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THERMAL CRACKING IN LARGE CONCRETE PLACEMENTS: THEORY AND APPLICATIONS

ANCOLD 2002 Conference on Dams Page 1

THERMAL CRACKING IN LARGE CONCRETE PLACEMENTS:

THEORY AND APPLICATIONS

N. Vitharana1

and R. Wark2

ABSTRACT

Large concrete placements such as those encountered in dam construction are subjected to severe stress

conditions at early age due to the heat generated in the process of hydration of cement. These thermal stresses

can be higher than those experienced by the structures during its service life. Being young concrete, its tensile

strength is much lower than that of hardened concrete. Consequently if measures are not taken, thermal stresses

could lead to the cracking of concrete.

In the traditional approaches, the criterion of limiting the maximum temperature rise (in some cases both

average and differential) is specified as the sole criterion to avoid thermal cracking. However, practical

experiences has shown this approach to be superficial, being either conservative or unconservative dependingon the conditions. This is due to the fact that thermal cracking occurs when thermal stresses exceed the current

tensile strength of the concrete and accordingly temperature limits have no direct relevance. Also, traditional

deemed-to-satisfy criterion of limited maximum temperature rise, based on experience(s) at past projects, would

not be valid for todays conditions where cement types and construction techniques are very different.

Consequently, some modern design and construction practices, such as Japanese standards, permit or require

designers and contractors to develop their own procedures/criteria with respect to thermal crack control in

concrete structures.

A method involving the calculation of a thermal cracking index (ratio of thermal stress/tensile strength) would

be a better and more rational approach. However, this method needs to consider structural, hydration, material,

thermal and exposure parameters. As early-age concrete is in a semi-plastic state with the involvement of many

interactive parameters such as creep, temperature, ambient conditions, strength gain etc., the evaluation of the

thermal cracking index entails complex procedures.

This paper presents the details of a numerical procedure (THERMAL) developed to predict the time-history of

thermal cracking index. Examples are also presented to show where this procedure has been successfully

applied.

Keywords: Thermal stresses, heat-of-hydration, concrete placements, creep, young concrete, strength gain.

1Principal Engineer, GHD Pty Ltd, BSc(Eng)Hons, PhD(Struct), MBA, PG-Dip(Geotech), MASCE

2Technical Director (Dams), GHD Pty Ltd, BEng(Hons),BAppSc(Maths), MEngSc, FIE(Aust)

1 INTRODUCTION

Hydration-induced cracking in concretestructures, particularly in dams and water-

retaining structures, could result in multifariouseffects: accelerated corrosion of reinforcing andprestressing steel, unsightly appearance, loss of

water-tightness, accelerated deterioration due tofreezing-thawing and alkali-aggregate reaction,development of hydrostatic pressures insidecracks in dams, impairment of structuralintegrity, stability and load redistributions. Thenegligence or cursory treatment of early-agestresses could incur heavy costs requiringextensive repair work or even the total early

replacement of structures such as that occurred

with Kinzua dam stilling basin [HOLLAND,1991].

Engineers are well conversant with the analysisand design of structures against applied loadssuch as dead and live loads because appropriateanalysis methods and design criteria are welldocumented. Deformation-induced loadings

such as temperature and shrinkage are treatedwith rule-of-thumb or deemed-to-satisfy

approaches, if considered at all. The adoptionof a minimum reinforcing steel ratio (AS 3600)to avoid early-age thermal cracking in massivestructures is questionable as bond transfer inyoung concrete is very limited [VITHARANA,

1995c].


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A popular method in tender specifications forthe construction of dams and large structures is

to place limits on the temperature rise anddifferential. This is achieved by reducing theconcrete placing temperature, internal andexternal cooling, multi-lift concrete placementetc. In recent decades, the introduction ofdifferent cement types and constructionmethods requires the validity of thesetraditional approaches to be assessed.Moreover, the cracking of young concreteoccurs when the thermal stresses exceed theavailable tensile strength at a given timefollowing the casting of concrete. Therefore,the adoption of the traditional method of limited

temperature rise could lead to either superfluous

or inadequate measures to guard against thermalcracking of concrete structures.

The tendency to cracking at an early age inconcrete structures is determined by the ratio of

thermal stress/tensile strength of the hardeningconcrete. The generation of thermal stresses

depends on many factors, mainly: cement typeand content, placing temperature, ambientconditions, concrete material properties,maturity of concrete, structure type,construction sequence and restrained

conditions, creep characteristics etc. Unlike thecase of applied loadings, thermal stresses are

relaxed by the creep effects (non-elasticadditional strains under sustained stresses)which are high in young concrete due to itssemi-plastic state. Also, it is important toconsider construction sequences as this wouldmodify both the thermal and structuralresponses of a structure. Therefore, theevaluation of thermal crack occurrence at early

ages entails a complex procedure. In recenttimes, constitutive models for young concrete

have been developed [THERMAL, 1995]although there is a great need to understandtheir behaviour at fundamental levels.

During 1994-1996, the first Author was on anindustrial fellowship at the HokkaidoDevelopment Bureau, in the most NorthernIsland of Japan. This organisation isresponsible for providing construction advice toall construction activities on the Island.Consequently, it has been conducting research

and development activities in conjunction withmajor contractors particularly for cold weather

conditions as winter temperatures can be as lowas 40

oC and construction periods can be as

short as six months. It has been developing

thermal stress prediction models for RCC damsas early as the 1970s. The temperature and

stress development predicted for a 70m highRCC dam with interruption for the winter isshown in Fig. 1. With such models, it is easierto observe the effect of various parameters anddetermine appropriate measures to be takenwhen construction conditions are varying. TheAuthor was involved in the development ofconstitutive models for young concrete(including large scale tensile creep tests for damconcrete) and a methodology for predicting thetendency to cracking under heat-of-hydration,drying and autogenous shrinkage. Thecomputer program developed [THERMAL,

1995] is suitable for 3-dimensional structures.

(a) Temperature prediction

(b) Thermal stress prediction

Fig. 1 Temperature and stress predictions for a70 m high RCC dam

2 MECHANISM BEHIND

EARLY-AGE THERMAL

CRACKING

Many interactive factors are involved in the

early-age thermal cracking of concretestructures (eg, hydration, thermal, material,

environmental and structural). Therefore, it isdifficult to make firm conclusions on theinfluence of each factor as they depend on a


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given situation. The major difference betweenapplied and thermal loadings is that thermal

loading depends on the member stiffness (ie,Youngs modulus of elasticity). Unfortunately,this has not been recognised in the design ofreinforced concrete structures subject toambient thermal loading where the structuralengineers have been misled and considerthermal stresses under ultimate limit stateconditions in which the member stiffness isvery small (being in a plastic stage) resulting innegligible thermal stresses [VITHARANA,1998].

During the hydration phase of concrete, semi-plastic concrete has a very low value forYoungs modulus of elasticity. The

incremental compressive thermal stress, due totemperature change Tt within a time step T,is given by: t = Et Tt where is thecoefficient of expansion for concrete and Et isYoungs modulus of concrete at an age t sincecasting concrete. This stress is further reducedby the concurrent creep which is high for young

concrete. The temperature rise takes placewithin the first few days and the resulting

compressive stresses would be about 1 MPa intypical applications. Once the rate of hydration

retards, the temperature begins to drop (orcooling phase) depending on the rate at whichheat is lost to the surrounding. The temperaturedrop takes place under an increased Et valueand the net stress therefore becomes tensile. Ifthis tensile stress exceeds the available tensilestrength, cracking would occur. Thismechanism is schematically shown in Fig. 2.For comparison, the development of thermalstresses under the same temperature cycle inmatured concrete with a constant E value is alsoshown. As can be seen with constant E, the

induced stress at the end of the temperaturecycle would return to zero.

This highlights that if an accurate assessment ofthe tendency to cracking is required, it isnecessary to consider both the heat generatingcharacteristics as well as the material and

strength properties of young concrete. Inparticular, creep effects should be considered as

it would reduce the magnitude of initialcompressive stresses as well as the subsequenttensile stresses.

Fig. 2 Mechanism behind early-age cracking

(a) Temperature and Youngs modulus

(b) Tensile stress and tensile strength

3 HEAT-OF-HYDRATION

In recent decades, finite element methods have

been developed for heat-transfer analysis and

structural analysis with various levels of

sophistication from simple elastic methods to

non-linear methods [CRICHTON, 1999].

However, little attention has been paid to the

use of appropriate hydration models. The

generation of thermal stresses depends on both

the rate and the total heat generated during the

hydration process. In most thermal stress

modellings, an adiabatic hydration model (in

which heat transfer is not lost to/or gained fromthe surrounding) is used in line with the

traditional approaches. Typical adiabatic

temperature generation curves are given in

[ACI207, 1973] for different placing

temperatures (low to high-heat cements) used

prior to the 1960s and these had been used often

in temperature predictions.

Hydration of cement is a thermally-activated

process and consequently the rate of cement

hydration depends on the reaction temperature

(ie, at which the reaction takes place). Thereaction temperature in turn depends on the

heat-transfer characteristics within the concrete


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mass. Therefore, heat-of-hydration and heat-

transfer processes should be coupled in which

the time-history of the reaction temperature is

taken into account and such analysis is known

as coupled or non-linear heat-of-hydrationanalysis [HARADA, 1991]. As will be shown

later, the arbitrary use of adiabatic hydration

models could result in significant errors in

predicted temperatures.

3.1 HYDRATION

CHARACTERISTICS

3.1.1 Adiabatic models

JCI[1986] covers a wide range of cement typesand blend ratios and parameters are given for

developing adiabatic curves for different

cements placed at different temperatures.

These models were calibrated against several

hundred concrete placements throughout Japan.

The adiabatic temperature rise of concrete can

be generally described by:

)1()()(t

t eTT

= (1)

where T(t) = temperature rise (oC) at time t since

casting concrete, T() = total (ultimate)temperature rise directly proportional to the unit

cement content S (kg/m3

of concrete), and =hydration constant representing the rate of

hydration dependent only on the concrete

placing temperature To for a given cement type.

The cumulative heat generated Q(t) (J/m3) up to

time t, for a given cement content S is given by:

)()( tt TcQ = (2)

where c = specific heat of concrete (J/kg/oC),

and = mass density of concrete (kg/m3

). Therate of heat generation, which is to be used in

transient heat-transfer analysis, Qo (J/m3/s) is

given by:

)()()()( t

to

TTceTcQ ==

(3)

Typical values of and T() for differentcement types (eg, ordinary Portland cement,

moderate-heat Portland, Portland fly ash,

Portland blast-furnace slag, and high-strength

concrete) are given in [JCI, 1986] and these are

very useful for design and constructionengineers to make a preliminary assessment of

the heat generating characteristics of a given

concrete mix design. Heat-of-hydration

adiabatic calorimeter tests or large concrete

blocks with insulated surfaces can alternatively

be used when an accurate evaluation of the

parameters is required.

3.1.2 Non-adiabatic models

Adiabatic hydration models give the fastest rate

of heat generation as they are based on the

implicit criterion that the heat already produced

by hydration accelerates this hydration process

in turn. Although this may be valid to the

interior of massive structures such as dams, it

would become invalid near the surfaces of

massive structures or in other members such as

walls, foundations etc.Non-adiabatic hydration models should

consider the time-history of the reaction

(process) temperature. As the reaction

temperature varies within a structural member

due to the concurrent heat-transfer, the

hydration process at a given time varies within

the structural member. In recent times,

sophisticated hydration models have been

developed and tested [HARADA, 1991], but

they are not suitable for routine applications. A

simplified model can be developed from themethod suggested by RASTRUP [1954] based

on heat measurement on cement samples under

constant-temperature (isothermal) conditions

(ie, heat-of solution tests). Supplemented by

laboratory testing and field measurements, it

was shown that [VITHARANA, 1995c] this

model can be easily calibrated and incorporated

in heat-transfer analysis.

With this model, the reaction temperature is

taken into account by relating the actual

reaction time t to an equivalent (or maturity)age te which is determined based on the time-

history of the reaction temperature. The

cumulative heat generated Q(t) (kJ/kg of

cement) up to the reaction time t (days) is

expressed as:

}][exp{)(n

et tmBAQ+=

dtt rTT

e = )(1.02 where A, B, m and n are hydration constants

dependent on the cement type and m alsodepends on Tr (the reference temperature). As

can be seen, the actual reaction time t is


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converted to the equivalent time te with the

known reaction temperatures from the heat-

transfer analysis. With this approach, the

hydration and heat-transfer process are coupled

implicitly. The rate of heat generation can beobtained numerically from the above equation

within a given time increment. Typical values

for hydration constants were developed

[VITHARANA, 1995c], and for normal

Portland cement: A=12.56, B=328.7 (both in

kJ/kg of cement) and n=0.42 and m=-1.029 for

the reference temperature Tr= 20oC.

Fig. 3a shows the adiabatic and isothermal heat

generation characteristics for different placing

temperatures for the same concrete mix. The

heat generation within a concrete mass lies inbetween these two extreme conditions. In order

to highlight the significance of the inaccuracy

in using adiabatic hydration models, Fig 3b

shows the typical thermal stress development in

a 200mm and 600mm thick reinforced concrete

walls in moderate ambient conditions

[VITHARANA, 1995b], an overestimation of

the tensile stresses by about 40% .

Fig. 3a Adiabatic and isothermal heat generation

Fig. 3b Thermal stress development with adiabatic

and non-adiabatic models

A general disadvantage with this model is the

difficulty in conducting isothermal heat-of-

solution tests. However, adiabatic tests are easy

to perform and test results are widely available

for different placing temperatures [JCI,ACI207]. Therefore, an indirect method was

developed to synthesise the hydration curves

for varying temperature curves [VITHARANA,

1995b]. This method ignores the temperature-

time history as a direct variable, but the rate of

heat generation is related to the cumulative heat

already generated by adjusting the placing

temperature with the known reaction

temperature at a given time. This was shown to

provide accurate temperature predictions for

non-adiabatic environments.

4 HEAT-TRANSFER

ANALYSIS

Both finite difference and finite element

methods are well established and standard

analysis procedures are available. Finite

difference methods can be formulated easily

and can be implemented even in an Excel

spreadsheet. It is very important to consider the

heat transfer boundaries, particularly the wind

speed and direct solar radiation. The insulationeffect provided by wood formwork should be

considered and sudden removal can generate

surface thermal stresses significant enough to

cause surface cracking. Due to space

limitation, these will not be discussed in detail

here.

Schmidts method is used [ACI207] by design

and construction engineers to predict

temperature rise in concrete placements.

Although this method is useful for preliminary

estimates of thermal stresses, there are severalimplicit assumptions that do not allow an

accurate assessment of temperatures. It

assumes that the surface temperature is equal to

the ambient temperature and field

measurements show that this is not true due to

heat-transfer resistance along the boundaries.

In addition, it uses adiabatic hydration models.

5 MATERIAL PARAMETERS

The development of material and strength

properties in young concrete depends not onlyon the age (time since casting concrete) but also

on the reaction temperature which varies with


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time as well as within the structural element

due to the transient heat-transfer. Although the

age has been considered using measured data or

maturity functions [ACI209, CEB] in recent

thermal stress predictions, the influence ofreaction temperature has received less attention

although strength development is also a

thermally-activated process.

5.1 TENSILE STRENGTH AND

YOUNGS MODULUS

The development of tensile strength can be

related to the compressive strength with good

accuracy using the ACI209[1986] strength

development function. The tensile strength ft at

age t (days) at a standard curing temperature of

20 oC is given by:

where ft(28) is the tensile strength at the age of

28 days, and A and B are material constants

dependent of the cement type, curing method,

admixtures etc. Typical values of A and B are:

A=4.0 and B=0.85 for normal-strength concrete

and A=2.30 and B=0.92 high-strength concrete.

The Youngs modulus of elasticity E at age t

can also be related to the corresponding value

of E at 28 days with the above function.

Similar to the heat of hydration, the strength

gain is much faster at elevated temperatures and

this should be considered in the determination

of the thermal cracking index. The major

implication is that the strength gain near the

surface would be much slower compared with

the interior.

The temperature-dependency of ft and E can be

incorporated by using an equivalent (ormaturity) age te which considers the time-

history of the curing temperature. The CEB-

FIP[1991] formulations can be used for this

purpose. The development of compressive and

tensile strength of high-strength concrete in

standard (20 oC) and adiabatic conditions is

shown in Fig. 4. As can be seen, within the

first 3-4 days, the strength gain under adiabatic

condition is about 30% higher than that under

the standard temperature. This highlights the

fact that the purpose of undertaking detailed

thermal stress analysis would be lost if the

development of basic material properties is not

considered rationally.

Fig. 4 Compressive and tensile strength

development under adiabatic and constant

temperatures

5.2 CREEP BEHAVIOUR

The analysis of thermal stresses due to heat-of-

hydration should be carried out in an

incremental fashion in time steps (Section 6).

With known thermal strain developed during a

particular time step, the stress can be calculated

with the appropriate Youngs modulus (Section

3). However, the time-dependent deformational

behaviour of concrete should be considered as

this reduces the magnitude of thermal stresses,usually known as creep-relaxation. In the

analysis procedure developed in THERMAL,

the principle of superposition is used to take

into account the stress-history. Separate creep

factors and characteristics are developed for the

incremental stress developed at each point in

the concrete section during each time step.

It is also important to consider effect of

temperature, before and after a particular

thermal stress is generated, on the creep

behaviour of concrete. The elevatedtemperature generated, before stressing,

determines the effective age te (calculated in

Section 3) thus reducing the creep strain. After

stressing, creep is accelerated with elevated

temperature. Based on an experimental

program [VITHARANA, 1995c], it was

concluded that reasonable accuracy can be

obtained by using the formulations given in

CEB-FIP with some modifications to reflect on

the very early-age concrete behaviour. The

total (elastic and creep) strain (t) at age t undera stress applied at age t = to is given by:

)28(

5.0

tt fBtA

tf

+

=


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)(

),()( ]1[toc

tototE

+=

where Ec(to) = Youngs modulus of elasticity at

age to (considering maturity due to reactiontemperature prior to to), o = ultimate creepfactor and (t,to) = function defining time-dependency of creep strain since to. The

ultimate creep value is proportional to the factor

k which defines the age at stressing and the

function (t,to) is given by:

]1.0[

12.0

otk

+=

3.0

),(

+= ta

ttot

6 THERMAL STRESS

CALCULATION

As the analysis is to be carried out in time steps

in an incremental fashion, free strains are to be

determined for each element/node before they

are introduced as induced strains to the

structural analysis program, either a FEM or a

simple beam analysis. THERMAL thencalculates the stresses depending on the external

restraints/boundary conditions of the structure.

FEM analysis would be time consuming and

not always necessary. Simple beam-theory

based calculations can be performed to obtain

accurate results (Section 7). The steps involved

in determining thermal stresses are:

Calculate the incremental thermal strainDt for each point within the concretesection developed during time t and t+t

(ie, within time step n): Dt = [T(t+t)-T(t)].Other induced-strains such as drying and

autogenous shrinkage can also be included.

Calculate the incremental creep strainDcr,to occurring within time step n due to aunit stress ( = 1) applied at an age to:

{ })())(

, tottott

toc

otocr

ED + =

Determine the algebraic sum of the

incremental creep strains Dcr fromincremental stresses generated at the 1sttime step to the current step n.

Calculate incremental thermal stress D(t)developed during time t and t+t. In FEManalysis, strains can be used as the input

directly. The total stress (t) at a given

time is given by the algebraic sum of D(t)up to the current time step n.

crtttttct DTTERFD = ++ )()()()(

The factor RF represents the degree ofrestraint on the free induced strain if this is

known for the structure based on previous

elastic analysis. Axial and flexural

restraint factors RF are given in

[JCI(1986), ACI207 (1973)].

The tendency to cracking is determined bycalculating the Thermal Cracking Index

ratio: Tci = (t)/ft , where ft is the availabletensile strength in hardening concrete. The

time-history of this ratio can be prepared

for the critical points in the structure and

then the probability of cracking can be

determined. Probability of cracking vs Tci

is given in JCI(1986) based extensive field

data. This is a very useful tool for

estimating the probability of cracking for

different concrete mix designs particularly

for comparing temperature controlmeasures and the cost involved.

Thermal stresses are generated in concrete

sections even if they are unrestrained externally

or by structural indeterminacy. These

unrestrained stresses are known as primary

stresses and are caused by non-linear

temperature distributions. Secondary stresses

are generated due to the axial and flexural

restraints. It would be worthwhile and

economical to understand the structure

behaviour in simple axial and flexural actions inconjunction with appropriate restraint factors

before undertaking complex finite element

analysis (eg, walls, massive foundations, box-

girder bridges, dams etc). The influence of

degree of restraint on thermal stresses can also

be used to compare different concrete mix

designs [VITHARANA, 1998] before

specifying them for a particular project.

A typical case for a wall is shown in Fig. 5 in

which unrestrained primary stresses are set up

in the vertical direction while restrained(primary and secondary) stresses are set up in

the horizontal direction due to base fixity.


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0

10

20

30

40

50

60

0 10 20 30 40 50 60 70 80

Time (days)

Temperature(C)

Base-measured

Base-predicted

Mid-measured

Mid-predicted

These can be calculated based on the usual

plane sections remain plane hypothesis. Similar

approach has been used to predict stresses in

multi-lift foundation construction (Section 7)

founded on different foundation conditions.

Fig. 5 Thermal stress generation in a wall section

7 APPLICATIONS

It is difficult to firm general conclusions /

recommendations on the development of

thermal stresses due to the involvement of

many factors including ambient conditions. In

the previous sections, the major aspects were

outlined with brief discussions on their

influence(s) on the development of thermal

stresses and tendency to cracking.

7.1 TOKACHI BRIDGE BASE

The spread foundation for this 750 m long

bridge was constructed in the winter conditionswith minimum daily temperatures varying

between 20 to 30 oC. With a plan area of

27mx32m and a thickness of 6m, this was the

second largest substructure of this type ever

constructed in Japan [ISHIKAWA, 1994]. The

construction was carried out in 3-lifts of 2m

thick pours with a 14-day gap between each

pour to permit the dissipation of hydration-

induced heat. Artificial surface heating and

surface insulation was undertaken to prevent

high temperature differentials. The cementcontent used is 280 kg/m

3with normal Portland

cement. During construction, temperature and

strains were measured continuously up to 70

days.

THERMAL was used to compare the

temperatures monitored and to predict thermal

stresses including creep relaxation. Inparticular, it was required to determine whether

the time gap between successive pours (which

is significant compared with the available

construction time) helped reducing thermal

stresses generated compared with a single pour.

In an earlier analysis [ISHIKAWA, 1994], the

measured strains have been converted to

thermal stresses using an effective modulus to

account for the creep relaxation and therefore it

may not be accurate as it neglects early-age

non-linear creep.The temperature predicted by THERMAL is in

close agreement with those measured, Fig. 6.

In particular, the thermal stresses predicted by

THERMAL highlight the importance of

considering the creep-relaxation. The tensile

strength development included the effect of

varying reaction temperatures during the

hardening process. It is interesting to note that

1-dimensional heat transfer and beam analyses

were also undertaken (with zero axial restraint

on the axial deformation) and the predictionswere accurate within 5-10% compared with

finite element analysis results. Therefore, it is

strongly recommended that simple analysis

procedures be undertaken, wherever possible,

which may take a fraction of the time and

resources required for finite element analysis.

Fig. 6a Tockachi Bridge: Temperature in 6 m

thick base


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-3

-2

-1

0

1

2

3

4

0 10 20 30 40Time (days)T

ensilestress(MPa)

Base-measured

Base-predicted

Mid-measured

Mid-predicted

Tensile strength

Fig. 6b Tockachi Bridge: Stress generation in 6 m

thick base

7.2 IS MULTI-LIFT

CONSTRUCTION ALWAYS

HELPFUL ?

It is generally believed that by lowering the

average temperature rise, the tendency to

cracking can be minimised. One such method

adopted in concreting for thick foundations is

multi-lift (or multi-layer) concrete placement

with a time gap between each lift [GYRAX,

1994]. A range of parameters was consideredfor foundation thicknesses varying from 1-5m

for typical Hokkaido weather conditions

[VITHARANA, 1995a].

Fig. 7 shows the temperature and thermal

stresses predicted for single and multi-lift

constructions with a 3-day time gap between

each 1-m thick concrete placement. As can be

seen, although temperature is lower with the

multi-lift construction, the induced thermal

stresses are not lower and single-lift concrete

can be undertaken without an increased risk ofthermal cracking. This is due to the fact that

stiffnesses of layers are different in multi-lift

construction thus developing internal restraints

within the structure. The influence of creep is

significant with a stress reduction of about 45-

50%. Following this study, several single-lift

concrete placements up to 5m thick were

undertaken without any thermal cracking.

Fig. 7 Temperature and stress developments in single and multi-lift foundations


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0.0

1.0

2.0

3.0

4.0

0 2 4 6 8 10

Time (days)

Stressandstrength(MPa)

0

10

20

30

40

50

Temperature(C)

Tensile strength (MPa)

Tensile stress (MPa)

Temperature (C)

Measured tem (C)

7.3 CANNING DAM ANCHOR

BLOCK

The anchor block for the Canning dam post-

tensioned cables is about 3 m thick [WARK,

2001]. The concrete mix design was based on

previous similar dam constructions (North

Dandalup and Lower South Dandalup) and

extensive laboratory tests undertaken to

minimise the tendency to alkali-aggregate-

reaction. The selected mix had a unit cement

content of 370 kg/m3

(normal Portland cement

= 190, Fly ash = 150 and Silica fume = 30).

With a high Fly ash content, the strength gain

was slower with a 180-day characteristics

compressive strength of 50 MPa (comparedwith the 28-day strength of 31 MPa).

According to the original construction

programme, concreting was to be done during

the summer months and the anticipated main

issue would be the control temperature rise.

The construction specification therefore placed

stringent conditions: a placing temperature of

20oC and a limit of 30

oC on the maximum

temperature rise in the pour. However, delay to

the construction programme required the

concrete placement to be undertaken during thewinter months and there were concerns that the

construction methodology would be unable to

limit the temperature differential to 20oC as per

the specifications.

THERMAL was used to simulate the

temperature monitored in Block No. 20 and

then to predict thermal stresses and the

tendency to cracking. There were also concerns

that uplifting at the old-new interface may

occur with a high differential temperature

gradient. The results at a point 200mm belowthe top of the block are shown in Fig. 8. As can

be seen, the thermal stresses were well below

the tensile strength. As an additional measure,

a couple of layers of ceiling insulation were

used during cooler nights as an insulation layer.

THERMAL considered the influence of creep-

relaxation and further analyses showed that the

predicted stresses would have been higher than

those shown in Fig. 8 by about 30% if creep

effects had been neglected.

Fig. 8 Temperature, stress and strength

development in Canning Dam anchor block (3 m

thick)

7.4 HARVEY DAM INTAKE TOWERFOUNDATION

The foundation for the Harvey Dam intake

tower has a diameter of 20m and a thickness of

2m. The foundation is founded on rock with

grouted dowels for enhancing its stability under

extreme loads. The foundation is provided with

reinforcement to resist the lateral loadings

induced by earthquakes and minimum steel for

crack control.

The total cementitious content was 370 kg/m3

(normal Portland cement =50%, Fly ash =40%and Silica Fume =10%). Temperature

development was monitored during the

construction at several locations. The

construction was carried out in March 2001

when maximum and minimum ambient

temperatures were 32 and 10oC respectively.

THERMAL was used to develop the heat-of-

hydration and heat-transfer in a 1-dimensional

analysis. The results are shown in Fig. 9 for the

mid-thickness of the foundation, comparing

with measured temperatures and they are inexcellent agreement.


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0

10

20

30

40

50

60

0 24 48 72 96 120 144

Time (hrs)

Temperature(oC)

Ambient

Measured

Predicted

Fig. 9 Temperature prediction for Harvey Dam

intake tower foundation (2 m thick)

8 CONCLUSIONS

Heat-of-hydration in hardening concrete would

be severe enough to cause cracking in concrete

placements. If accurate results are to be

obtained, correct modelling of the hydration

process is as important as heat-transfer analysis.

The traditional criterion of limiting the

maximum and differential temperature rise

would not always be correct. The primary aim

should be to determine the thermal stresses and

then compare with the strength gain. Stress-

relaxation due to early-age creep is significant.

As shown with practical applications,

THERMAL can be used to predict the

developments of temperature and thermal

stresses with good accuracy.

9 ACKNOWLEDGEMENTS

The first Author would like to thank the staff at

the Hokkaido Development Bureau, Sapporo,

Japan for all the assistance given during his stay

and Dr K Sakai (Head, Structural Engineering)

and Dr N Sato (Head, Dams Engineering) for

guidance and advice on numerous occasions

and financial support. Mr Jonathan Jensen is

thanked for formatting this paper.

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