pressure is necessary to obtain a properly  consolidated mix of such proportions. More recently, Lawrence made compacts of cement  powder in a die assembly under a very high pressure (up to 672 MPa (or 97 500 psi)), using  the techniques of powder metallurgy. Upon subsequent hydration for 28 days,  compre
ssive strengths up to 375 MPa (or 54 500 psi) and tensile strengths up to 25 MPa  (or 3600 psi) were measured. The porosity of such mixes and therefore the "equivalent"  water / cement ratio are very low.   On the other hand, if the water / cement ratio is higher than about 0.38, all the  cement can hydrate but capillary pores will also be present
.	Some of the capillaries will  contain excess water from the mix, others will fill by imbibing water from outside.  Fig. 1.9 shows the relative volumes of unhydrated cement, products of hydration,  and capillaries for mixes of different water / cement ratios.    Fig. 1.9. Composition of cement paste at different stages of hydration     As a more 
specific example, let us consider the hydration of a paste with a  water / cement ratio of 0.475, sealed in a tube. Let the weight of dry cement be 126 g,  which corresponds to 40 ml. The volume of water is then 0.475 x 126 = 60 ml. These mix  proportions are shown in the left-hand diagram of Fig. 1.10, but in reality the cement and  water are of 
course intermixed, the water forming a capillary system between the  unhydrated cement particles.    Fig. 1.10. Diagrammatic representation of the volumetric proportions of cement paste  at different stages of hydration     Let us now consider the situation when the cement has hydrated fully. The  non-evaporable water is 0.23 x 126 = 29.0 ml and t
he gel water is wg, such that -    (formula),    whence, the volume of gel water is 24.0 ml, and the volume of hydrated cement is 85.6 ml.  There are thus 60 - (29.0 + 24.0) = 7.0 ml of water left as capillary water in the paste. In  addition, 100 - (85.6 + 7.0)= 7.4 ml form empty capillaries. If the paste had access to water  during curing these 
capillaries would fill with imbibed water.   This then is the situation at 100 per cent hydration when the gel / space ratio is  0.856, as shown in the right-hand diagram of Fig. 1.10. As a further illustration, the centre  diagram shows the volumes of different components when only half the cement has  hydrated. The gel / space ratio is then    (
formula).    Capillary Pores    We can thus see that at any stage of hydration the capillary pores represent that part of the  gross volume which has not been filled by the products of hydration. Since these products  occupy more than twice the volume of the original solid phase (i.e. cement) alone, the  volume of the capillary system is reduced w
ith the progress of hydration.   Thus the capillary porosity of the paste depends on both the water / cement ratio of  the mix and the degree of hydration. The rate of hydration of the cement is of no  importance per se, but the type of cement influences the degree of hydration achieved at a  given age. As mentioned before, at water / cement ratio
s	higher than about 0.38 the volume  of the gel is not sufficient to fill all the space available to it so that there will be some  volume of capillary pores left even after the process of hydration has been completed.   Capillary pores cannot be viewed directly but their size has been estimated from  vapour pressure measurement to be of the order
 of 1.3 um (5x10'-5 in.). They vary in shape  but, as shown by measurement of permeability, form an interconnected system randomly  distributed throughout the cement paste. These interconnected capillary pores are mainly  responsible for the permeability of the hardened cement paste and for its vulnerability to  frost.   However, hydration increas
es the solid content of the paste and in mature and  dense pastes the capillaries may become blocked by gel and segmented so that they turn  into capillary pores interconnected solely by the gel pores. The absence of continuous  capillaries is due to a combination of a suitable water / cement ratio and a sufficiently long  period of moist curing; 
the degree of maturity required for different water / cement ratios  for ordinary Portland cements is indicated in Fig. 1.11. The actual time to achieve the  required maturity depends on the characteristics of the cement used, but approximate  values of the time required can be gauged from the data of Table 1.6. For water / cement  ratios above ab
out 0.7 even complete hydration would not produce enough gel to block  all the capillaries. For extremely fine cement, the maximum water / cement ratio would be  higher, possibly up to 1.0; conversely, for coarse cements, it would be below 0.7. The  importance of eliminating continuous capillaries is such that this might be regarded a  necessary c
ondition for a concrete to be classified as "good".    Fig. 1.11. Relation between the water / cement ratio and the degree of hydration at which  the capillaries cease to be continuous    Table 1.6: Approximate Age Required to produce Maturity at which Capillaries become segmented      Gel Pores    Let us now consider the gel itself. From the fact
 that it can hold large quantities of  evaporable water it follows that the gel is porous, but the gel pores are really  interconnected interstitial spaces between the gel particles.    The gel pores are much smaller than the capillary pores: between 15 and  20 A in diameter. This is only one order of magnitude greater than the size of  molecules 
of water. For this reason, the vapour pressure and mobility of adsorbed  water are different from the corresponding properties of free water. The amount  of reversible water indicates directly the porosity of the gel.   The gel pores occupy about 28 per cent of the total volume of gel, the  material left after drying in a standard manner being con
sidered as solids. The  actual value is characteristic for a given cement but is largely independent of  the water / cement ratio of the mix and of the progress of hydration. This would  indicate that gel of similar properties is formed at all stages and that  continued hydration does not affect the products already in existence. Thus as  the tota
l	volume of gel increases with the progress of hydration the total volume  of gel pores also increases. On the other hand, as mentioned earlier, the volume  of capillary pores decreases with the progress of hydration.   Porosity of 28 per cent means that the gel pores occupy a space equal to  about one-third of the volume of the gel solids. The ra
tio of the surface of the  solid part of the gel to the volume of the solids is equal to that of spheres  about 90 A in diameter. This must not be construed to mean that gel consists of  spherical elements; the particles are mostly fibrous, and bundles of such fibres  form a cross-linked network containing some more or less amorphous interstitial 
 material.   Another way of expressing the porosity of the gel is to say that the volume  of the pores is about three times the volume of the water providing a layer one  molecule thick over the entire solid surface in the gel.   From measurements of water adsorption, the specific surface of gel has been  estimated to be of the order of 5.5 x 10'8
 m2 per m3, or approximately  200 000 m2 / kg. Recent low-angle X-ray scattering measurements have yielded values  of the order of 600 000 m2 / kg, indicating a large internal surface within the  particles. By contrast, unhydrated cement has a specific surface of some 200 to  500 m2 / kg.   In connection with the pore structure, it may be relevant
 to note that  high-pressure steam-cured cement has a specific surface of some 7000 m2 / kg only.  This indicates an entirely different particle size of the products of hydration  at a high pressure and temperature and, in fact, steam curing seems to result in  an almost entirely micro-crystalline material.   The specific surface of normally cured
 cement depends on the curing  temperature and on the chemical composition of cement. It has been suggested  that the ratio of the specific surface to the weight of non-evaporable water  (which in turn is proportional to the porosity of the hydrated cement paste) is  proportional to    (formula),    where the symbols in brackets represent the perc
entages of the compounds present  in the cement. There seems to be little variation between the numerical  coefficients of the last three compounds, and this indicates that the specific  surface of the paste varies little with a change in the composition of cement.  The rather lower coefficient of C3S is due to the fact that it produces a large  q
uantity of micro-crystalline Ca(OH)2, which has a very much lower specific  surface than the gel.   The proportionality between the weight of water forming a monomolecular  layer over the surface of the gel and the weight of non-evaporable water in the  paste (for a given cement) means that gel of nearly the same specific surface is  formed throug
hout the progress of hydration. In other words, particles of the  same size are formed all the time and the already existing gel particles do not  grow in size. This is not, however, the case in cement with a high C2S content.      Mechanical Strength of Cement Gel    There are two classical theories of hardening or gain of strength of cement. Tha
t	 put forward by H. Le Chatelier in 1882 states that the products of hydration of  cement have a lower solubility than the original compounds, so that the hydrates  precipitate from a supersaturated solution. The precipitate is in the form of  interlaced elongated crystals with high cohesive and adhesive properties.   The colloidal theory propoun
ded by W. Michaelis in 1893 states that the  crystalline aluminate, sulpho-aluminate and hydroxide of calcium give the initial  strength. The lime-saturated water then attacks the silicates and forms a  hydrated calcium silicate which, being almost insoluble, forms a gelatinous mass.  This mass hardens gradually owing to the loss of water either b
y	external drying  or by hydration of the inner unhydrated core of the cement grains: in this manner  cohesion is obtained.   In the light of modern knowledge it appears that both theories contain  elements of truth and are in fact by no means irreconcilable. In particular,  colloid chemists have found that many if not most colloids consist of cry
stalline  particles but these, being extremely small, have a large surface area which gives  them what appear to be different properties from the usual solids. Thus colloidal  behaviour is essentially a function of the size of the surface area rather than  of the non-regularity of internal structure of the particles involved.   In the case of Port
land cement, it has been found that, when mixed with a  large quantity of water, cement produces within a few hours a solution  supersaturated with Ca(OH)2 and containing concentrations of calcium silicate  hydrate in a metastable condition. This hydrate rapidly precipitates in agreement  with Le Chatelier's theory; the subsequent hardening may be
 due to the withdrawal  of water from the hydrated material as postulated by Michaelis.   Further experimental work has shown that the calcium silicate hydrates are in fact  in the form of extremely small (sub-microscopic) interlocking crystals which, because of  their size, could be equally well described as gel. When cement is mixed with a small
	quantity of water the degree of crystallization is probably even poorer, the crystals being  ill-formed. Thus the Le Chatelier-Michaelis controversy is largely reduced to a matter of  terminology as we are dealing with a gel consisting of crystals.   The term cement gel is considered, for convenience, to include the crystalline  calcium hydroxid
e. Gel is thus taken to mean the cohesive mass of hydrated cement in its  densest paste, i.e. inclusive of gel pores, the characteristic porosity being about 28 per cent.   The actual source of strength of the gel is not fully understood but it probably arises  from two kinds of cohesive bonds. The first type is the physical attraction between sol
id surfaces,  separated only by the small (15 to 20 A) gel pores; this attraction is usually referred to as van der Waals'  forces.   The source of the second type of cohesion is the chemical bonds. Since cement gel  is of the limited swelling type (i.e. the particles cannot be dispersed by addition of water)  it seems that the gel particles are c
ross-linked by chemical forces. These are much stronger  than van der Waals' forces but the chemical bonds cover only a small fraction of the  boundary of the gel particles. On the other hand, a surface area as high as that of cement  gel is not a necessary condition for high strength development, as high-pressure steam-cured  cement paste, which 
has a low surface area, exhibits extremely good hydraulic properties.   We cannot thus estimate the relative importance of the physical and chemical  bonds but there is no doubt that both contribute to the very  considerable strength of the  hardened paste.      Water Held in Hydrated Cement Paste    The presence of water in hydrated cement has be
en repeatedly mentioned. The cement  paste is indeed hygroscopic owing to the hydrophilic character of cement coupled with the  presence of sub-microscopic pores. The actual water content of the paste depends on the  ambient humidity. In particular, capillary pores, because of  their comparatively large size,  empty when the ambient relative humid
ity falls below about 45 per cent, but water is  adsorbed in the gel pores even at very low ambient humidities.   We can thus see that water in hydrated cement is held with varying degree  of firmness. At one extreme there is free water; at the other, chemically combined  water forming a definite part of the hydrated compounds. Between these two c
ategories  there is gel water held in a variety of other ways.   The water held by the surface forces of the gel particles is called adsorbed water,  and that part of it which is held between the surfaces of certain planes in a crystal is called  interlayer or zeolitic water. Lattice water is that part of the water of crystallization which  is not
 chemically associated with the principal constituents of the lattice. The diagrammatic  representation of Fig. 1.12 may be of interest.    Fig. 1.12 Probable structure of hydrated silicates     Free water is held in capillaries and is beyond the range of the surface forces of  the solid phase.   There is no technique available for determining how
 water is distributed between  these different states, nor it is easy to predict these divisions from theoretical  considerations as the energy of binding of combined water in the hydrate is of the same  order of magnitude as the energy of binding of the adsorbed water. However, recent  investigations using the nuclear magnetic resonance technique
 suggest that gel water has  the same energy of binding as interlayer water in some swelling clays; thus the gel water  may well be in interlayer form.   A convenient division of water in the hydrated cement, necessary for investigation  purposes, though rather arbitrary, is into two categories: evaporable and non-evaporable.  This is achieved by 
drying the cement paste to equilibrium (i.e. to a constant weight) at a  given vapour pressure. The usual value is 1 Pa at 23 C, obtained over Mg(ClO4)2 .2H2O.  More recently, drying in an evacuated space which is connected to a moisture trap held at  -79 C has been used; this corresponds to a vapour pressure of 0.07 Pa. Alternatively, the  evapor
able water can be determined by the loss upon drying at a higher temperature,  usually 105 C, or by freezing out, or by removing with a solvent.   All these methods essentially divide water according to whether or not it can be  removed at a certain reduced vapour pressure. Such a division is perforce arbitrary because  the relation between vapour
 pressure and water content of cement is continuous. By  contrast with crystalline hydrates, no breaks occur in this relationship. However, in  general terms, the non-evaporable water contains nearly all chemically combined water  and also some water not held by chemical bonds. This water has a vapour pressure lower  than that of the ambient atmos
phere and the quantity of such water is in fact a continuous  function of the ambient vapour pressure.   The amount of non-evaporable water increases as hydration proceeds but in a  saturated paste non-evaporable water can never become more than one-half of the total  water present. In well-hydrated cement the non-evaporable water is about 18 per 
cent by  weight of the anhydrous material; this proportion rises to about 23 per cent in fully  hydrated cement. It follows from the proportionality between the amount of non-evaporable  water and the solid volume of the cement paste that the former volume can be used as a  measure of the quantity of the cement gel present, i.e. of the degree of h
ydration.   The manner in which water is held in a cement paste determines the energy of  binding. For instance, 1670 joules (400 calories) are used in establishing the bond of 1 gram  of non-evaporable water, while the energy of the water of crystallization of Ca(OH)2, is  3560 joules per gram (850 cal / g). Likewise, the density of the water var
ies; it is  approximately 1.2 for non-evaporable, 1.1 for gel, and 1.0 for free water. It has been  suggested that the increase in the density of the adsorbed water at low surface  concentrations is not the result of compression but is caused by the orientation of the  molecules in the adsorbed phase due to the action of the surface forces. A conf
irmation of  the hypothesis that the properties of adsorbed water are different from those of free water  is afforded by measurements of the absorption of microwaves by hardened cement paste.      Heat of Hydration of Cement    In common with many chemical reactions, the hydration of cement compounds is  exothermic, up to 500 joules per gram (120 
cal / g) of cement being liberated. Since the  conductivity of concrete is comparatively low, it acts as an insulator, and in the interior of  a large concrete mass, hydration can result in a large rise in temperature. At the same time  the exterior of the concrete mass loses some heat so that a steep temperature gradient may  be established, and 
during subsequent cooling of the interior serious cracking may result.  This behaviour is, however, modified by the creep of concrete.   At the other extreme, the heat produced by the hydration of cement may prevent  freezing of the water in the capillaries of freshly placed concrete in cold weather, and a  high evolution of heat is therefore adva
ntageous. It is clear, then, that it is advisable to  know the heat-producing properties of different cements in order to choose the most  suitable cement for a given purpose.   The heat of hydration is the quantity of heat, in joules per gram of unhydrated  cement, evolved upon complete hydration at a given temperature. The most common  method of
 determining the heat of hydration is by measuring the heats of solution  of unhydrated and hydrated cement in a mixture of nitric and hydrofluoric acids: the  difference between the two values represents the heat of hydration. This method is  described in BS 4550: Part 3: Section 3.8: 1978, and is similar to the method of ASTM  Standard C 186-78.
 While there are no particular difficulties in this test, care should be  taken to prevent carbonation of the unhydrated cement as the absorption of one per cent  of CO2 results in an apparent decrease in the heat of hydration of 24.3 joules per gram  (5.8 cal / g) out of a total of between 250 and over 420 joules per gram (60 and 100 cal / g).   
The temperature at which hydration takes place greatly affects the rate of heat  development, as shown by the data of Table 1.7 which gives the heat developed in 72  hours at different temperatures.    Table 1.7: Heat of Hydration Developed After 72 Hours at Different Temperatures     Strictly speaking, the heat of hydration, as measured, consists
 of the chemical heat  of the reactions of hydration and the heat of adsorption of  water on the surface of the gel  formed by the processes of hydration. The latter heat accounts for about a quarter of the  total heat of hydration. Thus  the heat of hydration is really a composite quantity.   For practical purposes it is not necessarily the total
 heat of hydration that matters  but the rate of heat evolution. The same total heat produced over a longer period can be  dissipated to a greater degree with a consequent smaller rise in temperature. The rate of  heat development can be easily measured in an adiabatic calorimeter, and typical  time - temperature curves obtained under adiabatic co
nditions are shown in Fig. 1.13.    Fig 1.13 Temperature rise in 1:2:4 concrete (water / cement ratio of 0.60) made with different  cements and cured adiabatically     For the usual range of Portland cements, Bogue observed that about one-half of  the total heat is liberated between 1 and 3 days, about  three-quarters in 7 days, and 83 to  91 per 
cent of the total heat in six  months. In fact, the heat of hydration depends of the  chemical composition  of the cement, and the heat of hydration of cement is very nearly a  sum of the heats of hydration of the individual compounds when hydrated separately. It  follows that, given the compound composition of cement, its heat of hydration can be
	calculated with a fair degree of accuracy. Typical values of the heat of hydration of pure  compounds are given in Table 1.8.    Table 1.8: Heat of Hydration of Pure Compounds     It may be noted that there is no relation between the heat of hydration and the cementing  properties of the individual compounds. Woods, Steinour and Starke tested a 
number of  commercial cements and, using the method of least squares, calculated the contribution of  individual compounds to the total heat of hydration of cement. They obtained equations of  the type -    Heat of hydration of 1 gram of cement =    (formula)    where the terms in brackets denote the percentage by weight of the individual compound
s	 present in cement.   Since in the early stages of hydration the different compounds hydrate at different  rates, the rate of heat evolution, as well as the total heat, depends on the compound  composition of the cement. It follows that by reducing the proportions of the compounds  that hydrate most rapidly (C3A and C3S) the high rate of heat li
beration in the early life of  concrete can be checked. The fineness of the cement also influences the rate of heat  development, an increase in fineness speeding up the reactions of hydration and therefore  the heat evolved, but the total amount of heat liberated is not affected by the fineness of  cement.   The influence of C3A and C3S can be ga
uged from Figs. 1.14 and 1.15. As  mentioned before, for many uses of concrete a controlled heat evolution is advantageous  and suitable cements have been developed. One such cement is low heat Portland cement  discussed in more detail in the next chapter. The rate of heat development of this and  other cements is shown in Fig. 1.16.    Fig. 1.14.
 Influence of C3A content on heat evolution (C3S content approximately constant)    Fig. 1.15. Influence of C3S content on heat evolution (C3A content approximately constant)    Fig. 1.16 Development of heat of hydration of different cements cured at 21 C (70 F)  (water / cement ratio of 0.40)     The quantity of cement in the mix will also affect
 the total heat development: thus  the richness of the mix can be varied in order to help the control of heat development.      Influence of the Compound Composition on Properties of Cement    In the preceding section it was shown that the heat of hydration of cement is a simple  additive function of the compound composition of cement. It would se
em, therefore, that  the various hydrates retain their identity in the cement gel, which can be considered thus to  be a fine physical mixture or to consist of copolymers of the hydrates. A further  corroboration of this is obtained from the measurement of specific surface of hydrated  cements containing different amounts of C3S and C2S: the resul
ts agree with the specific  surface areas of hydrated neat C3S and C2S. Likewise, the water of hydration agrees with  the additivity of the individual compounds.   This argument does not, however, extend to all properties of hardened cement  paste, notably to shrinkage, creep, and strength; nevertheless, the compound composition  gives some indica
tion of the properties to be expected. In particular, the composition  controls the rate of evolution of heat of hydration and the resistance of cement to sulphate  attack, so that limiting values of oxide or compound composition of different types of  cement are prescribed by some specifications. The limitations of ASTM Standard C 150-78a  are le
ss restrictive than they used to be (see Table 1.9).    Table 1.9: ASTM Specification C 150-78a: Compound Composition Limits for Cement     The difference in the early rates of hydration of C3S and C2S - the two silicates  primarily responsible for the strength of cement paste - has been mentioned earlier. A  convenient approximate rule assumes th
at C3S contributes most to the strength  development during the first four weeks and C2S influences the gain in strength from four  weeks onwards. At the age of about one year the two compounds, weight for weight,  contribute approximately equally to the ultimate strength. Neat C3S and neat C2S have  been found to have a strength of the order of 7
0	MPa (10 000 psi) at the age of 18  months, but at the age of 7 days C2S had no strength while the strength of C3S was about  40 MPa (6000 psi). The development of strength of neat compounds is shown in Fig. 1.17.    Fig. 1.17. Development of strength of pure compounds     As mentioned on page 14, the calcium silicates appear in commercial cement
s	in  "impure" form. These impurities may strongly affect the rate of reaction and of strength  development of the hydrates. For instance, the addition of 1 per cent of Al2O3 to pure C3S  increases the early strength of the paste, as shown in Fig. 1.18. According to Verbeck, this  increase strength probably results from activation of the silicate 
crystal lattice due to  introduction of the alumina (or magnesia) into the crystal lattice with resultant activating  structural distortions.    Fig. 1.18. Development of strength of pure C3S and C3S with 1 per cent of Al2O3       The influence of the other major compounds on the strength development of  cement has been established less clearly. C
3A contributes to the strength of the cement  paste at one to three days, and possibly longer, but causes retrogression at an advanced  age, particularly in cements with high C3A or (C3A + C4AF) content. The role of C3A is  still controversial.   The role of C4AF in the development of strength of cement is also debatable, but  there certainly is n
o	appreciable positive contribution. It is likely that colloidal hydrated  CaO.Fe2O3 is deposited on the cement grains, thus delaying the progress of hydration of  other compounds.   From knowledge of the contribution to strength of the individual compounds  present it might be possible to predict the strength of cement on the basis of its compoun
d	 composition. This would be in the form of a formula of the type -    strength = (formula),    where the symbols in brackets represent the percentage by weight of the compound, and a,  b, etc. are constant parameters representing the contribution of one per cent of the  corresponding compound to the strength of the cement paste.   The use of suc
h	a formula would make it easy to forecast at the time of  manufacture the strength of cement and would reduce the need for conventional testing. In  practice, however, the influence of different compounds is not always significant and has  been found to depend on age and on the curing conditions. In general terms, an increase in  the C3S content 
increases strength up to 28 days; Fig. 1.19 shows the 7-day strength of  standard mortar made with cements of different composition and obtained from different  works. The C2S content has a positive influence on strength at 5 and 10 years only, and  C3A a positive influence up to 7 or 28 days but a negative influence later on. The  influence of al
kalis is considered on page 48).    Fig. 1.19. Relation between 7-day strength of cement paste and the C3S content in cement  (Each mark represents cement from one plant)     Prediction of the effects of compounds other than silicates on strength is  unreliable. According to Lea, these discrepancies may be due to the presence of glass in  clinker,
 discussed more fully in the succeeding section. In other words, the relations  observed are statistical in nature, and deviations arise from the fact that we are ignoring  some of the variables involved. It can be argued, in any case, that all constituents of  hydrated Portland cement contribute in some measure to strength in so far as all produc
ts  of hydration fill space and thus reduce porosity.   Furthermore, there are some indications that the additive behaviour cannot be fully  realized. In particular, Powers suggested that the same products are formed at all stages of  hydration of the paste; this follows from the fact that for a given cement the surface area  of hydrated cement is
 proportional to the amount of water of hydration, whatever the  water / cement ratio and age. Thus the fractional rates of hydration of all compounds in a  given cement would be the same. This is probably the case only after the rate of diffusion  through the gel coating has become the rate-determining factor, but not at early ages, say  up to 7 
days. Confirmation of equal fractional rate of hydration was obtained by Khalil and  Ward.   More important still, the composition is not the same at different points in space.  This arises from the fact that for diffusion to take place from the face of the still  unhydrated part of the cement grain to the space outside (see p. 13), there must be 
a  difference in ion concentration: the space outside is saturated but that inside is  supersaturated. This diffusion varies the rate of hydration. It is likely, therefore, that  neither the suggestion of equal fractional rates of hydration nor the assumption that each  compound hydrates at a rate independent of other compounds is valid. Probably,
 the exact  composition of the individual compounds in the anhydrous cement is affected by the oxide  composition of the cement as a whole and therefore the compounds in different cements  are not identical. Likewise, the oxide composition influences to some extent the rate of  formation of the individual hydrates, but our understanding of the hyd
ration rates is still  unsatisfactory.   For instance, the amount of heat of hydration per unit weight of hydrated material  has been found to be constant at all ages (see Fig. 1.20), thus suggesting that the nature of  the products of hydration does not vary with time. It is therefore reasonable to use the  assumption of equal fractional rates of
 hydration within the limited range of composition  of ordinary and rapid hardening Portland cements. However, other cements which have a  higher C2S content than ordinary or rapid hardening cements do not conform to this  behaviour. Measurements of heat of hydration indicate that C3S hydrates earlier, and  some C2S is left to hydrate later.    Fi
g. 1.20. Relation between the heat of hydration and the amount of non-evaporable  water for ordinary Portland cement     Furthermore, the initial framework of the paste established at the time of setting  affects to a large degree the subsequent structure of the products of hydration. This  framework influences especially the shrinkage and develop
ment of strength. It is not  surprising, therefore, that there is a definite relation between the degree of hydration and  strength. Figure 1.21 shows, for instance, an experimental relation between the  compressive strength of concrete and the combined water in a cement paste with a  water / cement ratio of 0.25. These data agree with Powers' obs
ervations on the gel / space  ratio, according to which the increase in strength of a cement paste is a function of the  increase in the relative volume of gel, regardless of age, water / cement ratio, or compound  composition of cement. However, the total surface area of the solid phase is related to the  compound composition, which does affect t
he actual value of the ultimate strength.    Fig. 1.21. Relation between compressive strength and combined water content	The effects of the minor compounds on the strength of cement paste have not  been thoroughly investigated as these compounds were not thought to be of importance as  far as strength is concerned.  K2O is believed to replace 
one molecule of CaO in C2S with  a consequent rise in C3S content above that calculated. Tests on the influence of alkalis  have shown that the increase in strength beyond the age of 28 days is  strongly affected by  the alkali content: the greater the amount of alkali  present the lower the gain in strength.  This has been confirmed by two statis
tical evaluations of strength of several hundred  commercial statistical cements. The poor gain in strength between 3 and 28 days can be  attributed more specifically to water-soluble K2O present in the cement. On the other  hand, in the total absence of alkalis, the early strength of cement paste can be abnormally  low. Accelerated strength tests
 (see Page 570) have shown that, up to 0.4 per cent of  Na2O, strength increases with an increase in the alkali content (Fig. 1.22). The role of  alkalis is important with reference to cements made by the dry process, which sometimes  contain relatively large amounts of alkalis.    Fig. 1.22. Effect of alkali content on accelerated strength     Th
e alkalis are known to react with the so-called alkali-reactive aggregates  (see p. 158) and cements used under such circumstances often have their alkali content  limited to 0.6 per cent (measured as equivalent soda).  Such cements are referred to as  low-alkali cements.   We can see then that the alkalis are an important constituent of cement, b
ut fuller  information on their role is yet to be obtained. It may be noted that the use of pre-heaters  in modern cement plants has led to an increase in the alkali content of cement made from  given raw materials. Limiting the alkali content too severely results in an increased energy  consumption.  A more efficient dust collection also increase
s the alkali content of the  cement because the dust contains a large amount of alkalis; this may be as high as 15 per  cent, in which case the dust has to be discarded.    Effects of Glass in Clinker    It may be recalled that during the formation of cement clinker in the kiln some 20 to 30  per cent of the material becomes liquid; on subsequent 
cooling, crystallization takes place  but there is always some material which undercools to glass. In fact, the rate of cooling of  clinker greatly affects the properties of cement: if cooling were so slow that full  crystallization could be achieved (e.g. in a laboratory), B-C2S might become converted to  Y-C2S, this conversion being accompanied 
by expansion and powdering, known as  dusting. Furthermore, Y-C2S hydrates too slowly to be a useful cementitious material.  However, Al2O3, MgO and the alkalis may stabilize B-C2S, even on very slow cooling in  all practical cases.   Another reason why some glass is desirable is the effect of glass on the crystalline  phases. Alumina and ferric o
xide are completely liquefied at clinkering temperatures, and  on cooling produce C3A and C4AF. The extent of glass formation would thus affect these  compounds to a large degree while the silicates, which are formed mainly as solids, would  be relatively unaffected. It may be noted too that glass may also hold a large proportion of  minor compoun
ds such as the alkalis and MgO, which are thus not available for expansive  hydration. Thus a rapid cooling of high-magnesia clinkers is advantageous. Since the  aluminates are attacked by sulphates their presence in glass would also be an advantage.  C3A and C4AF in glass form can hydrate to a solid solution of C3AH6 and C3FH6 which  is resistant
 to sulphates.   On the other hand, there are some advantages of a lower glass content. In some  cements, a greater degree of crystallization leads to an increase in the amount of C3S  produced. Also, a high glass content adversely affects the grindability of clinker.   It can be seen, then, that a strict control of the rate of cooling of clinker 
so as to  produce a desired degree of crystallization is very important. The range of glass content in  commercial clinkers, determined by the heat of solution method, is between 2 and 21 per  cent. An optical microscope indicates much lower values.   It may be recalled that the Bogue compound composition assumes that the clinker  has crystallized
 completely to yield its equilibrium products, and, as we have seen, the  reactivity of glass is different from that of crystals of similar composition.   It can be seen then that the rate of cooling of clinker, as well as, possibly, other  characteristics of the process of cement manufacture, affects the strength of cement and  defies attempts to
 develop a formula of the type mentioned in the preceding section.  Nevertheless, if one process of manufacture is used and the rate of cooling of clinker is  kept constant, there is a definite relation between compound composition and strength.      Tests on Physical Properties of Cement    The manufacture of cement requires stringent control, an
d a number of tests are  performed in the cement works laboratory to ensure that the cement is of the desired  quality and that it conforms to the requirements of the relevant national standards. It is  desirable none the less for the purchaser or for an independent laboratory to make  acceptance tests or, more frequently, to examine the propertie
s of a cement to be used for  some special purpose. Tests on the chemical composition and fineness have already been  described; further tests prescribed by BS 4550 : Part 3 : 1978 for ordinary and rapid  hardening Portland cements are given below. Other relevant standards are mentioned  when other types of cement are discussed in Chapter 2.    Co
nsistence of Standard Paste    For the determination of the initial and final setting times and for the Le Chatelier  soundness test, neat cement paste of a standard consistence has to be used. It is, therefore,  necessary to determine for any given cement the water content of the paste which will  produce the desired consistence.   The consistenc
e is measured by the Vicat apparatus shown in Fig. 1.23, using a 10 mm  diameter plunger fitted into the needle holder. A trial paste of cement and water is mixed  in a prescribed manner and placed in the mould. The plunger is then brought into contact  with the top surface of the paste and released. Under the action of its weight the plunger  wil
l penetrate the paste, the depth of penetration depending on the consistence. This is  considered to be standard, in the meaning of BS 4550 : Part 3 : Section 3.5 : 1978, when  the plunger penetrates the paste to a point 5 + 1 mm from the bottom of the mould. The  water content of the standard paste is expressed as a percentage by weight of the dr
y  cement, the usual range of values being between 26 and 33 per cent.    Fig. 1.23. Vicat apparatus    Setting Time    The physical processes of setting were discussed on page 19; here, the actual  determination of setting times will be briefly dealt with. The setting times of cement are  measured using the Vicat apparatus (Fig. 1.23) with differ
ent penetrating attachments.   For the determination of the initial set a round needle with a diameter  1.13 +/-  0.05 mm is used. This needle, acting under a prescribed weight, is used to penetrate a  paste of standard consistence placed in a special mould.  When the paste stiffens  sufficiently for the needle to penetrate only to a point 5 +/- 1
 mm from the bottom, initial set  is said to have taken place.  Initial set is expressed as the time elapsed since the mixing  water was added  to the cement. A minimum time of 45 minutes is prescribed by BS 12 :  1978  for ordinary and rapid hardening Portland cements and for Portland  blast-furnace  cement; for low heat Portland cement (BS 1370 
: 1979) the  minimum setting time is 60  minutes. The initial setting time of high-alumina cement is prescribed by BS 915 : 1972 as  between 2 and 6 hours.   Final set is determined by a similar needle fitted with a metal attachment hollowed  out so as to leave a circular cutting edge 5 mm in diameter and set  0.5 mm behind the tip  of the needle.
 Final set is said to have taken place when the needle, gently lowered to the  surface of the paste, makes an  impression on it but the circular cutting edge fails to do so.  The final setting  time is reckoned from the moment when mixing water was added to the  cement, and is required by the relevant British Standards to be not more than 10 hours
 for  ordinary, rapid hardening, low heat, and blast-furnace Portland cements. Rather  surprisingly, in 1978, the Swiss standard SIA 215 deleted its requirements for final set.  For high-alumina cement BS 915 : 1972 specifies the final setting time as not more than 2  hours after the initial set.   It may sometimes be useful to take advantage of t
he observation that for  the majority of American commercial ordinary and rapid hardening Portland  cements at room temperature, the initial and final setting times are related. An  approximate (within +/- 15 min) simple relation has been observed: final setting time (min)  = 90 + 1.2 x initial setting time (min).   Since the setting of cement is 
affected by the temperature and the humidity of the  surrounding medium, these are specified by BS 4550 : Part 3 : Section 3.6 : 1978: mixing  room temperature of 20 +/- 2 C (68 +/- 4 F) and minimum relative humidity of 65 per cent;  curing room temperature of 20 +/- 1 C (68 +/- 2 F) and relative humidity of air of not less than  90 per cent.   Te
sts have shown that setting of cement paste is accompanied by a change in the  ultrasonic pulse velocity through it (cf. page 581) but it has not been possible to develop  an alternative method of measurement of setting time of cement. Attempts at using  electrical measurements have also been unsuccessful, mainly because of the influence of  admix
tures on electrical properties.	It should be remembered that the speed of setting and the rapidity of hardening,  i.e. of gain of strength, are entirely independent of one another. For instance, the  prescribed setting times of rapid hardening cement are no different from those for ordinary  Portland cement, although the two cements harden at di
fferent rates.   It may be relevant to mention here that the setting time of concrete can also be  determined, but this is a different property from the setting time of cement. ASTM  Standard C403-77 lays down the procedure for the former, which uses a Proctor  penetration needle applied to mortar sieved from the given concrete. The definition of 
this  setting time is arbitrary as there is no abrupt advent of setting in practice. The Russians  have attempted to define the setting time of concrete by the minimum resistance between  two embedded metal electrodes between which is passed a high-frequency electric current.    Soundness    It is essential that a cement paste, once it has set, do
es not undergo a large change in  volume. In particular, there must be no appreciable expansion, which under conditions of  restraint could result in a disruption of the hardened cement paste. Such expansion may  take place due to the delayed or slow hydration or other reaction of some compounds  present in the hardened cement, namely free lime, m
agnesia, and calcium sulphate.   If the raw materials fed into the kiln contain more lime that can combine with the  acidic oxides, the excess will remain in a free condition. This hard burnt lime hydrates only  very slowly and, since slaked lime occupies a larger volume than the original free calcium  oxide, expansion takes place. Cements which e
xhibit this expansion are known as  unsound.   Lime added to cement does not produce unsoundness because it hydrates rapidly  before the paste has set. On the other hand, free lime, present in clinker is intercrystallized  with other compounds and is only partially exposed to water during the time before the  paste has set.   Free lime cannot be d
etermined by chemical analysis of cement since it is not  possible to distinguish between unreacted CaO and Ca(OH)2 produced by a partial  hydration of the silicates when cement is exposed to the atmosphere. On the other hand, a  test on clinker, immediately after it has left the kiln, would show the free lime content  since no hydrated cement is 
then present.   A cement can also be unsound due to the presence of MgO, which reacts with  water in a manner similar to CaO. However, only periclase (crystalline MgO) is  deleteriously reactive, and MgO present in glass is harmless.   Calcium sulphate is the third compound liable to cause expansion: in this case,   calcium sulphoaluminate is form
ed. It may be recalled that a hydrate of calcium  sulphate - gypsum - is added to cement clinker in order to prevent flash set, but if gypsum is present  in excess of the amount that can react with C3A during setting, unsoundness in the form  of a slow expansion will result. For this reason, BS 12 : 1978 limits very strictly the  amount of gypsum 
that can be added to clinker, but the limits are well on the safe side as  far as the danger of unsoundness is concerned.   Since unsoundness of cement is not apparent until after a period of months or  years it is essential to test the soundness of cement in an accelerated manner: a test  devised by Le Chatelier is prescribed by BS 4550: Part 3: 
Section 3.7: 1978. The Le  Chatelier apparatus, shown in Fig. 1.24, consists of a small brass cylinder split along its  generatix. Two indicators with pointed ends are attached to the cylinder on either side of  the split; in this manner, the widening of the split, caused by the expansion of cement, is  greatly magnified and can be easily measured
.	The cylinder is placed on a glass plate, filled  with cement paste of standard consistence, and covered with another glass plate. The  whole assembly is then immersed in water at 20 +/- 1 C (68 +/- 2 F) for 24 hours. At the  end of that period, the distance between the indicators is measured, and the mould is  immersed in water again and brought
 to the boil in 25 to 30 minutes. After boiling for one  hour the assembly is taken out, and, after cooling, the distance between the indicators is  again measured. The increase in this distance represents the expansion of the cement, and  for Portland cements is limited to 10 mm. If the expansion exceeds this value a further test  is made after t
he cement has been spread and aerated for 7 days. During this time, some of  the lime may hydrate or even carbonate, and a physical breakdown in size may also take  place. At the end of the 7-day period, the Le Chatelier test is repeated and the expansion  of aerated cement must not exceed 5 mm. A cement which fails to satisfy at least one of  the
se tests should not be used.    Fig. 1.24. Le Chatelier apparatus     The Le Chatelier test detects unsoundness due to free lime only. Magnesia is rarely  present in large quantities in the raw materials from which cement is manufactured in  Britain, but is encountered in other countries. For this reason, in the United States for  instance, soundn
ess of cement is checked by the autoclave test, which is sensitive to both  free magnesia and free lime. In this test, prescribed by ASTM Standard C151-77, a neat  cement bar 25 mm (or 1 in.) square in cross-section and with a 250 mm (or 10 in.) gauge  length is cured in humid air for 24 hours. The bar is then placed in an autoclave (a high-pressu
re  steam boiler), which is raised to a temperature of 216 C (420 F) (steam pressure of  2 +/- 0.07 MPa (295 psi)) in 60 +/- 15 min, and maintained at this temperature for 3 hours.  The expansion of the bar due to autoclaving must not exceed 0.8 per cent. The high steam  pressure accelerates the hydration of both magnesia and lime.   The results o
f	the autoclave test are affected not only by the compounds causing  expansion, but also by the C3A content, by additives to cement, and are also subject to  other anomalies. The test gives, therefore, no more than a broad indication of the risk of  long-term expansion in practice.   No test is available for the detection of unsoundness due to an 
excess of calcium  sulphate, but its content can be easily determined by chemical analysis.    Strength of Cement    The mechanical strength of hardened cement is the property of the material that is perhaps  most obviously required for structural use. It is not surprising, therefore, that strength  tests are prescribed by all specifications for c
ement.   The strength of mortar or concrete depends on the cohesion of the cement paste,  on its adhesion to the aggregate particles, and to a certain extent on the strength of the  aggregate itself. The last factor is not considered at this stage, and is eliminated in tests on  the quality of cement by the use of standard aggregates.   Strength t
ests are not made on a neat cement paste because of difficulties of  moulding and testing with a consequent large variability of test results. Cement-sand  mortar and, in some cases, concrete of prescribed proportions and made with specified  materials under strictly controlled conditions are used for the purpose of determining the  strength of ce
ment.   There are several forms of strength tests: direct tension, direct compression, and  flexure. The latter determines in reality the tensile strength in bending because, as is well  known, cement paste is considerably stronger in compression than in tension. Since the  flexure test is not used in Great Britain and little used elsewhere it wil
l	not be further  discussed.   The direct tension test used to be commonly employed but pure tension is rather difficult  to apply so that the results of such a test show a fairly large scatter. Furthermore, since  structural techniques are designed mainly to exploit the good strength of concrete in  compression, the tensile strength of cement is 
often of lesser interest than its compressive  strength. For these reasons, the tension test has gradually given way to compression tests.   However, the tension test still exists in some countries as a permitted test for a  one-day strength of rapid hardening Portland cement, and the details of the test as  prescribed in the 1971 edition of BS 12
 may be of interest. In this test, a 1:3 cement-sand  mortar with a water content of 8 per cent of the weight of the solids is mixed and moulded  into a briquette of the shape shown in Fig. 1.25. The sand is the standard Leighton  Buzzard sand obtained from a quarry near the village of that name in Bedfordshire. This  sand consists of pure siliceo
us material and is practically all of one size; all particles are  nearly spherical and are smaller than an 850 um (No. 20 ASTM) sieve and at least 90 per  cent of the sand is retained on a 600 um (No. 30 ASTM) sieve.    Fig. 1.25. Briquette for the tension test of mortar     The briquettes are moulded in a standard manner, cured for 24 hours at a
	temperature between 18 and 20 C (64 and 68 F) in an atmosphere of at least 90 per cent  relative humidity, and tested in direct tension, the pull being applied through special jaws  engaging the wide ends of the briquette. BS 12 : 1958 prescribes the minimum one-day  strength of rapid hardening Portland cement as 2.1 MPa (300 psi), taken as the 
average  value for six briquettes.   There are two standard methods of testing the compressive strength of  cement: one uses mortar, the other concrete.   In the mortar test, a 1:3 cement-sand mortar is used. The sand is again the  standard Leighton Buzzard sand, and the weight of water in the mix is 10 per cent of the  weight of the dry materials
. Expressed as a water / cement ratio, this corresponds to 0.40  by weight. A standard procedure, prescribed by BS 4550 : Part 3 : Section 3.4 : 1978, is  followed in mixing, and 70.7 mm (2.78 in.) cubes are made using a vibrating table with a  frequency of 200 Hz applied for two minutes. The cubes are demoulded after 24 hours  and further cured i
n water until tested in a wet-surface condition. The BS 12 : 1978  requirements for minimum strengths (average values for three cubes) are given in Table  1.10. It is expected that the mortar test will be deleted when BS 4550 is next revised.    Table 1.10: BS 12: 1978 Requirements for Strength of Cement     The vibrated mortar test gives fairly r
eliable results but it has been suggested that  mortar made with one-size aggregate leads to a greater scatter of strength values than  would be obtained with concrete made under similar conditions. Moreover, the values of  strength obtained in a test should approximate the level of strength generally found in  concrete, and this would require the
 use of a water / cement ratio higher than that used in  the mortar test. It can also be argued that we are interested in the performance of cement  in concrete and not in mortar, especially one made with a one-size aggregate and never  used in practice. For these reasons, a test on concrete was introduced in 1958 and is now  covered by BS 4550 : 
Part 3 : Section 3.4 : 1978.   In the concrete test, one of three water / cement ratios can be used, viz. 0.60,0.55  and 0.45. The amounts of coarse and fine aggregate, which have to come from particular  quarries, are specified in BS 4550 : Parts 4 and 5 : 1978. Batches of 100 mm (or 4 in.)  cubes are made by hand in a prescribed manner; the temp
erature and humidity conditions  of the mixing room, curing chamber, compression testing room, and the temperature of  the water curing tank are specified. The BS 12: 1978 requirements for the minimum values  of the average strength of three cubes at each age are given in Table 1.10. Apart from  satisfying these minima, the strength at later ages 
has to be higher than at an earlier age, as  strength retrogression might be a sign of unsoundness or other faults in the cement. The  requirement of strength increase with age applies also to the vibrated mortar cubes.   The influence of cement on the properties of mortar and concrete is qualitatively  the same, and the relation between the stren
gths of corresponding specimens of the two  materials is linear. This is shown, for instance, in Fig. 1.26: mortar and concrete of fixed  proportions, each with a water / cement ratio of  0.65 were used. The strengths are not the  same for the specimens of each pair, at least in part because specimens of different shape  and size were used, but th
ere may also be an inherent quantitative difference between the  strengths of  mortar and concrete due to the greater amount of entrapped air in mortar.    Fig. 1.26. Relation between the strengths of concrete and mortar of the same  water / cement ratio     It may be interesting to know to what extent commercial cements comply with  the standard 
requirements. Data on samples of cement which failed to comply in  Switzerland in the period 1965-1975 are given in Table 1.11; of the total of over 7000  samples tested, 1.1 per cent failed.  We can see thus that failure is rare and, with respect to  strength, virtually non-existent.    Table 1.11: Distribution of Samples of Cement Failing to Com
ply with Standard Tests