CIRIA C660: Control of Contraction Induced Cracking In Concrete

F.Papworth & P.Bamforth,

Synopsis: In the UK, BS8007 has provided the basis for the design for early-age thermal cracking. This is to be replaced by EN1992-3 and in conjunction with the replacement of the general design code, BS8110 by EN1992-1-1 this has led to significant changes in some aspects of design. BS8007 was supported by CIRIA 91 which provided background to the design method and data for use in the design process. This has been updated and replaced by CIRIA C660 which brings the design into line with the Euro-codes and provides current information required to support the design process. The significant changes are described and their implications are discussed. AS3600 and AS3735 crack control provisions are based on the principles of BS 8110 and may need revision to take account of the changes outlined herein.

Introduction

Cracking can occur due to restraint to contraction as concrete: cools from its peak hydration temperature (heat of hydration contraction) cools to lowest ambient temperatures (ambient temperature contraction) shrinks as it dries (drying shrinkage) shrinks due to a change in chemical nature during hydration (autogenous shrinkage)

Where the contraction is not the same on the top and bottom surfaces curling can occur and this can lead to additional cracking stresses.

Contraction cracking is not necessarily inconsistent with good practice and in many cases it may be either unnecessary or uneconomical to avoid cracking entirely. Measures to minimise or avoid contraction cracking are available i.e.

selection of cementitious materials to slow hydration and give time for heat to escape selection of cementitious materials to reduce cementitious hydration and heat selection of aggregates to reduce the contraction and increase the strain capacity taking measures to reduce the concrete delivered, & hence maximum insitu, temperatures cooling of the concrete insitu to reduce the maximum temperature insulating the concrete to reduce temperature differentials planning the construction sequence to minimise restraint prestressing.

In the UK, design for early-age thermal cracking has been dealt with using BS8007 (1) and together with modifications to suit Australian conditions (2) this provided a basis for the recommendations of AS3735-2001 (3). BS8007 was supported by CIRIA 91 (4) which provided background to the design method and data for use in the design process. CIRIA 91 has been updated and replaced by CIRIA C660 (5) to take account of new knowledge of the performance of a range of concreting materials; the increasing use of higher strength concrete; and changes in the design process arising from the introduction of Euro-codes, in particular EN1992-1-1 (6) which will replace BS8110 (7) as the general design code in the UK in 2009; and EN1992-3 (8) which will replace BS8007 for water retaining structures in 2011. In addition to bringing the design into line with Euro-codes, CIRIA C660 differs from CIRIA 91 in the following respects.

Values of temperature drop (T1) for Portland cement (GP cement) have been revised and additional information is provided on concretes containing fly ash and ground granulated blast-furnace slag (ggbs). Silica fume can be treated as GP cement and provision made for reduced cement content.

Information is provided on autogenous shrinkage Additional information is given on different forms of restraint, how they may be calculated and

how they affect crack width Tensile strain capacity is dealt with more comprehensively A method for reinforcement design has been developed to deal with cracking caused by

temperature differentials in thick sections Guidance is given on methods for minimising the risk of cracking Advice is provided on specification, testing and in situ monitoring.

The Design Approach

BS8007 uses a strain based approach (9) and this has been maintained in EN1992-3. It is generally assumed that compressive stresses induced during heating are relieved by creep. Hence the restrained contraction, εr that may lead to cracking, is related to the drop from the peak temperature in the section to the mean ambient temperature T1, the coefficient of thermal expansion of concrete c

the restraint R, and a creep coefficient K according to the equation εr = c. T1. K. R

Temperature drop T1

C660 provides T1 data in chart form and in a spreadsheet calculator for various combinations of GP cement with fly ash or ggbs based on semi-adiabatic test results (10). A comparison of predicted and measured results (11, 12, 13) validated the model for a range of concrete mix types and temperatures. Figure 1 gives examples of the heat generated for different cementitious systems as proposed by CIRIA. Figure 2 shows the calculated adiabatic curve for 300kg/m3 of GP cement (CEM 1) at different concrete delivered temperatures (CDT’s). C660.

Figure 1 – Heat of Reference Cements Figure 2 – Adiabatic Temperature Rise

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However, the charts were derived with ultimate heat outputs for UK cement, fly ash and slag, a delivered concrete temperature of 20oC and a mean ambient temperature of 15oC. Hence the charts are unlikely to be useful in Australia and New Zealand. In the spreadsheet calculator the delivered concrete temperature and ambient temperatures are variable and hence adiabatic can be calculated for Australian and New Zealand climates. The heat out puts for UK cement, fly ash and slag might be inappropriate for Australia and New Zealand. The semi adiabatic temperature rise using an insulated hot box (eg 1m cubic from plywood with 25mm insulation) should be measured for the proposed concrete and the adiabatic temperature rise calculated and compared it with the adiabatic given by the CIRIA C660 spreadsheet. If there is a significant difference between the actual adiabatic and the C660 adiabatic the later should be modified before going on to calculate the insitu temperature rise.

Once the adiabatic temperature is confirmed the insitu temperature rise can be calculated using the CIRIA C660 spreadsheet.

Figure 3 – Maximum, Surface and Differential Temperatures for a 1200mm wall using GP cement, 18mm ply and a CDT of 10C

Figure 4 – Maximum Temperature for a 1200mm wall using GP cement, 18mm ply and various CDT’s

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Figure 3 shows a typical individual calculation for ambient temperature of 17-34C and Figure 4 shows the comparison of peak temperatures for various CDT’s. An increase in CDT of 25C leads to an increase in peak temperature of approximately the same amount.

Autogenous shrinkageAutogenous shrinkage is not a new phenomenon but it has been assumed hitherto that it will only occur in concretes with very low w/c ratios (14). EN1992-1-1 estimates autogenous shrinkage based on the strength alone and assumes that it occurs to some degree in all structural concretes and this is supported by research data (Appendix 4 of C660). EN1992-1 takes no account of the cement type but there is evidence that mineral additions affect autogenous shrinkage with an increase using silica fume or ggbs and a reduction with fly ash. Further research in needed as high strength concretes become more widely used. The calculator for autogenous shrinkage is of the form shown in Figure 5.

Figure 5 – Autogenous Shrinkage CalcualtionStrength class C40/50 MPaUltimate autogenous shrinkage ε ca (∞ ) 75

Autogenous shrinkage ε ca at 3 days 22 microstrain90 days 63 microstrain

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RestraintIn line with BS8007, EN1992-3 provides a single coefficient of 0.5 to take account of restraint and creep. This simple approach must be assumed to deal with the worst case, but prevents benefit being taken in situations where the worst case does not occur. Based on published data (15, 16, 17) C660 recommends a creep coefficient K = 0.65. This implies a worst case R = 0.5/0.65 ≈ 0.8. This is not surprising as a new element does have considerable inherent stiffness when cooling commences and, in general, will not be totally dominated by the element against which it is cast. C660 describes a basis for calculating edge restraint (18) which has been validated by comparison with measured restraints in walls. Values at the joint are commonly in the range for 0.4 to 0.7 indicating values of R x K in the range from 0.26 to 0.46 and hence that the coefficient of 0.5 given in EN1992-3 is safe in the majority of cases.

Tensile strain capacityThe tensile strain capacity εctu is the maximum strain that the concrete can withstand without a continuous crack forming. A comprehensive review of data (19) demonstrated a linear relationship between εctu under rapid loading and the ratio of the tensile strength fctm to the elastic modulus Ecm in compression. Using property data and age functions provided in EN1992-1-1, values for εctu have been estimated and adjusted to take account of sustained loading during the early-age thermal cycle. Values for strength class C30/37 (EN1992 presents strength class as both cylinder/cube, hence for C30/37, f’c = 30 MPa) estimated on this basis are given in Table 2. To estimate the strain capacity for other strength classes, the value obtained for C30/37 is multiplied by 0.63 + (1.25f’c/100) for f’c in the range from 20 to 50 MPa. For higher strength concrete the value obtained for 50 MPa is used.

Table 2 Estimated Values of Εctu (In Microstrain) for Strength Class C30/37 under Sustained Loading Using Different Aggregate Types [Early-age = 3 Days, Long-term ≥ 28 Days]

Aggregatetype

Basalt Flint gravel

Quartzite Granite, Gabbro

Limestone, Dolerite

Sandstone Lightweight

Early-age 63 65 76 75 85 108 115

Long term 90 93 108 108 122 155 165

Comparison of the Design Methods of BS8007 and EN1992The design approach adopted by EN1992 is broadly similar to that of BS8007 but there are some significant and important differences as follows;

Different values of surface zone are used to estimate the minimum area of reinforcement Different surface zones are used to estimated the steel ratio for calculating of crack width EN1992-1-1 includes cover in the expressions for crack spacing and width The term fct/fb (tensile strength/bond strength) has been replaced by the coefficient k1 Cracking develops depends on whether the element is subject to edge restraint or end

restraint (20) and this is reflected in different expressions for calculating crack width Autogenous shrinkage is assumed to occur in all grades of structural concrete.

Calculating the minimum area of reinforcementThe minimum area of reinforcement As required to control the crack width is determined by ensuring that the stress transferred to the steel after a crack has occurred is below the yield strength of the steel. Expressions used by BS8007 and EN1992-1-1 are shown in Table 3.

Table 3 Expression for Estimating the Minimum Area of Reinforcement

BS8007 BS1992-1-1

Ac is the surface zone of 250mm or h/2, whichever is less

Act is the area of concrete in tension

fct is the tensile strength of the concrete fct,eff is the tensile strength of the concretefy is the yield strength of the steel fyk is the yield strength of the steel

k allows for non-uniform and self-equilibrating stress which leads to a reduction in restraint forces.kc takes account of the stress distribution in the section

A review of the development of the approach of both BS8007 and EN1992-1-1 was undertaken to understand the bases for the assumed surface zones. In doing so it was identified that the assumption that cracking is initiated from the surface may not be correct (21, 22). Under conditions of external restraint it is most likely that cracking will be initiated at the point where the temperature drop is the greatest, i.e. at the centre of the section (Figure 3) transferring stress from the full section to the reinforcement when a crack occurs.

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t0t1

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Figure 3 Cross-Section Through a Thick Wall Subject External Restraint

C660 has therefore developed appropriate surface zones taking account of both the temperature profile and the fact that in practice some compressive stresses must be relieved by a drop in temperature before tensile stress are generated. A comparison of the values recommended by C660 with those of the existing recommendations of EN1992-1-1 and BS8007 is shown in Figure 4.

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Figure 4 Surface Zones used in Estimating the Minimum Area of Reinforcement in Sections that are Dominated by External Restraint and Subject to Tension Through the Full Thickness

Estimating crack widths for elements subject to continuous edge restraintThe expressions for calculating crack width for elements subject to continuous edge restraint are given in Table 4. It should be noted that in EN1992-1-1 the characteristic crack width, wk is estimated, this being a value with only a 5% chance of being exceeded. This value is expected to be about 30% higher than the mean value (22, 23).

Table 4 Expressions for the Calculation of Crack Width

BS8007 EN1992-1-1

NO cover term c is the cover (mm)

fct/fb is the ratio of the tensile strength of the concrete to the bond strength = 0.67

k1 is a coefficient which takes account of the bond properties of the reinforcement = 0.8 increased in C660 to 1.14

φ is the bar diameter (mm)

ρ is the steel ratio based on a surface zone of 250mm of h/2, whichever is less

ρe,eff is the effective steel ratio based on a surface zone to a depth of 2.5 (c + φ/2) or h/2, whichever is less

εcr is the crack inducing strain

Hence, and,

The second term in the EN1992-1-1 expression appears to be very similar to that of BS8007. However, the way in which ρe,eff is calculated leads to very different results being based on a surface zone he,ef = 2.5(c + φ/2) or h/2 whichever is smaller, compared with a BS8007 value of h/2 or 250mm. For a 500mm thick wall, if c = 40mm and φ = 20mm he,ef = 2.5(40 +20/2) = 125mm, only half the value of 250mm used by BS8007. As the value of ρp,eff is inversely proportional to he,ef this will result in ρp,eff

being double the value used by BS8007, thus halving the value of the second term in the crack width expression. This difference is partially offset by a cover term but the net effect is for crack widths, estimated using EN1992-1-1, to be significantly lower than crack widths estimated using BS8007. With no other changes this would lead to a significant reduction in crack control reinforcement compared with that currently used as shown in Figure 5 (a).

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Figure 5 The Ratio of Reinforcement Requirements for Design to EN1992 and BS8007 (C30/37 Concrete; Plywood Formwork; Limiting Early-age Crack Width to 0.15 Mm; Cover as Shown)

Observations by the author suggest that the requirements of BS8007, while having been generally applicable, have occasionally led to crack widths in excess of those predicted and on this basis it would be unsafe to adopt a design that significantly reduces the current reinforcement requirements. The factors used in the design were therefore investigated and the bond coefficient k1 has been increased from 0.8 to 1.14 by applying the EN1992-1-1 factor of 0.7 applied when “good” bond cannot be guaranteed (0.8/0.7 = 1.14). The calculations shown in Figure 5 (a) have been repeated with the revised coefficient and the results are shown in Figure 5 (b). With the revised coefficient the steel requirements are closer to those of BS8007 with the normal range of cover. Higher steel ratios than those suggested by BS8007 are generally associated with high cover.

EXAMPLE 1: 300mm retaining wall on a rigid foundation (continuous edge restraint)

A high cement content including GP, fly ash and silica fume (>500kg/m3) and high cover (65mm) were used for durability in an aggressive saline environment. To achieve a mean crack width of 0.15mm AS3735 recommends >0.64% steel. This was achieved using 16mm bars at 200mm centres (=0.67%). The maximum allowable placing temperature was 35oC.

RestraintRestraint

EATC occurred, with a mean crack width of 0.25 mm and individual cracks up to 0.3mm. The large crack widths were attributed to the fact that the advice of AS3735 “to use higher steel ratios in extremely hot climates or for concretes with high cement contents” had not been followed, although AS3735 provides no specific guidance on how the increased steel ratio may be calculated. The design was checked using CIRA C660 as follows;

Strain developmentTemperature drop T1 (estimated for specified concrete) 39 oCCoefficient of thermal expansion αc (default value) 12 µε/oCRestraint R 0.77Autogenous shrinkage εca (based on compressive strength) 33 µεCreep coefficient K (default value) 0.65Restrained strain εr = K R (T1 αc + εca) 251 µεTensile strain capacity εctu (at 3 days) 100 µεCrack-inducing strain εcr = εr – 0.5 εctu 201 µεEstimated crack width wk See Table 5 (Characteristic value) 0.28 mm

The estimated characteristic crack width of 0.28mm indicates a mean value of about 0.21mm.

Increased reinforcement and control over the mix temperature was recommended

OPTIONSPlacing

temp (oC)Steel ratio (%)

Crack width

wk MeanDesign to AS3735 35 16mm@200 = 0.67 0.28 0.21Reduce placing tem 30 16mm@200 = 0.67 0.23 0.16Increase steel ratio 35 16mm@150 = 0.89 0.22 0.16Reduce placing temp and increase steel ratio 30 16mm@150 = 0.89 0.18 0.14

The high cover also contributed to the high crack widths. Estimated mean crack widths are 0.19mm and 0.16 mm for cover of 50mm and 40mm respectively for design to AS3735

The influence of coverThe net effect of cover alone on the crack width is shown in Figure 6. This has been recognised for many years. For example, Campbell-Allen & Hughes (2) recommended that “the placing of such reinforcement shall be as near to the surface of the concrete as is consistent with the requirements of adequate cover”. However, in relation to control of EATC, the effect of cover has previously not been quantified. Acknowledging that the crack will taper from the surface to the reinforcement it may be appropriate, when using high cover for durability, to design for a crack width at a cover of, say, 50mm, and to accept that the crack width at the surface will be wider. This approach is recommended by the author, although not included in C660, as it ismay be justified byas there is considerable evidence which indicates that protection of steel is related more to the quality and depth of the cover than to the crack width (24). As an example this approach might lead to acceptance of 0.2mm cracks at 65mm cover as opposed to 0.15mm cracks at 50mm cover.

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Figure 6 The Effect of Cover in a 300mm Wall Subject to a 30oc Temperature Drop and 70% Restraint

Estimating crack widths for elements subject to continuous END restraintThe method of BS8007 (and AS 3600 and AS3735) EATC assumes continuous edge restraint. EN1992-3 also recognises end restraint and C660 provides a design approach for both conditions. End restraint leads to substantially more reinforcement to control cracking because the restraint itself plays no part in preventing the cracks from widening (as it does under conditions of edge restraint). The crack width wk is determined by the tensile strength of the concrete fct; the stress transferred to the steel determined by the steel ratio ρ, the modular ratio αe; the elastic modulus of the steel Es; and the length over which debonding occurs Sr,max (for k and kc see Table 4). Hence,

(1)

Even when the minimum steel ratio is exceeded, crack widths may be significantly wider than achieved under conditions of edge restraint, although fewer cracks may occur (Table 5).

Table 5 Estimated Crack Widths (300mm Section, End Restraint, 16mm Bars at 150mm)

Cylinder Strength f’c (MPa) 20 25 37 45 50 55 60

Crack width (mm) 0.42 0.48 0.54 0.60 0.66 0.71 0.76

It is important, therefore, to recognise the nature of the restraint when designing reinforcement to control cracking. End restraint typically occurs in the following situations;

Suspended slabs cast between rigid core walls or columns Ground slabs cast on piles The top of infill walls with a low length/height ratio such that the edge restraint from the

base is not effective at the top Large area ground slabs cast onto membranes which are either restrained locally, e.g. by

columns, or by a build up of friction when the area is very large.

Implications for AS 3735AS 3735-2001 (3) provides values for minimum steel ratio pmin for the control of crack widths. For ‘restrained concrete’ Clause 3.2.2 (b) gives values that are related to the bar diameter (Table 6) and according to the Supplement to AS 3735 Supp 1 – 2001 (25) the values have been derived for a mean crack width of 0.15mm

Table 6 Minimum Percentage Reinforcement for Fully Restrained Concrete (Table 3.1 AS 3735)

Bar diameter (mm) 8-12 16 20 24 28 32

pmin (%) 0.48 0.64 0.80 0.96 1.12 1.28

EXAMPLE 2: A 350mm suspended slab (end restraint)

A 1.2m deep slab included an area of reduced thickness to 350mm. The design called for 20mm bars at 175mm, providing 1.03% on each face. This exceeded the 0.8% minimum requirement of AS3735 but for conditions of end restraint the crack width estimated using equ. 1 was >0.5mm.

RestraintRestraint

1.2m thick end wall

1.2m deep beam restrained at end

Crack widths were estimated with increased bar sizes, keeping the same spacing to simplify laps. The wall was internal and a mean crack width of 0.2mm was deemed to be acceptable. C660 indicated that this would be achieved by the use of 28mm bars at 175mm centres (2%) with wk = 0.23mm. This is about double the proposed area of reinforcement and demonstrates the importance of recognising when end restraint exists.

. Working backwards using the approach of BS8007 upon which AS 3735 was based, a maximum allowable temperature drop, T1 may be calculated. To do this it has been assumed that the coefficient of thermal expansion = 12 microstrain/oC; restraint (including the creep coefficient) = 0.5; and tensile strain capacity = 75 microstrain. This leads to an estimated value of T1 = 36oC above which the mean crack width is likely to exceed 0.15mm. This value may be compared with estimated T1 values. The results are shown in Table 7 for walls from 300mm to 500mm thick with cement contents ranging from 300 to 450 kg/m3 and for placing temperatures of 20oC and 30oC.

It is apparent that there are some conditions for which AS3735 may provide insufficient reinforcement, particularly when concreting in the summer months when placing temperatures may be 30oC or more. AS 3735 acknowledges the difficulty in evaluating precisely the amount of reinforcement required due to the numerous and highly variable factors which influence cracking and notesd that in extremely hot climates or for concretes with high cement contents, the recommended minimum steel requirements may be higher than shown in Table 6.

A similar analysis to EN1992 is more difficult as the crack width is also dependent on the cover. Estimates of the maximum T1 values required to ensure a mean surface crack width of ≤ 0.15mm are given in Table 8 for three wall thicknesses, each using a different bar diameter. It can be seen that to maintain a surface crack width of ≤ 0.15mm, lower T1 values are acceptable when there is high cover; or conversely, as shown in Figure 6, the area of reinforcement must be increased.

Table 7 Estimated Values of T1 (OC) using the CIRIA C660 Model for UK Portland Cement

Cement (kg/m3)

Placing temp = 20oC Placing temp = 30oCh = 300mm h = 400mm h = 500mm H = 300mm h = 400mm h = 500mm

300 25 28 31 30 34 37350 28 32 36 35 39 42400 32 37 40 39 44 48450 36 41 45 44 50 54

Table 8 Estimated Maximum Temperature Drop T1 to Achieve a Crack Width ≤ 0.15mm

Cover (mm)

h = 300 mm h = 400mm h = 500mmΦ = 16 mm Φ = 20mm Φ = 24mm

30 43 51 5840 35 42 4850 30 36 4160 29 31 3670 28 28 32

Also to be taken into account is the fact that concrete strengths are now commonly higher than assumed in BS8007 which was developed specifically for concrete with a characteristic cylinder strength of about 30MPa and with an assumed early-age tensile strength of 1.6 MPa. When applied to much higher strength concretes the tensile stress transferred to the steel is proportionally higher and hence more steel is required to maintain the steel stress at an acceptable level.

EXAMPLE 3: Long Deep Slab on Blinding and Compacted Fill

Frictio

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Force to cause slip 1.1m deep slab, 52m x 10m. May appear

to be a continuous edge restraint but it is actually END restraint. The resistance to slip increases with the length of slab until the forces required to cause slip exceed the tensile strength of the section, at which point cracking will occur. When the crack occurs the local frictional resistance is low and does not prevent the crack from opening.

Concrete on compacted fill has a coefficient of friction of 1- 2. Assuming 2, this would require >30m from a free end to develop a tensile stress in excess of 1.5 MPa in the concrete. A 50MPa concrete was used and the early-age tensile strength and the tensile strength would be higher, hence the 52m slab (with 26m to the nearest free end at mid span) was considered to be unlikely to crack due to base friction.0

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The maximum temperature differential through the thickness of the slab is 40oC. The estimated crack width achieved using proposed 28mm bars at 200mm centres was less than 0.1 mm.

To avoid cracking the temperature differential would have to be reduced to below 28oC (assuming the use of a granite aggregate resulting in concrete with a coefficient of thermal expansion of 10 µε/oC.

This may be achieved with the use of insulation with the additional benefit of slowing the rate of cooling enabling the slab to achieve a higher tensile strength, thus marginally reducing the (already low) risk of cracking though base restraint.

SummaryThe introduction of Euro-codes EN1992-1-1 for general design and EN1992-3 for water retaining structures has resulted in some changes in the design process for reinforced concrete in Europe. These changes have implications for the control of early age thermal cracking, previously provided by BS8007. To reflect these, and other changes which have occurred since the publication of CIRIA 91 (revised edition, 1992) CIRIA has published report C660 “Early-age thermal crack control in concrete”. AS3735 has used the approach of BS8007 with modifications to suit Australian conditions but in some circumstances crack widths have been larger than permitted. It may therefore be appropriate to reconsider the recommendation of AS3735 also.

References 1. British Standards Institution, BS 8007:1987, Design of Concrete Structures for Retaining

Aqueous Liquids”.

2. Campbell-Allen, D and Hughes, G W, Reinforcement to Control Thermal and Shrinkage Cracking. Transaction of the Institution of Engineers, Australia, Civil Engineering, August 1981, Vol. CE23. No. 3.

3. Australian Standards, AS3735-2001, “Concrete Structures Retaining Liquids”

4. Harrison, T A, “Early-age Thermal Crack Control in Concrete” CIRIA Report 91, 1992

5. Bamforth, P B, “Early-age Thermal Crack Control in Concrete” CIRIA C660, 2007

6. EN1992-1-1:2004, Eurocode 2. Design of Concrete Structures. General Rules and Rules for Buildings.

7. British Standards Institution, BS8110-2: 1985 “Structural Use of Concrete”.

8. EN1992-3:2006 “Eurocode 2: Design of Concrete Structures – Part 3: Liquid Retaining and Containment Structures”

9. Hughes, B (1971) “Control of Early Age Thermal and Shrinkage Cracking in Restrained Reinforced Concrete Walls” CIRIA Technical Note 21, 1971

10. Dhir, R, Paine, K A and Zheng, L “Design Data for Use where Low Heat Cements are Used” DTI Research Contract No. 39//680, CC2257, University of Dundee, Report No CTU2704, November 2006

11. Anson, M and Rowlinson, P M “Early-age Strain and Temperature Measurements in Concrete Tank Walls ”Magazine of Concrete Research, Vol. 40, No. 145, December 1988

12. Concrete Society “In Situ Strength of Concrete – An Investigation into the Relationship between Core Strength and the Standard Cube Strength” Report of a Working Party of the Concrete Society, Project Report No. 3, 2004

13. Fan, S C Aw, K M and Tan, Y M “Peak Temperature-rise for Early-age Concrete under Tropical Climatic Conditions” Journal of the Institution of Engineers, Singapore, Vol.44, Issue 1, 2004

14. Pigeon, M Bissonnette, B Marchand, J Boliy, D and Barcelo, L “Stress Relaxation of Concrete under Autogenous Early-Age Restrained Shrinkage” American Concrete Institute, Special Publication, SP-227-16, 2005, ACI Detroit Michigan

15. Altoubat, S A and Lange D A “Creep, Shrinkage and Cracking of Restrained Concrete at Early Age” ACI Materials Journal, July/August 2001 Vol. 98, No.4. 323-331

16. Bamforth, P B “Early Age Thermal Cracking in Concrete” Institute of Concrete Technology, Technical Note TN/2, 1982, Camberley, Surrey

17. Vitharana, V and Sakai, K “Early Age Behaviour of Concrete Sections under Strain Induced Loadings” Proceedings of 2nd International Conference on ‘Concrete under severe conditions’ CONSEC ’95, Sapporro, Japan, 2-4 August 1995, Ed K Satai, N Banthai and O E, Gjorv, E&F Spon, 1571-1581

18. ACI Committee 207 “Effect of Restraint, Volume Change and Reinforcement on Cracking of Mass Concrete” ACI Manual of Concrete Practice, Part 1, 207.2R-73 (reapproved 1986), Detroit, Michigan.

19. Tasdemir, M A Lydon, F D and Barr B I G “The Tensile Strain Capacity of Concrete” Magazine of Concrete Research, 1996, 48, No. 176, Sept., 211-218

20. Beeby, W and Forth, J P “Control of Cracking in Walls Restrained Along Their Base Against Early Thermal Movements” University of Dundee, International Congress on Global Construction, Ultimate Concrete Opportunities, 5-7 July, 2005, Thomas Telford pp123-132, ISBN: 0727733877

21. Anchor, R D, Hill, A W and Hughes, B P, “Handbook on BS 5337:1976 (The structural use of concrete for retaining aqueous liquids)” Viewpoint. Publications, Cement & Concrete Association, Slough. 1979

22. Narayanan, R S and Beeby A W, “Designers’ Guide to EN 1992-1-1 and EN 1992-1-2 Eurocode 2: Design of Concrete Structures. General rules and rules for buildings and structural fire design” Thomas Telford.

23. Beeby, A W “Fixings in cracked concrete – The Probability of Coincident Occurrence and Likely Crack Width” CIRIA Technical Note 136, 1990

24. Bamforth, P B, Price, W F and Emerson, M “An International Review of Chloride Ingress into Structural Concrete” Contractor Report 359, TRL Scotland, 1997

25. Australian Standards, AS3735 Supp 1 -2001, “Concrete Structures Retaining Liquids – commentary (Supplement to AS3735-2001)”