Engineering Structures 62–63 (2014) 148–163

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Engineering Structures

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Application of air cooled pipes for reduction of early age cracking riskin a massive RC wall

http://dx.doi.org/10.1016/j.engstruct.2014.01.0180141-0296/� 2014 Elsevier Ltd. All rights reserved.

⇑ Corresponding author. Tel.: +351 938404554; fax: +351 253 510 217.E-mail addresses: miguel.azenha@civil.uminho.pt (M. Azenha), rmlameiras@

civil.uminho.pt (R. Lameiras), christoph@civil.uminho.pt (C. de Sousa), barros@civil.uminho.pt (J. Barros).

Miguel Azenha ⇑, Rodrigo Lameiras, Christoph de Sousa, Joaquim BarrosISISE – Institute for Sustainability and Innovation in Structural Engineering, University of Minho, School of Engineering, Civil Engineering Dept., Azurém Campus, 4800-058Guimarães, Portugal

a r t i c l e i n f o a b s t r a c t

Article history:Received 27 May 2013Revised 2 December 2013Accepted 13 January 2014Available online 15 February 2014

Keywords:Cement hydrationService life conditionsThermal shrinkageCrackingNumerical simulation

The construction of massive concrete structures is often conditioned by the necessity of phasing castingoperations in order to avoid excessive heat accumulation due to cement hydration. To accelerate con-struction and allow larger casting stages (usually increasing lift height), it is usual to adopt internal cool-ing strategies based on embedding water pipes into concrete, through which water is circulated tominimize temperature development. The present paper reports the use of horizontally placed ventilatedprestressing ducts embedded in a massive concrete wall for the same purpose, in line with a preliminarySwedish proposal made in the 1990s. The application herein reported is a holistic approach to the prob-lem under study, encompassing extensive laboratory characterization of the materials (including a tech-nique developed for continuous monitoring of concrete E-modulus since casting), in situ monitoring oftemperatures and strains, and 3D thermo-mechanical simulation using the finite element method. Basedon the monitored/simulated results, it is concluded that the air-cooling system is feasible and can effec-tively reduce early cracking risk of concrete, provided adequate planning measures are taken.

� 2014 Elsevier Ltd. All rights reserved.

1. Introduction

The combined effect of the exothermic nature of cement hydra-tion reactions and the relatively low thermal diffusivity of concreteleads concrete structures to endure temperature increases at earlyages, and eventually return to thermal equilibrium with the sur-rounding environment. These early temperature variations inducevolumetric changes in concrete that are partially restrained byadjoining previously cast members, or even due to non-uniformtemperature distributions within a concrete member itself. Suchrestraint to deformation may induce stresses that can be relevantenough to induce early age thermal cracking in concrete, whichis usually unacceptable in view of aesthetics, durability and evenstructural performance. Contractors usually attempt to avoid thisthermal cracking by adopting concrete compositions and construc-tion schedules that maintain temperature gradients in concrete be-low prescribed limits, both along time and space [1]. It hashowever been recognized that such approach leads to erroneousconclusions, as several important issues are disregarded [1], such

as the degree of restraint to deformation and the actual mechanicalproperties of concrete. In view of the limitations of these temper-ature-based criteria, it has been widely acknowledged [2] thatmore realistic crack risk assessments can be made through multi-physics approaches that encompass numerical simulation of tem-peratures and corresponding stresses in concrete: thermo-mechanical analyses. The use of thermo-mechanical simulationmodels allows the evaluation of alternative construction scenarios(for casting procedures, concrete mixes, environmental condi-tions), and thus permits the optimization of construction withoutcompromising the safety in regard to thermal cracking. The numer-ical studies for the assessment of the optimum construction strat-egy frequently involve diminishing the temperature rise inconcrete at early ages. In fact, if the thermal variation is dimin-ished, the corresponding volumetric changes also decrease, as wellas the developed stresses. The diminishment of early temperaturerises is usually achieved by partial replacement of cement by addi-tions as fly ash [3,4], or by cooling water/aggregates before mixingoperations [5–7], or even by introducing internal cooling pipes inconcrete with cooling fluids (usually water) [8–12]. An attemptto use air as the cooling fluid in cooling pipes has been made inthe 1990s by Hedlund and Groth [8,9], who have shown the feasi-bility of such technique in thick columns. Nonetheless, no furtherapplication of such technique was found in the literature, except

M. Azenha et al. / Engineering Structures 62–63 (2014) 148–163 149

for an in situ use of air for localized cooling reported by Ishikawaet al. [12].

Even though thermo-mechanical simulation approaches havebeen applied to concrete structures for decades [13–17], they in-volve several complexities and problematic issues particularly inview of the assessment of material properties and model parame-ters (thermal and mechanical), which frequently demand specificlaboratory characterization: heat of hydration, thermal boundarycoefficients, creep of concrete, evolution of E-modulus and tensilestress, among others. Even though several scientific works havebeen done either on thermal simulation [18], thermo-mechanicalsimulation [19] or monitoring the concrete behaviour at early ages[20–23], some combine the numerical simulation with experimen-tal data obtained in laboratory [24–26] or with temperature mon-itoring for partial validation [27–31], whereas others go furtherand additionally include in situ strain monitoring for validation[32]. Nonetheless, in the scope of internal cooling of concretethrough the use of embedded pipes, no works were found to adoptholistic approaches that simultaneously include material charac-terization, in situ monitoring of temperature/strain and thermo-mechanical simulation. The works that focus on concrete coolingwith embedded pipes are mostly limited to thermal [33], or ther-mo-mechanical analyses only [34–37], and thermal [38–40] orthermo-mechanical analyses together with partial validationthrough in situ monitoring [12,41].

The present paper pertains to a case study of the thermal stres-ses in the central wall of the entrance organ of a dam spillway.Such wall is 27.5 m long, with a maximum width of 2.8 m andheight of 15.0 m, with attention being given to the most unfavour-able construction phase in which a 2.5 m tall batch had to be made(total 150 m3 of concrete). Due to the materials and equipmentavailable at the construction site, it was decided to attempt inter-nal cooling of concrete with air-cooled prestressing ducts placedlongitudinally along the wall.

The present paper regards to the in-depth study of the earlyage performance of concrete in the wall, encompassing labora-tory thermal and mechanical characterization of concrete, aswell as in situ monitoring of temperatures/strains and the corre-sponding thermo-mechanical simulation with the finite elementmethod.

The extensive laboratory characterization included quantifica-tion of the heat of hydration evolution (isothermal and semi-adia-batic calorimetry), evaluation of compressive/tensile strength andE-modulus through cube/cylinder testing, creep testing at severalages and assessment of mechanical activation energy. In particularregard to E-modulus testing, a methodology that allowscontinuous measurement of concrete E-modulus since casting(EMM-ARM [42]) was applied. This is a pioneering use of thismethodology for the purpose of supporting stress simulation onconcrete since its very early ages, with important advantages inregard to previous approaches that tend to extrapolate values ofE-modulus at very early ages.

A relatively complete monitoring program has been carried outin situ, involving the use of 20 temperature sensors and 7 vibratingwire strain gauges embedded in concrete. Particular attention wasgiven to the evaluation of the effectiveness of the cooling system,with temperatures being measured at several points along the pre-stressing ducts, and with air velocity measurements taken withhandheld anemometers.

Bearing in mind the information gathered with the laboratorycharacterization and in situ monitoring, a thermo-mechanicalsimulation was carried out with recourse to a three dimensionalfinite element model. Such simulation model included the explicitmodelling of the cooling ducts, as well as the phased constructionof the wall. The simulation model was made with DIANA software[43].

2. Thermo-mechanical model

The thermo-mechanical simulation approach presented herehas strong similarities with that described in a previous work[30]. Nonetheless, some particularities are distinct, namely: (i) so-lar radiation is explicitly considered according to a model based onthe incidence angle of the sun beams; (ii) the effect of internalcooling ducts is taken into account (iii) the evolution of mechanicalproperties is simulated according to the ‘equivalent age concept’(instead of the degree of hydration concept). The following sub-sections pertain to a general description of the modelling strategywith specific emphasis on topics (i)–(iii) mentioned above.

2.1. Thermal model

The calculation of temperature fields in concrete is based on theheat balance equation, whose solution is made through the finiteelement method [44]:

kr � ðrTÞ þ _Q ¼ qc _T ð1Þ

where k is the thermal conductivity, qc is the volumetric specificheat and T is the temperature. _Q is the volumetric heat generationrate due to cement hydration, formulated as an Arrhenius type law[45]:

_Q ¼ Af ðaÞe�EaRT ð2Þ

where A is a rate constant, Ea is the apparent activation energy, a isthe degree of heat development (ratio between the heat Q releasedup to time t and the total heat Qfinal released upon completion of ce-ment hydration), R = 8.314 J mol�1 K�1 is the Boltzmann’s constantand f(a) is a normalized function for heat.

Thermal boundary conditions are applied through a prescribedflux per unit area qT formulated as [46]:

qT ¼ hcrðTb � TeÞ ð3Þ

where hcr is a mixed convection–radiation boundary transfer coeffi-cient, Tb is the boundary surface temperature and Te is the environ-mental temperature.

The simulation of thermal inputs associated to solar radiation inconcrete structures can be made with significant accuracy throughthe adoption of models that are readily used in meteorological sci-ences [31,47]. Such models can take into account the effects of thespatial relationship between the earth and sun at a given time ofthe day/year and thus predict the solar radiation that reaches a cer-tain point on earth at sea level (i.e. low atmosphere). It is furtherpossible to compute the angle between the sunbeam and any arbi-trarily oriented/inclined surface, and evaluate the intake of energythroughout the day of such surface.

The calculation of the solar energy that reaches earth at sea le-vel, qm, is based on the solar constant, q0, which represents the to-tal radiation energy received from the sun at a distancecorresponding to 1 Astronomical Unit. Even though q0 variesslightly throughout the year by�7%, it is usually acceptable to con-sider q0 = 1367 W m�2. The estimation of qm can be done throughthe following empirical equation [47,48]:

qm ¼ q0 � e�Tl

0:9þ9:4�sinðhÞ ð4Þ

where Tl is the Linke turbidity factor that summarizes the turbidityof the atmosphere (attenuation of the direct beam solar radiation)and h is the solar elevation that corresponds to the angle betweenthe direction of the sunbeam and the idealized horizon. Tl is knownto usually vary between 3 and 7, whereas h can be calculated bytaking into account latitude, date and time. Further geometricalconsiderations allow the calculation of the angle between anincident sunbeam and the vector orthogonal to an arbitrarily

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oriented/inclined surface, termed as i (see detailed description ofmodels to calculate h and i in [44,47]).

Based on the knowledge of qm and i at a given instant, and con-sidering the absorvity of the material of the target surface (aS), it ispossible to calculate the radiation energy qS that is actuallyabsorbed:

qS ¼ aS � qm � cosðiÞ ð5Þ

Another particularity of the present application in regard to previ-ous works [30] is the use of prestressing ducts acting as coolingpipes. A formulation is thus necessary to describe the added inter-nal heat fluxes that are caused by the presence of an embeddedcooling pipe, which can be expressed in the following energy bal-ance equation, to be applied throughout the length z along the pipe[49]:

_m � cpdTc

dz¼ hc � P � ðTc � TwÞ ð6Þ

where _m is the mass flow rate of the coolant (air in this case), cp isthe specific heat of the coolant, hc is the boundary transfer coeffi-cient between the coolant and the surrounding concrete, P is theperimeter of the cooling pipe, Tc is the temperature in the coolingpipe and Tw is the bulk temperature of concrete around the coolingpipe. The mass flow rate of the coolant _m can be estimated throughthe product of the cooling fluid density (q) by the fluid mean veloc-ity (lm) and by the cross sectional area of the cooling pipe (Ac). Theimplementation of Eq. (6) into a finite element software [43] bringsfurther nonlinearities due to the interaction between the coolingfluid and the surrounding concrete, which results in progressiveheating of the cooling fluid along the pipe.

For updating age-dependent properties along time in themechanical model, the equivalent age of concrete teq is adopted.Its formulation is based on an Arrhenius type equation establishedfor a reference temperature Tref (usually 293.15�K) [50]. For a giveninstant t, the equivalent age can be calculated as:

teq ¼Z t

0e�Ea

R1

TðsÞ�1

Tref

� �ds ð7Þ

2.2. Mechanical model

The mechanical model is relatively similar to those adopted fortime-dependent mechanical analysis of hardened concrete, exceptfor some particularities associated to the facts that: (i) it is beingpreceded by a thermal analysis (de-coupled), with imposition ofstrains that are calculated with basis on the thermal dilation coef-ficient of concrete (aT) and the previously calculated temperaturefield; (ii) there is a strong evolution of mechanical propertiesthroughout the analysis, dully taken into account through theequivalent age concept; (iii) the strong viscoelastic behaviour ofconcrete at early ages makes it necessary to use creep formulationsthat can provide adequate estimates within such time span. In spe-cific regard to the last point (iii), basic creep of concrete was ac-counted for through the use of the Double Power Law (DPL),which has a reasonably good performance on both early age andlong term time spans [51]:

Jðt; t0Þ ¼ 1E0ðt0Þ

þ /1

E0ðt0Þðt0Þ�mðt � t0Þn ð8Þ

where J(t, t0) is the compliance function at time t for a load appliedat instant t0, E0(t0) is the asymptotic elastic modulus, and /1, m and nare material parameters. Since drying creep is negligible for anapplication that only envisages early age behaviour, it was disre-garded [52].

Bearing in mind that the aim of the thermo-mechanical simula-tions is to assess the risk of cracking, the post-cracking behaviour isnot considered relevant and it is thus not simulated. Thus, linearelastic behaviour (with creep) is considered for concrete both incompression and in tension.

3. The Paradela dam spillway: description and monitoring

3.1. Overview

The Paradela dam, located in the North of Portugal, is a rockfillgravity dam built in the 1950s, with 540 m longitudinal develop-ment and maximum height of 112 m from foundation. Due to re-cent hydraulic problems in one of the dam’s spillways, it wasnecessary to build a new complementary ‘ski-jump’ spillway onthe right margin of the river [53]. The case study reported in thispaper concerns the cooling measures and assessment of crackingrisk in the construction of the central wall of the spillway entrance.

3.2. Description of the spillway entrance

3.2.1. Geometry and construction phasingThe spillway functions in free surface conditions, and has two

main entrances at the top level of the dam, each with 5.5 m width,being separated by a hydro-dynamic shaped wall with maximumwidth of 2.8 m and 17.4 m height. A three dimensional representa-tion of the entrance region of the dam spillway is shown in Fig. 1a,whereas its corresponding plan view at approximately mid-heightof the wall is depicted in Fig. 1b. The reinforcement of the middlewall can be generally characterized by £16//200 mm placed verti-cally and £12//200 mm placed horizontally near each surface witha concrete cover of 60 mm.

The construction of the wall was generally performed with1.2 m tall construction phases, with empirically defined waitingperiods being defined by a target temperature in the core regionsduring the cooling period (approximately 27 �C, which corre-sponded to 17 �C above average daily temperature duringconstruction). In order to minimize such waiting periods, anair-cooling system based on ventilated prestressing ducts placedhorizontally was implemented, allowing lower peak temperaturesand faster return to temperature equilibrium with the surroundingenvironment. The main scope of the present paper is the study of aspecific construction phase that corresponds to the zone of embed-ment of the fixation parts of the sluice gates. Such fixation partswere approximately 2.5 m tall, and it was thus desirable toperform a 2.5 m tall construction phase, labelled as 9th phase inFig. 2a. Due to its larger thickness, this construction phase is thecritical one in terms of peak temperatures and cracking risk, beingtherefore the object of analysis.

3.2.2. MaterialsThe wall of the spillway entrance was generally cast using con-

crete of class C30/37 [54] with the composition labelled as S1-D32in Table 1. In the 9th construction phase, due to increased com-plexity of reinforcement near the downstream extremity of thewall (related to the salient concrete blocks), two slightly differentcompositions with higher fluidity were used in the vicinity of suchregion, as shown in Table 1 (S3-D32 and S3-D16). In spite of this,the areas of most interest to this study (thickest regions of the walland monitored sections) correspond to the upstream region. There-fore, and also taking into account the fact that the compositionshave similarities, all the characterizations and modelling in thescope of this work pertain to mix S1-D32. Steel reinforcementwas S400C [54], with characteristic yield stress of 400 MPa.

Fig. 1. Entrance of the dam spillway: (a) three dimensional representation and (b) plan view of the central wall.

Fig. 2. (a) Construction phasing of the central wall (elevation) and (b) overall photo during the construction of the 9th phase.

Table 1Mix proportions of concrete in the 9th phase of the spillway wall.

Components C30/37 S1-D32 (kg/m3) C30/37 S3-D32 (kg/m3) C30/37 S3-D16 (kg/m3) Concrete pour plan

Gravel 14–32 mm 449 400 –Gravel 10–16 mm 438 400 490Gravel 4–8 mm 306 316 387Sand 621 646 763CEM I 42.5R 224 238 280Fly ash 96 102 120Plasticizer 2.2 3.4 4Water 170 180 200

M. Azenha et al. / Engineering Structures 62–63 (2014) 148–163 151

3.2.3. Cooling systemIn view of the work reported by Hedlund and Groth [8,9], which

proposed the possibility of using ventilated prestressing ducts forconcrete cooling, and bearing in mind the easy availability of thecorresponding necessary equipment in the construction site ofthe wall, it was decided to test the feasibility of this kind of coolingtechnique. However, in view of practical limitations posed by con-tractor/owner, this pilot application of air-cooling system wasslightly different from that of Hedlund and Groth [8,9]. Instead ofplacing the tubes vertically along the wall, they were placed hori-zontally, even though this implied more limited cooling capacity asthe length of tube along freshly cast concrete is longer. In the par-ticular case of the 9th construction phase, a total of 6 prestressingducts of 90 mm diameter have been used, with their air intakebeing made horizontally at the downstream extremity of the wall,and the outtake made near the upstream extremity, on the top sur-face of the casting phase, in order to avoid a direct upstream–downstream potential leakage channel after construction. The

overall path of the ducts is shown in the schemes of Fig. 3, withthe ducts labelled from T1 to T6. For ventilation, a 0.60 m diameterfan was used, with 1200 m3/h ventilation capacity, that collectedair from the environment and blowed it into the ducts at an inter-nal air speed of approximately 8.6 m/s (measured with anemome-ter at the outtake of the ducts). The fan had to be placed in thedownstream extremity of the wall due to practical constraints ofthe contractor. The cooling system was only started at the age of14 h after the end of casting operations to avoid introducing poten-tially undesirable vibrations to the freshly cast concrete before itsstructural setting time. The ventilation system was disconnected8.6 days after casting in view of the similarity of temperature be-tween the wall’s core and the surrounding environment. After that,the prestressing ducts were filled with mortar using standard pro-cedures [55].

It is remarked that this cooling system had been previously testedin the 8th phase of casting, with three prestressing ducts placed atmid-height. Details on this test can be found elsewhere [56].

Fig. 3. Cooling system at the 9th construction phase: (a) plan view; (b) longitudinal section and (c) cross-section A0–A.

152 M. Azenha et al. / Engineering Structures 62–63 (2014) 148–163

3.3. Monitoring and material characterization

3.3.1. General remarksIn order to better understand the effectiveness of the cooling

system, its influence on the cracking risk, and assess the capabili-ties of the adopted numerical simulation strategy, an extensiveset of actions has been carried out, comprising in situ internal mon-itoring of temperatures/strains of concrete, in situ validation tests,as well as laboratory material characterization.

3.3.2. Temperature monitoringTemperatures inside the 9th construction phase have been

monitored with K-type thermocouples, aiming particularly atassessing temperature profiles in a region near the maximumwidth of the wall (i.e. at approximately 5.3 m from the upstreamextremity: section A–A0 as identified in Fig. 3a). The placement oftemperature sensors in section A–A0 is depicted in the scheme ofFig. 4a, where thermocouples are identified by the prefix TC. Tem-peratures in the locations labelled as VW in Fig. 4a have also beenmonitored with resistive temperature sensors, as these are thelocations of vibrating wire strain gauges, which contain internaltemperature sensors for strain compensation. The internal air tem-perature of ducts T1–T3 has been monitored both in section A–A0

and in neighbouring areas, as shown in Fig. 4b. Environmental tem-perature (dry-bulb) has been assessed with a thermocouple.

Monitoring was carried out since the instant of casting during aperiod of 10 days, and the measurement frequency was set to 1reading per each 30 min. The internally monitored temperaturesin concrete are shown in Fig. 5 for a vertical and a horizontal align-ment of sensors that pass through sensor VW3. It is remarked thatthe results of this figure and all upcoming figures of the paper

(either concerning experimental results or numerical simulationresults) have their corresponding time axis zeroed in regard tothe instant at which the casting operations of the studied phasewere finished. From Fig. 5 it can be seen that the initial tempera-ture of concrete was �15 �C, and the peak temperature wasapproximately 42 �C in the core regions (VW3 and VW5). Further-more, it can be observed that the ascending branch of temperaturedevelopment is clearly affected at the age of 14 h, when the coolingsystem is activated. In specific regard to the vertical profile of tem-peratures shown in Fig. 5a, the expectable behaviour was captured:the core region has the highest peak temperatures (VW3, VW5),whereas a decrease trend is seen towards the top surface. In fact,sensors TC6 and TC7 exhibit maximum temperatures of �36 �C,while VW6 (near the top surface) has the lowest peak temperature(circa 27 �C). Near the bottom surface of this construction phase,sensor VW1 highlights the importance of the heat storage effectcaused by the previously cast concrete: in fact, even though thetemperature peak is lower than that of the core regions, it occurslater and the heat loss rate observed afterwards is lower than inother regions. It should also be remarked that all sensors are al-most in equilibrium with environmental temperature by the ageof 8 days.

In regard to the temperature development in the sensors lo-cated along a horizontal alignment, shown in Fig. 5b, it can be no-ticed that the sensors located in the vicinity of vertical boundariesexhibit lower temperature variations (VW4 and TC5), with temper-ature peaks under 35 �C. It is interesting to observe that TC5 isplaced nearer the surface (15 cm apart) than the case of VW4(20 cm apart), and consequently the temperature peak of VW4 isslightly higher. By looking at the temperature evolution after theage of 8 days (heat of hydration has been dissipated), it can be seen

Fig. 4. Temperature and strain measurement devices: (a) location of sensors within section A0–A and (b) location of sensors within the tubes T1, T2 and T3. Units: [m].

Fig. 5. Monitored temperatures (two alignments that encompass sensor VW3): (a)vertical alignment and (b) horizontal alignment.

M. Azenha et al. / Engineering Structures 62–63 (2014) 148–163 153

that temperatures in VW4 remain higher than those of TC5 due to acombination of two main reasons: VW4 is located in the vicinity ofa surface oriented to southeast, which is bound to receive more en-ergy through solar radiation than the surface near TC5, which isoriented to northwest; VW4 is 5 cm deeper than TC5, thus beingslightly nearer the inner core (higher thermal inertia).

The temperature evolution along the air inside duct T3 mea-sured by sensors TC3, TC9 and TC10 (Fig. 4b) is shown in Fig. 6,where the environmental temperature and the temperature in

VW5 (hottest region in concrete) are also represented for compar-ative purposes. The information provided by such figure allows theclear identification of the instant at which air ventilation began(14 h), as the rate of temperature rise is clearly disrupted insidethe duct. Furthermore, the rising temperature tendency along theduct’s length is identifiable, as the temperature is consistentlyhigher in TC3 in regard to TC10, and in TC9 in regard to TC3. Takingas example the temperatures recorded at the instant of peak tem-perature (1.88 days), TC9 measured a temperature of 32.2 �C,whereas TC10 indicated a temperature of 29.4 �C. This representsa shift in temperature of approximately 3 �C in 3.5 m length ofduct. At three instants of this study, temperatures were measuredalso at the entrance of T3 (x = 0) and at x = 1 m through the use of ahandheld temperature probe (PT1000). By joining such data withthe results of TC3, TC9 and TC10, it was possible to plot a temper-ature profile along the duct for t = 3.84 d, t = 4.56 d and t = 6.58 d –see Fig. 6b. It can be observed that the temperature at the inlet ofthe tube matches the environmental temperature, and that theheating of the air along the tube is strongly dependent on the com-bination of environmental temperature and internal temperatureof concrete. In fact, in the most unfavourable situation shown byFig. 6b, air was heated from �11 �C at the air inlet to �26 �C at apoint located 25 m away from the air inlet (t = 4.56 d). This in-crease of air temperature is bound to reduce its cooling capacity.However, in spite of such diminishment of cooling capacity, thetemperature of the air in the hotter regions of the duct remainedat least 10 �C below that of concrete during the periods at whichtemperature in concrete was near its peak (see t = 1 d to t = 3 d inFig. 6a), showing that the heat removal potential was not negligibleat all.

The observed diminishment of cooling capacity along the lengthof the duct highlights the fact that the adopted configuration forthe tubes does not maximize cooling capacity, which would beconversely maximized if the length of tube inside concrete hadbeen minimized. Such goal could have been achieved by providinga vertical arrangement for the tubes and introducing more individ-ual smaller tubes.

3.3.3. Strain monitoringStrain measurement was carried out with vibrating wire strain

gauges of metallic casing with 14 cm reference length (TES/5.5/T –Gage Technique). Past laboratory tests and in situ applications[44,57,58] have shown that this kind of sensor is robust and ade-quate for strain measurement in concrete at early ages. The straingauges were placed at the locations identified in Fig. 4a (VW1–VW6), dully positioned in order to measure strains in the longitu-dinal direction of the wall. VW7 has distinct intents and shall be

Fig. 6. Monitored temperatures in duct T3: (a) Global results for TC3, TC9 and TC10 and (b) temperature profile along the duct at three selected instants. Units: [m].

Fig. 7. Methodology adopted for determination of the instant of solidarization ofVW sensor to concrete.

Fig. 8. Strains measured by sensors VW1–VW5.

154 M. Azenha et al. / Engineering Structures 62–63 (2014) 148–163

specifically addressed later. Measurements were taken with thesame datalogger and at the same sampling rate as adopted forthe temperature sensors.

One important issue to tackle is the ‘zeroing’ of the measuredstrains. In fact, before concrete sets and has enough stiffness todrive the sensor into the same deformation state, the measure-ments taken by the sensor do not have any relevant physical mean-ing. It is thus necessary to assess the instant of solidarization (i.e.the full bond) between concrete and the sensor. In a previous work[57], the zeroing operation has been made by assessing the instantat which two sensors with different casing (plastic and metallic),placed under the same conditions inside concrete, started yieldingthe same results. This means that both sensors are solidarized (asthe plastic sensor is bound to solidarize earlier due to its smallerstiffness). Since the plastic cased sensor was not available, an alter-native methodology for zeroing the data was implemented. Byinterpretation of the findings reported in [57] it can be consideredthat the solidarization instant coincides with a progressive changein the derivate of strain variation detected by the sensor, that canbe obtained by geometrical intersection of tangents of measuredstrains, as shown in Fig. 7. It was decided to use such ‘zeroing’ cri-terion in the scope of this research work. As a result of the applica-tion of such rule, the solidarization instant of each sensor VW1–VW5 was, respectively: 0.17, 0.19, 0.19, 0.20 and 0.17 days (dueto malfunctioning of the VW6, the strain results from this sensorare not available). The solidarization instants seem coherent, sincethey have a trend to increase with the distance of the sensor fromthe bottom surface of the casting block, due to the natural delay inits involvement by concrete during the casting process.

The measured strains in sensors VW1–VW5 are shown in Fig. 8.Even though the strain output is dependent on several factors thatinteract with each other (thermal deformation, restraint, creep), itis possible to find a set of common points and reasoning. Overall,all deformations are strongly commanded by the temperature var-iation, following the same kinetics. Sensors VW2, VW3 and VW5,which are located in regions near the core of the wall’s cross sec-tion and had similar temperature development histories, also havesimilar strain developments. This is bound to be caused by similarthermal deformations and restraints for these locations of mea-surement. The smallest deformations are recorded in VW1, whichis located in the bottom of the casting phase (i.e. near the existingconcrete) and thus having less temperature rise (thus less expan-sion), while being more restrained by the existing concrete belowat lower temperatures. Finally VW4, which is near the surface and

thus has lower temperature rises (maximum temperature of�33 �C), also has a smaller deformation variation when comparedto VW2, VW3 and VW5.

Fig. 10. Heat generation rate of a cement paste with CEM I 42.5R and w/c = 0.5.

M. Azenha et al. / Engineering Structures 62–63 (2014) 148–163 155

In order to assess free deformations of the concrete used in theconstruction (associated to unrestrained autogenous shrinkage andthermal deformations), strain was measured in a concrete cylinderplaced in specially devised conditions. The concrete cylinder(150 mm diameter and 300 mm tall), cast simultaneously withthe studied construction phase and by using the same concrete,was cast into a special mould internally coated with a soft mem-brane (with lids also coated with such material). A strain gaugewas placed inside the mould to measure longitudinal strains –see photo of the open mould in Fig. 9a. After casting concrete intosuch mould, it was placed horizontally inside the studied construc-tion phase (during its casting procedures) at the location that isidentified as VW7 in Fig. 4a. This kind of procedure, here termedas use of a ‘no-stress specimen’, has been reported in Choi et al.[59], and it allows to measure free deformations of concrete, whichcan in turn be used to assess the thermal dilation coefficient, TDC(provided that the temperature inside the concrete cylinder is rel-atively uniform, and autogeneous shrinkage deformations areknown). Unfortunately, due to undetermined causes, the outputof the sensor could not be read during the first 0.8 days, and thusthe reported data only starts at such age, as shown in Fig. 9b.

3.3.4. Heat generation and activation energyIn an extensive experimental program for the characterization

of the cements marketed in Portugal, Azenha [44] has reported alibrary of heat generation obtained through isothermal calorimetryunder several temperatures for plain cement pastes with w/c = 0.5.The cement used in the construction concerned in this paper wasalso characterized (same supplier and manufacturing plant), andthe resulting information for calorimetry tests under 20 �C, 30 �C,40 �C and 50 �C is shown in Fig. 10. A reasonable estimate of theheat generated by concrete can be obtained by multiplying theheat generation reported in Fig. 10 by the volumetric content of ce-ment, which is of 224 kg/m3. By using the speed method algorithm[44,60], the necessary data for the numerical simulation of heatgeneration according to Eq. (2) was obtained: Ea = 37.31 kJ/mol,A = 4.989 � 109 W/m3, Qpot = 8.295 � 107 J/m3, and function f(a)characterized by the following set of data [a; f(a)] = [0.00; 0.00],[0.05; 0.58], [0.10; 0.85], [0.15; 0.98], [0.20; 1.00], [0.30; 0.94],[0.40; 0.69], [0.50; 0.41], [0.60; 0.22], [0.70; 0.13], [0.80; 0.07],[0.90; 0.02], [1.00; 0.00]. Even though this data pertains to CEM I42.5R of the same company that supplied the cement to this con-struction site, there may be deviations caused by inevitable varia-tions in the characteristics of the cement. Also, the extrapolationprocedure mentioned above did not take into account the presenceof fly ash in the mix (96 kg/m3), which may have non-negligible

Fig. 9. In situ determination of TDC: (a) overview of installed ‘no-stress’ spe

effects on the heat generation potential and hydration kinetics.Therefore, in order to assess the potential importance of such devi-ations, a semi-adiabatic test was conducted in situ (simultaneouslywith the casting operations) in a 30 cm edge concrete cube, dulyisolated by 2.1 cm thick plywood and 12 cm of polystyrene boards.The results of such semi-adiabatic calorimetry test shall be ad-dressed in Section 4, upon the simulation of its temperature devel-opment through the finite element method.

3.3.5. Complementary laboratory characterizationCompressive strength evolution was assessed with concrete

cubes (150 mm edge) at the ages of 1, 3, 7 and 29 days, whereastensile strength was measured with splitting tests on cylinders(150 mm diameter and 300 mm tall) and at the same ages. It is re-marked that testing at the reference age of 28 days was not possi-ble due to laboratorial constraints. The evolution of both tensileand compressive strength for concrete cured at 20 �C (saturatedconditions) is shown in Fig. 11a (average results of three specimensat each age). In order to assess the activation energy suitable forcompressive strength maturity estimations, a set of concrete cubeswas cured at 40 �C with the compressive strength measured at thesame ages. The corresponding results are also shown in Fig. 11a. Byapplying the equivalent age concept [61], together with the ‘super-position method’ [60], it was possible to asses that the activationenergy based on mechanical testing has the value of 37 kJ/mol,which is rather consistent with the activation energy obtainedthrough isothermal calorimetry for the same cement (yet without

cimen and (b) evolution of strain and temperature evolutions in VW7.

Fig. 11. (a) Evolution of compressive and tensile strength of concrete. (b) Basic creep of concrete: specific creep for loading at the ages of 1.1, 3.3 and 7.3 days.

156 M. Azenha et al. / Engineering Structures 62–63 (2014) 148–163

fly ash), 37.31 kJ/mol [44]. Such coincidence in activation energyfor thermal and mechanical phenomena had already been reportedby Ulm and Coussy [62].

Basic creep was assessed in creep rigs on prismatic specimens(sealed) with dimensions 15 cm � 15 cm � 60 cm, loaded at 30–40% of the concrete compressive strength and internally monitoredwith vibrating wire strain gages. Such creep tests were conductedat the ages of 1.1, 3.3 and 7.3 days, and the corresponding specificcreep curves are shown in Fig. 11b.

The experimental program also included a single specimen formeasurement of autogenous shrinkage. Such specimen was a150 mm diameter and 300 mm long cylinder, internally instru-mented with a vibrating wire strain gage, which was kept in itsformwork during the experiment and sealed with a plastic filmon the top surface. Unfortunately, two factors contributed to ren-der the results of this specimen unusable for this research: onone hand, the monitoring only could be started at the age of 2 daysin the laboratory due to unavailability of datalogging system; onthe other hand, the measurements of autogenous taken sincet = 2 days were disturbed by an inefficient sealing, which promotedundesired drying of the specimen. This was not considered a criti-cal problem in view of the low values of autogenous shrinkage thatare usually expectable in concretes of low cement content and highw/c ratio [63–65].

Fig. 12. E-modulus of concrete assessed by compressive cyclic testing and EMM-ARM.

3.3.6. Continuous monitoring of concrete stiffnessThe evolution of elasticity modulus along time was measured

through compressive cyclic testing in concrete cylinders(150 mm diameter and 300 m tall) at the ages of 1, 3, 7, 15 and29 days. Concomitantly, E-modulus of concrete was continuouslyassessed through a methodology termed as EMM-ARM (ElasticityModulus Measurement through Ambient Response Method). Thismethodology has been developed by Azenha et al. [42] and consistsin a variant to classic resonant frequencies that allows the quanti-fication of E-modulus continuously since the instant of casting ofthe specimen inside the testing mould. The basic principle ofEMM-ARM is the following: the specimen is cast inside the testingmould, which is in turn placed in simply supported conditions andcontinuously subject to modal identification (using accelerome-ters) without any explicit excitation of the beam, as ambient vibra-tion suffices. As concrete hardens, the first resonant frequency ofthe composite beam evolves, and the stiffness of concrete can beinferred by applying the equations of motion. Details about the

testing setup and procedure applied for the concrete of this spill-way application can be found in [56,66,67], as it corresponds toan improved version of the originally devised test (a steel mouldis used). The collected results with EMM-ARM and cyclic compres-sion tests on cylinders are shown in Fig. 12, where the feasibility ofEMM-ARM is confirmed in view of the resemblance of results. Also,the richness of information that can be obtained through EMM-ARM represents an added value for the numerical simulation.

4. Numerical modelling

4.1. Geometry, mesh, materials, initial/boundary conditions and timeintegration

4.1.1. Geometry of the model and finite element meshA cross-sectional scheme of the model for simulation is shown

in Fig. 13a, where the construction stages considered in the analy-sis can be observed. The first stage of the model encompasses allconcrete until the 7th phase of concreting (inclusive), consideredas hardened concrete, together with the 8th phase of concrete eval-uated as freshly cast concrete. The second and third stages corre-spond to the 9th and 10th phases of concreting respectively. Thisstrategy diminishes the computational cost of the model without

Fig. 13. (a) Scheme of the simulation model and phasing; (b) section A0–A of the model and (c) finite element mesh.

M. Azenha et al. / Engineering Structures 62–63 (2014) 148–163 157

relevant effect on the accuracy of results. For similar reasons, theunderlying subgrade is not modelled, as it is far from the construc-tion phases of interest. Due to the geometrical symmetry of thewall, a longitudinal plane of symmetry is considered, identifiableby a ZX plane in Fig. 13.

The simulation was made with a 3D finite element model com-prising rectangular brick FE of 8 nodes (2 � 2 � 2 integrationscheme) for concrete in the thermal model, and coincident 20nodes brick FE (3 � 3 � 3 integration scheme) in the mechanicalanalysis. Convective boundaries were modelled with 4 node planarelements (2 � 2 integration scheme), and the cooling ducts wereconsidered with linear elements of 2 nodes (2 point integrationscheme) [43]. The schematic representation of geometry, castingphases and cooling duct location for section A0–A is shown inFig. 13b. The reader is reminded that phase 8 also had coolingducts, according to the description of Section 3.2.3. The longitudi-nal layout of all the ducts was considered straight on an horizontalplane, in correspondence to simplifications in the vicinity of theextremities of the wall.

It should be remarked that due to the phased analysis, the meshevolved along time with some elements/boundaries being acti-vated/de-activated (e.g. the convective top boundary of a givenphase is de-activated upon the beginning of the next castingphase). The mesh adopted for this simulation is shown inFig. 13c, with a total of 18,738 elements and 66,037 nodes. As con-sequence of the symmetry simplification, it was considered thatthe perimeter of the tube elements located in the symmetry plane

of the model was equal to half the actual perimeter of the tubes. Assolar radiation does not represent a symmetrical energy input tothe structure, the symmetry simplification adopted is not truly va-lid. However, as solar radiation has most of its effect near the sur-face, it was decided to keep the symmetry simplification byconsidering the southeast half of the wall, which is most subjectto solar radiation effects.

4.1.2. Materials (thermal and mechanical properties)The thermal conductivity and specific heat of concrete were

estimated with basis on the pondered average of the correspondingthermal properties of the constituent materials of the mix [44,68].The adopted values for k and qc for concrete were, respectively2.40 W/m K and 2.4 � 106 J/m3 K. Even though it is known thatthese thermal properties suffer variations during early ages [69–75], the adopted modelling approach considers them constant inview of the conclusions of the parametric analyses reported byAzenha [44], where a relatively small impact of considering evolv-ing k and qc was found on computed temperatures in hardeningconcrete.

In regard to the data for heat of hydration, the parameters men-tioned in Section 3.3.4 were used. The adequacy of these parame-ters was evaluated through the semi-adiabatic calorimeterdescribed in the same section, whose behaviour was simulatedthrough a FE simulation model that explicitly considered the ex-truded polystyrene (XPS) and wood walls of the calorimeter. Thematerial modelling parameters for concrete coincide with those

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herein described, whereas additional information and mesh repre-sentation are shown in Fig. 14a. The results of the simulation of thecalorimeter were quite coherent with those collected experimen-tally, as seen in Fig. 14b, leading to the confirmation that the strat-egy described in Section 3.3.4 to determine the heat generationand activation energy was adequate for defining the heat of hydra-tion modelling parameters adopted in the simulations.

The thermal dilation coefficient (TDC) of concrete was assessedwith basis on the ‘no-stress’ specimen described in Section 3.3.3.However, in order to obtain the thermal deformation and calculatethe TDC, it was necessary to subtract the autogenous shrinkagedeformation from the total deformation. As data on autogenousshrinkage was not available, an estimate of the autogenous shrink-age evolution based on Eurocode 2 [54] was used. Another issue totake into account is the fact that the TDC of concrete is not constantduring the first hours of age. In fact, several authors have dealtwith this subject, and it generally agreed that the initial TDC tendsto be larger than that of hardened concrete, and tends to decreasesharply within the first 12 to 24 h of age, reaching then the plateaulevel corresponding to hardened concrete [69,73,76]. The ‘no-stress’ specimen cast in the scope of this research cannot be usedto estimate TDC at concrete early ages, due to the absence of datain the first 8 h reported in Fig. 9b. Nonetheless, as the peak temper-ature occurred later than 24 h age, and full data is available fortemperatures and strains occurred in such period, calculationscould be made under the assumption that TDC was already at itsplateau value. The TDC was estimated between instants t = 2.0 d(peak temperature) and t = 8.0 d (local minimum) as shown inFig. 9b, and the autogenous shrinkage strain variation in such per-iod was estimated to be of 8.45 le (considering fcm = 42.3 MPa inEurocode 2 [54]). The estimated constant TDC to be used in thenumerical simulation was 11.07 le/�C. Nonetheless, since it isknown that TDC varies during the first 24 h of age, an alternativeformulation for the evaluation of the TDC was considered, basedon experimental evidence reported by Laplante and Boulay [77].Therefore, this alternative formulation considers that during thefirst 16 h the TDC varies according to TDC(t) = 0.16t2 � 4.88t+ 48.93 (t in hours), and remains constant after such age.

The E-modulus evolution of concrete for the simulation modelwas directly extracted from EMM-ARM data reported in Fig. 12,whereas creep modelling was made through the adjustment ofDPL parameters to the creep data experiments. The best-fit creepparameters and their adjustment to the experimental data areshown in Fig. 11b. Prestressing ducts were modelled with

Fig. 14. (a) FE model for the semi-adiabatic calorimeter (geometry and general data/unitof the calorimeter.

consideration of their inner perimeter of 283 mm (90 mm innerdiameter). Steel reinforcement was disregarded in temperaturecalculations due to its low interference in temperature develop-ment [44,52]. Regarding mechanical field simulations, steel wasnot considered because post-cracking behaviour was not sought.In the non-cracked stage, the similitude in TDC of steel in regardto that of hardened concrete strongly minimizes the restraint toconcrete thermal deformation, thus rendering the effect of rein-forcement negligible for the computation of thermal stresses atearly ages [52].

4.1.3. Initial/boundary conditions and construction phasingIn what concerns the boundary conditions in the thermal prob-

lem, it was assumed that an average wind speed of 2.5 m/s oc-curred (confirmed during 3 days with an anemometer) and theresulting convection/radiation coefficient for concrete surfaces incontact with the environment was estimated to be 15.25 W/m2 Kin accordance to the predictive formula of Branco et al. [78]. Forthe particular case of formworks surrounding concrete, an equiva-lent boundary convection coefficient was adopted according to anelectrical analogy [46]. Bearing in mind that the formworks weremade of wood (kwood = 0.175 W/m2 K) and their thickness was18 mm, the resulting equivalent boundary coefficient was 6 W/m2 K. Formwork was applied to the vertical surfaces of each castingstage during the first 7 days of age, and removed afterwards. Allconvective boundaries were subjected throughout the analysis tothe environmental temperature that was monitored in situ – seeFig. 5. The symmetry plane of the model was considered as an adi-abatic boundary.

In what regards to solar radiation intake, the several surfacedirections of the model were taken into account, and the absorbedradiation was calculated according to the model described in Sec-tion 2.1. Absorvity of concrete was considered as 0.6 [79], the lati-tude was 41.77�N, and the casting date was 28/03/2011. The Linketurbidity factor was considered as 2.5, and its feasibility was con-firmed by comparing computed solar radiation on horizontal sur-faces with solar radiation data from a nearby weather station (RioTorto Station) in conditions of clear skies. The effect of cloudinesswas taken into account by normalizing the predictions of theadopted solar radiation model according to information obtainedfrom the piranometer of the neighbouring weather station. Thediminishment of solar absorption caused by shadows cast by neigh-bouring objects was considered negligible throughout the entireday, as the wall was one of the tallest elements in the landscape.

s in millimetres). (b) Calculated and recorded temperature in the geometrical centre

M. Azenha et al. / Engineering Structures 62–63 (2014) 148–163 159

Bearing in mind the construction phases taken into account inthis calculation model (identified in Fig. 13a), the initial tempera-tures were considered as follows. For the existing concrete at thebeginning of analysis, it was assumed that concrete was alreadyin thermal equilibrium with the environment, and so, the averagedaily temperature of the preceding week (14.5 �C) was consideredfor the existing concrete. For the subsequent stages of construction,the initial temperature of concrete was obtained from in situ mon-itoring (average value), which was also of approximately 14.5 �C.

In regard to the cooling ducts, due to a limitation of the adoptedsoftware, the inlet temperature had to be considered constant,

Fig. 15. (a) Measured and simulated temperatures for VW1, VW2, VW4 and VW5. (b) Temmaximum temperatures are attained. (c) Calculated temperatures in the points where t

equal to the average environmental temperature during the timein which the tubes were active. The following temperatures wereconsidered for each phase: 8th: 7.13 �C and 9th: 13.91 �C. Theinternal convection coefficient in the ducts was obtained with ba-sis on their internal air speed of 8.6 m/s, which according to thestudies of Hedlund and Groth [8] should correspond to a convec-tion coefficient of 30.0 W/m2 K. The cooling duct elements wereactivated at each construction phase when the surrounding con-crete had age of 14 h.

Taking into account the directions of the axes of the coordinatesystem presented in Fig. 13a, the mechanical boundary conditions

perature distribution within the thicker section of the wall for the instants that thehe highest temperatures are attained in the thicker section of the wall.

160 M. Azenha et al. / Engineering Structures 62–63 (2014) 148–163

consisted in placing Z direction supports on the bottom surface ofthe wall and Y supports in all the elements of the symmetry plane.

The time step strategy adopted for the model consisted in con-sidering the initial time coincident with the casting instant of 8thphase. Casting of the 9th and 10th phases were considered at therelative instants t = 24 days and t = 41 days in accordance to theactual construction. All analyses were conducted with a constanttime step of 1 h duration, even though some localized adjust-ments were necessary in view of construction phasing and venti-lation activation. Nonetheless, all adjustments were carefullymade to assure that the duration of all time steps remained under1 h.

4.2. Results and discussion

The presentation and discussion of results is centred in the 9thconstruction phase, with particular emphasis for comparisons be-tween monitored and simulated temperatures/strains. The tem-perature simulations in concrete were quite coherent with themonitored ones, as it can be confirmed for a set of representativelocations (VW1, VW2, VW4 and VW5), whose results are shownin Fig. 15. In fact the largest deviations regarding the monitoredtemperatures during the entire calculation always remain under4 �C, thus providing confirmation of the feasibility of the modellingstrategy, particularly in regard to the cooling capacity of the venti-lated ducts. Based on the confidence gained on the thermal simu-lation model, a further numerical simulation was made, in whichthe effect of the cooling ducts was disregarded. The correspondingresults for the hottest region of both models are shown in Fig. 15b.It can be confirmed that the inclusion of the duct had a twofold ef-fect: not only was the peak temperature diminished by 5 �C withbenefits for cracking safety, but also the return of the internal tem-

Fig. 16. Measured and simulated strai

perature to thermal equilibrium with the environment was accel-erated, with advantages for construction phasing.

The fact that the calculated temperatures matched well themonitored ones is a solid starting point for the analysis of resultsof the mechanical simulation, as any detected deviations arebound to be solely attributed to issues in the mechanical simula-tion itself. The calculated and measured strains for the same setof sensors that has just been discussed for temperature develop-ment are shown in Fig. 16. The experimentally measured strainsin this figure are represented by their value according to thezeroing procedure mentioned in Section 3.3.3 (Fig. 7), but alsowith a lower and upper bound related to possible uncertaintiesin the instant for zeroing of the sensors output of ±2 h. It canbe seen that all the computed strains with consideration of con-stant TDC underestimate the peak strain at 1.96 day, but thepost-peak kinetics seems to have been well captured. As the con-stant TDC assumption may lead to underestimations of earlystrain development [57], a further calculation was made usinga plausible TDC evolution during the first 24 h, as discussed in4.1.2. The corresponding simulation results are shown inFig. 16, where a better overall fit is seen between experimentaland calculation data (particularly for core regions). Even thoughthe variable TDC was not based on experimental evidence ob-tained in the scope of this research, it is feasible to assume thata significant part of the strain deviations regarding experimentalresults can be explained by the variable TDC at early ages. It hasto be kept into consideration that another possible source ofdeviation of results may be related to the instant at which mea-sured strains were zeroed, which can be debatable. Nonetheless,the adequate prediction of strains that was attained is a goodindication of the feasibility of the computed stresses which areto be analyzed.

ns for VW1, VW2, VW4 and VW5.

Fig. 17. Principal tensile stresses at the point corresponding to the maximumtensile stress in the wall (x = 8.5 m, y = 0.36 m, z = 11.7 m) for three simulationmodels considered.

M. Azenha et al. / Engineering Structures 62–63 (2014) 148–163 161

The discussion of cracking risk is now addressed by comparingthe computed principal tensile stress at the most unfavourable partof the model (located in the core region: x = 8.5 m, y = 0.36 m,z = 11.7 m), as shown in Fig. 17. This figure contains the equivalentage-adjusted measured evolution of the tensile strength, and thecalculation results for the cases of constant and variable TDC, as wellas the case of inexistent cooling ducts and constant TDC. The firstcomment that can be made from Fig. 17, is that the considerationof constant or variable TDC had very marginal effect on the results.The reason for the very small difference is bound to be related withthe very low stress level that is induced during the first 24 h, as theE-modulus of concrete is still very small and creep/relaxation is veryhigh. Regardless of the comparison between these two models, itcan be observed from Fig. 17 that the ratio between the tensile stressand the tensile strength of concrete at the most unfavourable in-stant is of approximately 0.9, which corresponds to a significantcracking risk. Nonetheless, even though the cracking risk was higherthan the desirable one, the structure did not present thermal cracksneither later through-cracks (evaluations made until 2 years aftercasting). It should be remarked that this point of stress analysiswas the most unfavourable one within the structure, and signifi-cantly lower cracking risks were calculated for distinct regions,resulting in a global scenario of much more cracking safety (com-pared to a single-point analysis).

Interestingly, a simulation of the same construction situationwithout consideration of cooling ducts would yield to highercracking risk at the same location (as seen in Fig. 17), with the ratiobetween the tensile stress and the tensile strength of concretereaching 1.2. Therefore, if the calculations made here are consid-ered trustworthy, the use of the cooling ducts may have been thedifferentiating factor that avoided a cracking scenario in thisconcrete lift.

5. Conclusions

A case study regarding the assessment of the cracking risk of athick wall in the entrance of a dam spillway, internally cooled withair-filled prestressing ducts, was presented in this paper. The use ofventilated prestressing ducts is considered more straightforwardthan water ducts due to the often easy availability of ducts andfans in construction works associated to massive concretestructures.

The case study has involved a comprehensive experimentalpart, including laboratory characterization of materials and in situ

monitoring for temperatures and strains. A 3D thermo-mechanicalsimulation of the construction phasing was shown, with input dataduly based in the laboratory characterization program.

In regard to previous works reported in the literature, the workpresented in this paper has its main original contributions in thefollowing fields: (i) the EMM-ARM methodology for continuousmonitoring of concrete E-modulus since casting was applied forthe first time as a characterization tool for stress simulation in con-crete at early ages, thus enhancing the quality of input data; (ii)this is the first reported application of horizontally placed air-cool-ing pipes, with its efficiency being assessed and numerically simu-lated; (iii) the work reported here is relatively unprecedented inview of its holistic approach, with the authors being involved inall tasks of laboratory characterization (allowing sustainable esti-mates of material properties and modelling strategy for numericalsimulations), field monitoring and numerical simulation with ther-mo-mechanical analysis.

It is nonetheless acknowledged that a significant research effortis still necessary in regard to the creep monitoring and modellingat early ages, both in view of the effects of early hydration and inview of temperature effects on creep. Even though in-depth analy-ses of these particular issues of creep have not been included inthis paper, the authors consider them to be of critical importancefor adequate stress simulation in hardening concrete, thusdemanding further research works.

The numerical simulation results were compared to those col-lected by the in situ monitoring and good coherences were observedboth in terms of temperature and strain, providing good prospects inregard to the simulation capabilities of the models and the sound-ness of the experimentally obtained data. The monitoring/simula-tion results allowed concluding that the effectiveness of the aircooling system with horizontally placed pipes is limited in view ofthe significant heating that air suffers along the first meters of tube,thus diminishing its capacity of cooling further regions of concrete.Also, the duct efficiency ends up being quite dependent of the envi-ronmental temperature, which cannot be easily anticipated duringplanning. Nonetheless, in spite of the acknowledged limitations ofair cooling, it may prove quite feasible in relatively cool climatesand small lengths of embedment (e.g.: 10 m or less).

Furthermore, the risk of cracking on the studied constructionphase has resulted acceptable in most of its regions, even thougha non-negligible risk of internal cracking was observed in some re-gions. The fact that no surface or through cracking was observablein the construction corroborates the cracking risk evaluation. Itwas also concluded that the same construction phasing withoutthe use of the cooling ducts had a significantly higher crackingprobability, thus confirming the usefulness of the cooling system.

Finally, it is also worth remarking that a parametric analysisregarding the possibility of considering variable thermal dilationcoefficient of concrete along hydration has been carried out. Theoutcome of such parametric analysis seems to point out a relativelylow impact of admitting the thermal dilation coefficient as con-stant along hydration on the corresponding computed stresses,thus validating such simplification in this case study.

Acknowledgements

Funding provided by the Portuguese Foundation for Science andTechnology to the Research Unit ISISE, to the second authorthrough the PhD Grant SFRH/BD/64415/2009, and to the researchprojects PTDC/ECM/099250/2008 and QREN Number 5387, LEGO-USE, is gratefully acknowledged. The kind assistance of the con-tractor (Teixeira Duarte S.A.) and the owner (EDP – Eletricidadede Portugal) are also deeply appreciated. The contribution of Ânge-lo Costa to the experimental work here reported is also gratefullyacknowledged.

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