revised engstruct d 13 00590r1

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Elsevier Editorial System(tm) for Engineering Structures

Manuscript Draft

Manuscript Number: ENGSTRUCT-D-13-00590R1

Title: Application of air cooled pipes for reduction of early age cracking risk in a massive RC wall

Article Type: Research Paper

Keywords: Cement hydration; service life conditions; thermal shrinkage; cracking; numerical

simulation

Corresponding Author: Dr. Miguel Azenha, PhD

Corresponding Author's Institution: University of Minho

First Author: Miguel Azenha, PhD

Order of Authors: Miguel Azenha, PhD; Rodrigo Lameiras, MSc; Christoph de Sousa, MSc; Joaquim

Barros, PhD

Abstract: The construction of massive concrete structures is often conditioned by the necessity of

phasing casting operations in order to avoid excessive heat accumulation due to cement hydration. To

accelerate construction and allow larger casting stages (usually increasing lift height), it is usual to

adopt internal cooling strategies based on embedding water pipes into concrete, through which water

is circulated to minimize temperature development. The present paper reports the use of horizontally

placed ventilated prestressing ducts embedded in a massive concrete wall for the same purpose, in line

with a preliminary Swedish proposal made in the 1990's. The application herein reported is a holistic

approach to the problem under study, encompassing extensive laboratory characterization of the

materials (including a technique developed for continuous monitoring of concrete E-modulus since

casting), in-situ monitoring of temperatures and strains, and 3D thermo-mechanical simulation usingthe finite element method. Based on the monitored/simulated results, it is concluded that the air-

cooling system is feasible and can effectively reduce early cracking risk of concrete, provided adequate

planning measures are taken.


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Title of Paper:

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

M. Azenha, R. Lameiras, C. de Sousa, J. Barros

Ms. Ref. No.:ENGSTRUCT-D-13-00590

Dear Prof. Herbert Mang:

We have uploaded a revised version of the manuscript, where several modifications were introduced to

attend the reviewers comments and criticisms.

The file Letter_To_Reviewers is dedicated to the reviewers, explaining how their comments were

addressed in the revised manuscript. The revised version of the manuscript is supplied in the file

Revised_Manuscript (please note that the main changes associated to the revision are highlighted in

yellow colour to facilitate the location of changes in regard to the original manuscript).

I look forward to hearing from you soon.

Yours sincerely,

Miguel Azenha (the corresponding author)

(Research Assistant)

Miguel ngelo Dias Azenha, PhDSchool of Engineering University of MinhoCivil Engineering DepartmentCampus de Azurm, 4800-058 Guimares, PortugalTel: +351 253510248; E mail address:[email protected]

University of MinhoSchool of Engineering

Prof. Herbert A. Mang

Editor of

Engineering Structures

Date: November 29, 2013

tter To Editor

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To Reviewer #3

Concerning the remarks raised in the next comments (italic text), the following changes wereintroduced in the manuscript, or alternatively the following clarification comments are made:

Page 10, lines 24 - 26: "The cooling system was ... to avoid inducing undesirable vibrations ...concrete." This remark from the authors is not accurate as the concrete have been pokervibrated during pouring.

Thank you for your comment. Our remark was not accurate indeed. We have corrected thesentence in the revised version of the manuscript:

The cooling system was only started at the age of 14h after the end of castingoperations to avoid introducing potentially undesirable vibrations to the freshly castconcrete before its structural setting time.

Page 17, line 9: Comment why the 29-day strength was used instead of 28-day strength according

to Standard testing (EN 206-1) is missing.There were limitations due to problems that occurred in the laboratory, which did not allowus to test the specimens at 28 days age. However, we tested in the 29 thday of age and,which is likely to provide us very similar results to those that would have been obtained at28 days. Nevertheless, the following clarification has been added to the manuscript:

It is remarked that testing at the reference age of 28 days was not possible due tolaboratorial constraints.

Page 18, line 8: Add reference "(Hedlund, H., (2001), Hardening Concrete. Measurements andevaluation of non-elastic deformation and associated restraint stresses. Doctoral Thesis2000:25, ISBN 91-89580-00-1.)" concerning autogenous shrinkage and evaluation ofdeformations. Results used in thermal stress calculations.

The reference has been added.

Page 19, line 13: Calculation with suggested geometrically symmetry will not be a good estimation.Authors should explain the consequences of the engineering approach.

We believe that the reviewer may have misunderstood the drawings, owing to a mistake onour behalf in labelling the axes Z and Y in Figure 13b. We have corrected such mistake, andthe revised version of Figure 13 is shown below. Except for some details quite near the baseslab, the wall is geometrically symmetric along its middle plane. This holds true even withconsideration of the cooling ducts, which are also symmetric in regard to this plane. Interms of solar radiation intake, the wall is not symmetrical indeed. However, such lack ofsymmetry and the corresponding simplified approach was already dully justified at the end

of section 4.1.1 of the original manuscript.

As we believe that the reviewers comment may have been induced by a misunderstandingof our text and Figure 13, we have corrected the figure as mentioned above, and tried tomake our text clearer in the revised manuscript:

Due to the geometrical symmetry of the wall, a longitudinal plane of symmetry isconsidered, identifiable by a ZX plane in Figure 13.


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a) b)

c)

Figure 13 a) Scheme of the simulation model and phasing; b) Section A-A of the model; c) Finiteelement mesh

Page 20, line 6: "... was equal to half actual diameter". This is an engineering approach that shouldbe explained better as the correct solution using the full diameter, but correcting the heattransfer coefficient of the tube by an reduction of 0,8.

The consideration of half diameter for the tubes located in the symmetry plane was not anengineering approach. Please look at the figure below, were we identify the situation at across-sectional level. In fact, the real situation has half a tube in each side of the symmetryplane, which means that the modelling of the tube, which is made through a bar elementcoincident with the symmetry axis should only promote a heat transfer corresponding to

half a perimeter of the tube. We have considered half diameter, which actually correspondsto a tube with half the perimeter. However, our text had some inaccuracies that may havebeen misleading: (i) we mentioned symmetry axis instead of symmetry plane; (ii) wementioned half a diameter instead of half perimeter for the tube in the symmetry plane.Therefore, in the revised manuscript, the corresponding text has been rephrased to:

As consequence of the symmetry simplification, it was considered that theperimeter of the tube elements located in the symmetry plane of the model wasequal to half the actual perimeter of the tubes.


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All figures indication the time = 0 must be better explained what zero represents.

All graphs in the manuscript have the time zeroed to coincide with the end of castingoperations of the studied phase of the wall. This clarification has been introduced in therevised manuscript upon the presentation of the first graph into which it is applicable (i.e.Figure 5):

It is remarked that the results of this figure and all upcoming figures of the paper(either concerning experimental results or numerical simulation results) have theircorresponding time axis zeroed in regard to the instant at which the casting

operations of the studied phase were finished.

Figure 17: The authors should comment on the lower tensile strength tested on cylinders comparedwith the evaluated tensile stresses, which is higher than the assumed tensile strength.

The results of Figure 17 in which the tensile strength is lower than the evaluated tensilestresses correspond to a hypothetical scenario of not having used the cooling ducts. Thecorresponding comment is made in the last paragraph of Section 4 and is replicated below:

Interestingly, a simulation of the same construction situation without considerationof cooling ducts would yield to higher cracking risk at the same location (as seen inFigure 17), with the ratio between the tensile stress and the tensile strength ofconcrete reaching 1.2. Therefore, if the calculations made here are consideredtrustworthy, the use of the cooling ducts may have been the differentiating factorthat avoided a cracking scenario in this concrete lift.

Yours sincerely,

Miguel Azenha (the corresponding author)(University of Minho, PhD)


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Highlights

Internal cooling of a massive concrete wall with ventilated prestressing ducts; Extensive laboratory characterization of the materials; In-situ monitoring of temperatures and strains; 3D thermo-mechanical simulation using the finite element method Good coherence between field monitoring and numerical simulation.

ghlights

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Abstract

The construction of massive concrete structures is often conditioned by the necessity of phasing

casting operations in order to avoid excessive heat accumulation due to cement hydration. To

accelerate construction and allow larger casting stages (usually increasing lift height), it is usual

to adopt internal cooling strategies based on embedding water pipes into concrete, through which

water is circulated to minimize temperature development. The present paper reports the use of

horizontally placed ventilated prestressing ducts embedded in a massive concrete wall for the

same purpose, in line with a preliminary Swedish proposal made in the 1990s. The application

herein reported is a holistic approach to the problem under study, encompassing extensive

laboratory characterization of the materials (including a technique developed for continuous

monitoring of concrete E-modulus since casting), in-situmonitoring of temperatures and strains,

and 3D thermo-mechanical simulation using the finite element method. Based on the

monitored/simulated results, it is concluded that the air-cooling system is feasible and can

effectively reduce early cracking risk of concrete, provided adequate planning measures are

taken.

bstract

ck here to download Abstract: Abstract.docx
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Application of air cooled pipes for reduction of early age cracking risk in1

a massive RC wall2

By Miguel Azenha1, Rodrigo Lameiras

2, Christoph de Sousa

3and Joaquim Barros

43

Abstract: The construction of massive concrete structures is often conditioned by the necessity4

of phasing casting operations in order to avoid excessive heat accumulation due to cement5

hydration. To accelerate construction and allow larger casting stages (usually increasing lift6

height), it is usual to adopt internal cooling strategies based on embedding water pipes into7

concrete, through which water is circulated to minimize temperature development. The present8

paper reports the use of horizontally placed ventilated prestressing ducts embedded in a massive9

concrete wall for the same purpose, in line with a preliminary Swedish proposal made in the10

1990s. The application herein reported is a holistic approach to the problem under study,11

encompassing extensive laboratory characterization of the materials (including a technique12

developed for continuous monitoring of concrete E-modulus since casting), in-situmonitoring of13

temperatures and strains, and 3D thermo-mechanical simulation using the finite element method.14

Based on the monitored/simulated results, it is concluded that the air-cooling system is feasible15

and can effectively reduce early cracking risk of concrete, provided adequate planning measures16

are taken.17

Keywords: Cement hydration, service life conditions, thermal shrinkage, cracking, numerical18simulation19

1 INTRODUCTION20The combined effect of the exothermic nature of cement hydration reactions and the21

relatively low thermal diffusivity of concrete leads concrete structures to endure temperature22

1Assistant Professor, ISISE Institute for Sustainability and Innovation in Structural Engineering, University ofMinho, School of Engineering, Civil Engineering Dept., Azurm Campus, 4800-058 Guimares, Portugal, Phone:(+351)938404554, Fax: (+351) 253 510 217, E-mail:[email protected].

2PhD Student, ISISE Institute for Sustainability and Innovation in Structural Engineering, University of Minho,School of Engineering, Civil Engineering Dept., Azurm Campus, 4800-058 Guimares, Portugal, Phone: (+351)938928308, Fax: (+351) 253 510 217, E-mail:[email protected].

3PhD Student, ISISE Institute for Sustainability and Innovation in Structural Engineering, University of Minho,School of Engineering, Civil Engineering Dept., Azurm Campus, 4800-058 Guimares, Portugal, Phone: (+351)938928308, Fax: (+351) 253 510 217, E-mail:[email protected].

4Full Professor, ISISEInstitute for Sustainability and Innovation in Structural Engineering, University of Minho,

School of Engineering, Civil Engineering Dept., Azurm Campus, 4800-058 Guimares, Portugal, Phone:(+351)93.840.4554, Fax: (+351) 253 510 210, E-mail:[email protected].

evised Manuscript

ck here to download Manuscript: Revised_Manuscript.docx Click here to view linked References
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increases at early ages, and eventually return to thermal equilibrium with the surrounding1

environment. These early temperature variations induce volumetric changes in concrete that are2

partially restrained by adjoining previously cast members, or even due to non-uniform3

temperature distributions within a concrete member itself. Such restraint to deformation may4

induce stresses that can be relevant enough to induce early age thermal cracking in concrete,5

which is usually unacceptable in view of aesthetics, durability and even structural performance.6

Contractors usually attempt to avoid this thermal cracking by adopting concrete compositions7

and construction schedules that maintain temperature gradients in concrete below prescribed8

limits, both along time and space [1]. It has however been recognized that such approach leads to9

erroneous conclusions, as several important issues are disregarded [1], such as the degree of10

restraint to deformation and the actual mechanical properties of concrete. In view of the11

limitations of these temperature-based criteria, it has been widely acknowledged [2] that more12

realistic crack risk assessments can be made through multi-physics approaches that encompass13

numerical simulation of temperatures and corresponding stresses in concrete: thermo-mechanical14

analyses. The use of thermo-mechanical simulation models allows the evaluation of alternative15

construction scenarios (for casting procedures, concrete mixes, environmental conditions), and16

thus permits the optimization of construction without compromising the safety in regard to17

thermal cracking. The numerical studies for the assessment of the optimum construction strategy18

frequently involve diminishing the temperature rise in concrete at early ages. In fact, if the19

thermal variation is diminished, the corresponding volumetric changes also decrease, as well as20

the developed stresses. The diminishment of early temperature rises is usually achieved by21

partial replacement of cement by additions as fly ash [3, 4], or by cooling water/aggregates22

before mixing operations[5-7], or even by introducing internal cooling pipes in concrete with23

cooling fluids (usually water) [8-12]. An attempt to use air as the cooling fluid in cooling pipes24

has been made in the 1990s by Hedlund and Groth [8, 9], who have shown the feasibility of25

such technique in thick columns. Nonetheless, no further application of such technique was26


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found in the literature, except for an in-situuse of air for localized cooling reported by Ishikawa1

et al.[12].2

Even though thermo-mechanical simulation approaches have been applied to concrete3

structures for decades [13-17], they involve several complexities and problematic issues4

particularly in view of the assessment of material properties and model parameters (thermal and5

mechanical), which frequently demand specific laboratory characterization: heat of hydration,6

thermal boundary coefficients, creep of concrete, evolution of E-modulus and tensile stress,7

among others. Even though several scientific works have been done either on thermal simulation8

[18], thermo-mechanical simulation [19]or monitoring the concrete behaviour at early ages [20-9

23], some combine the numerical simulation with experimental data obtained in laboratory [24-10

26]or with temperature monitoring for partial validation [27-31], whereas others go further and11

additionally include in-situ strain monitoring for validation [32]. Nonetheless, in the scope of12

internal cooling of concrete through the use of embedded pipes, no works were found to adopt13

holistic approaches that simultaneously include material characterization, in-situmonitoring of14

temperature/strain and thermo-mechanical simulation. The works that focus on concrete cooling15

with embedded pipes are mostly limited to thermal [33], or thermo-mechanical analyses only16

[34-37], and thermal [38-40] or thermo-mechanical analyses together with partial validation17

through in situmonitoring [12,41].18

The present paper pertains to a case study of the thermal stresses in the central wall of the19

entrance organ of a dam spillway. Such wall is 27.5m long, with a maximum width of 2.8m and20

height of 15.0m, with attention being given to the most unfavourable construction phase in21

which a 2.0m tall batch had to be made (total 150m3 of concrete). Due to the materials and22

equipment available at the construction site, it was decided to attempt internal cooling of23

concrete with air-cooled prestressing ducts placed longitudinally along the wall.24

The present paper regards to the in-depth study of the early age performance of concrete in25

the wall, encompassing laboratory thermal and mechanical characterization of concrete, as well26


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as in-situ monitoring of temperatures/strains and the corresponding thermo-mechanical1

simulation with the finite element method.2

The extensive laboratory characterization included quantification of the heat of hydration3

evolution (isothermal and semi-adiabatic calorimetry), evaluation of compressive/tensile strength4

and E-modulus through cube/cylinder testing, creep testing at several ages and assessment of5

mechanical activation energy. In particular regard to E-modulus testing, a methodology that6

allows continuous measurement of concrete E-modulus since casting (EMM-ARM [42]) was7

applied. This is a pioneering use of this methodology for the purpose of supporting stress8

simulation on concrete since its very early ages, with important advantages in regard to previous9

approaches that tend to extrapolate values of E-modulus at very early ages.10

A relatively complete monitoring program has been carried out in-situ, involving the use of11

20 temperature sensors and 7 vibrating wire strain gauges embedded in concrete. Particular12

attention was given to the evaluation of the effectiveness of the cooling system, with13

temperatures being measured at several points along the prestressing ducts, and with air velocity14

measurements taken with handheld anemometers.15

Bearing in mind the information gathered with the laboratory characterization and in-situ16

monitoring, a thermo-mechanical simulation was carried out with recourse to a three dimensional17

finite element model. Such simulation model included the explicit modelling of the cooling18

ducts, as well as the phased construction of the wall. The simulation model was made with19

DIANA software [43].20

21

2 THERMO-MECHANICAL MODEL22The thermo-mechanical simulation approach presented here has strong similarities with that23

described in a previous work [30]. Nonetheless, some particularities are distinct, namely: (i)24

solar radiation is explicitly considered according to a model based on the incidence angle of the25


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sun beams; (ii) the effect of internal cooling ducts is taken into account (iii) the evolution of1

mechanical properties is simulated according to the equivalent age concept (instead of the2

degree of hydration concept). The following sub-sections pertain to a general description of the3

modelling strategy with specific emphasis on topics (i) to (iii) mentioned above.4

2.1 Thermal model5The calculation of temperature fields in concrete is based on the heat balance equation,6

whose solution is made through the finite element method [44]:7

k T Q cT (1)

where kis the thermal conductivity, cis the volumetric specific heat and Tis the temperature.8

Q is the volumetric heat generation rate due to cement hydration, formulated as an Arrhenius9

type law [45]:10

RT

Ea

efAQ

)( (2)

where A is a rate constant, Ea is the apparent activation energy, is the degree of heat11

development (ratio between the heat Qreleased up to time tand the total heat Qfinalreleased upon12

completion of cement hydration),R= 8.314 Jmol-1K-1is the Boltzmanns constant and f() is a13

normalized function for heat.14

Thermal boundary conditions are applied through a prescribed flux per unit area qT15

formulated as [46]:16

T cr b eq h T T (3)

where hcr is a mixed convection-radiation boundary transfer coefficient, Tb is the boundary17

surface temperature and Teis the environmental temperature.18

The simulation of thermal inputs associated to solar radiation in concrete structures can be19

made with significant accuracy through the adoption of models that are readily used in20

meteorological sciences [31, 47]. Such models can take into account the effects of the spatial21


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relationship between the earth and sun at a given time of the day/year and thus predict the solar1

radiation that reaches a certain point on earth at sea level (i.e. low atmosphere). It is further2

possible to compute the angle between the sunbeam and any arbitrarily oriented/inclined surface,3

and evaluate the intake of energy throughout the day of such surface.4

The calculation of the solar energy that reaches earth at sea level, qm, is based on the solar5

constant, q0, which represents the total radiation energy received from the sun at a distance6

corresponding to 1 Astronomical Unit. Even though q0 varies slightly throughout the year by7

~7%, it is usually acceptable to consider q0=1367 Wm-2. The estimation of qm can be done8

through the following empirical equation [47, 48]:9

)sin(4.99.0

0

h

T

m

l

eqq

(4)

where Tl is the Linke turbidity factor that summarizes the turbidity of the atmosphere10

(attenuation of the direct beam solar radiation) and his the solar elevation that corresponds to the11

angle between the direction of the sunbeam and the idealized horizon. Tl is known to usually12

vary between 3 and 7, whereas hcan be calculated by taking into account latitude, date and time.13

Further geometrical considerations allow the calculation of the angle between an incident14

sunbeam and the vector orthogonal to an arbitrarily oriented/inclined surface, termed as i (see15

detailed description of models to calculate hand iin [44, 47]).16

Based on the knowledge of qmand iat a given instant, and considering the absorvity of the17

material of the target surface (S), it is possible to calculate the radiation energy qs that is18

actually absorbed:19

)cos(iqq mSS (5)

Another particularity of the present application in regard to previous works [30]is the use of20

prestressing ducts acting as cooling pipes. A formulation is thus necessary to describe the added21

internal heat fluxes that are caused by the presence of an embedded cooling pipe, which can be22


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expressed in the following energy balance equation, to be applied throughout the length zalong1

the pipe [49]:2

cp c c wdT

m c h P T T dz

(6)

where m is the mass flow rate of the coolant (air in this case), cp is the specific heat of the3

coolant, hcis the boundary transfer coefficient between the coolant and the surrounding concrete,4

Pis the perimeter of the cooling pipe, Tcis the temperature in the cooling pipe and Twis the bulk5

temperature of concrete around the cooling pipe. The mass flow rate of the coolant m can be6

estimated through the product of the cooling fluid density () by the fluid mean velocity (m) and7

by the cross sectional area of the cooling pipe (Ac). The implementation of equation (6) into a8

finite element software [43] brings further nonlinearities due to the interaction between the9

cooling fluid and the surrounding concrete, which results in progressive heating of the cooling10

fluid along the pipe.11

For updating age-dependent properties along time in the mechanical model, the equivalent12

age of concrete teqis adopted. Its formulation is based on an Arrhenius type equation established13

for a reference temperature Tref (usually 293.15K) [50]. For a given instant t, the equivalent age14

can be calculated as:15

dett

TTR

E

eq

ref

a

0

11

(7)

2.2

Mechanical model16The mechanical model is relatively similar to those adopted for time-dependent mechanical17

analysis of hardened concrete, except for some particularities associated to the facts that: (i) it is18

being preceded by a thermal analysis (de-coupled), with imposition of strains that are calculated19

with basis on the thermal dilation coefficient of concrete (T) and the previously calculated20

temperature field; (ii) there is a strong evolution of mechanical properties throughout the21

analysis, dully taken into account through the equivalent age concept; (iii) the strong viscoelastic22


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behaviour of concrete at early ages makes it necessary to use creep formulations that can provide1

adequate estimates within such time span. In specific regard to the last point (iii), basic creep of2

concrete was accounted for through the use of the Double Power Law (DPL), which has a3

reasonably good performance on both early age and long term time spans [51]:4

nm ttttEtE

ttJ ),

(),

(),

(),

(

1),

,(

0

1

0

(8)

where ),,( ttJ is the compliance function at time t for a load applied at instant,

t , ),(0 tE is the5

asymptotic elastic modulus, and 1, m and n are material parameters. Since drying creep is6

negligible for an application that only envisages early age behaviour, it was disregarded [52].7

Bearing in mind that the aim of the thermo-mechanical simulations is to assess the risk of8

cracking, the post-cracking behaviour is not considered relevant and it is thus not simulated.9

Thus, linear elastic behaviour (with creep) is considered for concrete both in compression and in10

tension.11

3 THE PARADELA DAM SPILLWAY: DESCRIPTION AND MONITORING123.1 Overview13

The Paradela dam, located in the North of Portugal, is a rockfill gravity dam built in the14

1950s,with 540m longitudinal development and maximum height of 112m from foundation.15

Due to recent hydraulic problems in one of the dams spillways, it was necessary to build a new16

complementary ski-jump spillway on the right margin of the river[53]. The case study reported17

in this paper concerns the cooling measures and assessment of cracking risk in the construction18

of the central wall of the spillway entrance.19


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3.2 Description of the spillway entrance13.2.1 Geometry and construction phasing2

The spillway functions in free surface conditions, and has two main entrances at the top level3

of the dam, each with 5.5m width, being separated by a hydro-dynamic shaped wall with4

maximum width of 2.8m and 17.4m height. A three dimensional representation of the entrance5

region of the dam spillway is shown in Figure 1a, whereas its corresponding plan view at6

approximately mid-height of the wall is depicted inFigure 1b. The reinforcement of the middle7

wall can be generally characterized by 16//200mm placed vertically and 12//200mm placed8

horizontally near each surface with a concrete cover of 60mm.9The construction of the wall was generally performed with 1.2m tall construction phases,10

with empirically defined waiting periods being defined by a target temperature in the core11

regions during the cooling period (approximately 27C, which corresponded to 17C above12

average daily temperature during construction). In order to minimize such waiting periods, an13

air-cooling system based on ventilated prestressing ducts placed horizontally was implemented,14

allowing lower peak temperatures and faster return to temperature equilibrium with the15

surrounding environment. The main scope of the present paper is the study of a specific16

construction phase that corresponds to the zone of embedment of the fixation parts of the sluice17

gates. Such fixation parts were approximately 2.5m tall, and it was thus desirable to perform a18

2.5m tall construction phase, labelled as 9thphase inFigure 2a. Due to its larger thickness, this19

construction phase is the critical one in terms of peak temperatures and cracking risk, being20

therefore the object of analysis.21

3.2.2 Materials22The wall of the spillway entrance was generally cast using concrete of class C30/37 [54]with23

the composition labelled as S1-D32 inTable 1. In the 9thconstruction phase, due to increased24

complexity of reinforcement near the downstream extremity of the wall (related to the salient25


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concrete blocks), two slightly different compositions with higher fluidity were used in the1

vicinity of such region, as shown inTable 1 (S3-D32 and S3-D16). In spite of this, the areas of2

most interest to this study (thickest regions of the wall and monitored sections) correspond to the3

upstream region. Therefore, and also taking into account the fact that the compositions have4

similarities, all the characterizations and modelling in the scope of this work pertain to mix S1-5

D32. Steel reinforcement was S400C [54], with characteristic yield stress of 400MPa.6

3.2.3 Cooling system7In view of the work reported by Hedlund and Groth [8, 9], which proposed the possibility of8

using ventilated prestressing ducts for concrete cooling, and bearing in mind the easy availability9

of the corresponding necessary equipment in the construction site of the wall, it was decided to10

test the feasibility of this kind of cooling technique. However, in view of practical limitations11

posed by contractor/owner, this pilot application of air-cooling system was slightly different12

from that of Hedlund and Groth [8, 9]. Instead of placing the tubes vertically along the wall, they13

were placed horizontally, even though this implied more limited cooling capacity as the length of14

tube along freshly cast concrete is longer. In the particular case of the 9thconstruction phase, a15

total of 6 prestressing ducts of 90mm diameter have been used, with their air intake being made16

horizontally at the downstream extremity of the wall, and the outtake made near the upstream17

extremity, on the top surface of the casting phase, in order to avoid a direct upstream-18

downstream potential leakage channel after construction. The overall path of the ducts is shown19

in the schemes of Figure 33, with the ducts labelled from T1 to T6. For ventilation, a 0.60m20

diameter fan was used, with 1200m3/h ventilation capacity, that collected air from the21

environment and blowed it into the ducts at an internal air speed of approximately 8.6m/s22

(measured with anemometer at the outtake of the ducts). The fan had to be placed in the23

downstream extremity of the wall due to practical constraints of the contractor. The cooling24

system was only started at the age of 14h after the end of casting operations to avoid introducing25

potentially undesirable vibrations to the freshly cast concrete before its structural setting time.26


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The ventilation system was disconnected 8.6 days after casting in view of the similarity of1

temperature between the walls core and the surrounding environment. After that, the2

prestressing ducts were filled with mortar using standard procedures [55].3

It is remarked that this cooling system had been previously tested in the 8th

phase of casting,4

with three prestressing ducts placed at mid-height. Details on this test can be found elsewhere5

[56].6

3.3 Monitoring and material characterization73.3.1 General remarks8

In order to better understand the effectiveness of the cooling system, its influence on the9

cracking risk, and assess the capabilities of the adopted numerical simulation strategy, an10

extensive set of actions has been carried out, comprising in-situ internal monitoring of11

temperatures/strains of concrete, in-situ validation tests, as well as laboratory material12

characterization.13

3.3.2 Temperature monitoring14Temperatures inside the 9th construction phase have been monitored with K-type15

thermocouples, aiming particularly at assessing temperature profiles in a region near the16

maximum width of the wall (i.e. at approximately 5.3m from the upstream extremity: section A-17

Aas identified inFigure 3a). The placement of temperature sensors in section A-Ais depicted18

in the scheme ofFigure 4a, where thermocouples are identified by the prefix TC. Temperatures19

in the locations labelled as VW inFigure 4ahave also been monitored with resistive temperature20

sensors, as these are the locations of vibrating wire strain gauges, which contain internal21

temperature sensors for strain compensation. The internal air temperature of ducts T1-T3 has22

been monitored both in section A-A and in neighbouring areas, as shown in Figure 4b.23

Environmental temperature (dry-bulb) has been assessed with a thermocouple.24


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Monitoring was carried out since the instant of casting during a period of 10 days, and the1

measurement frequency was set to 1 reading per each 30 minutes. The internally monitored2

temperatures in concrete are shown in Figure 5 for a vertical and a horizontal alignment of3

sensors that pass through sensor VW3. It is remarked that the results of this figure and all4

upcoming figures of the paper (either concerning experimental results or numerical simulation5

results) have their corresponding time axis zeroed in regard to the instant at which the casting6

operations of the studied phase were finished. From Figure 5 it can be seen that the initial7

temperature of concrete was ~15C, and the peak temperature was approximately 42C in the8

core regions (VW3 and VW5). Furthermore, it can be observed that the ascending branch of9

temperature development is clearly affected at the age of 14 hours, when the cooling system is10

activated. In specific regard to the vertical profile of temperatures shown in Figure 5a, the11

expectable behaviour was captured: the core region has the highest peak temperatures (VW3,12

VW5), whereas a decrease trend is seen towards the top surface. In fact, sensors TC6 and TC713

exhibit maximum temperatures of ~36C, while VW6 (near the top surface) has the lowest peak14

temperature (circa 27C). Near the bottom surface of this construction phase, sensor VW115

highlights the importance of the heat storage effect caused by the previously cast concrete: in16

fact, even though the temperature peak is lower than that of the core regions, it occurs later and17

the heat loss rate observed afterwards is lower than in other regions. It should also be remarked18

that all sensors are almost in equilibrium with environmental temperature by the age of 8 days.19

In regard to the temperature development in the sensors located along a horizontal alignment,20

shown in Figure 5b, it can be noticed that the sensors located in the vicinity of vertical21

boundaries exhibit lower temperature variations (VW4 and TC5), with temperature peaks under22

35C. It is interesting to observe that TC5 is placed nearer the surface (15cm apart) than the case23

of VW4 (20cm apart), and consequently the temperature peak of VW4 is slightly higher. By24

looking at the temperature evolution after the age of 8 days (heat of hydration has been25

dissipated), it can be seen that temperatures in VW4 remain higher than those of TC5 due to a26


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combination of two main reasons: VW4 is located in the vicinity of a surface oriented to1

southeast, which is bound to receive more energy through solar radiation than the surface near2

TC5, which is oriented to northwest; VW4 is 5cm deeper than TC5, thus being slightly nearer3

the inner core (higher thermal inertia).4

The temperature evolution along the air inside duct T3 measured by sensors TC3, TC9 and5

TC10 (Figure 4b) is shown in Figure 6, where the environmental temperature and the6

temperature in VW5 (hottest region in concrete) are also represented for comparative purposes.7

The information provided by such figure allows the clear identification of the instant at which air8

ventilation began (14 hours), as the rate of temperature rise is clearly disrupted inside the duct.9

Furthermore, the rising temperature tendency along the ducts length is identifiable, as the10

temperature is consistently higher in TC3 in regard to TC10, and in TC9 in regard to TC3.11

Taking as example the temperatures recorded at the instant of peak temperature (1.88 days), TC912

measured a temperature of 32.2C, whereas TC10 indicated a temperature of 29.4C. This13

represents a shift in temperature of approximately 3C in 3.5m length of duct. At three instants of14

this study, temperatures were measured also at the entrance of T3 (x=0) and at x=1m through the15

use of a handheld temperature probe (PT1000). By joining such data with the results of TC3,16

TC9 and TC10, it was possible to plot a temperature profile along the duct for t=3.84d, t=4.56d17

and t=6.58d see Figure 6b. It can be observed that the temperature at the inlet of the tube18

matches the environmental temperature, and that the heating of the air along the tube is strongly19

dependent on the combination of environmental temperature and internal temperature of20

concrete. In fact, in the most unfavourable situation shown by Figure 6b, air was heated from21

~11C at the air inlet to ~26C at a point located 25m away from the air inlet (t=4.56d). This22

increase of air temperature is bound to reduce its cooling capacity. However, in spite of such23

diminishment of cooling capacity, the temperature of the air in the hotter regions of the duct24

remained at least 10C below that of concrete during the periods at which temperature in25


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concrete was near its peak (see t=1d to t=3d in Figure 6a), showing that the heat removal1

potential was not negligible at all.2

The observed diminishment of cooling capacity along the length of the duct highlights the3

fact that the adopted configuration for the tubes does not maximize cooling capacity, which4

would be conversely maximized if the length of tube inside concrete had been minimized. Such5

goal could have been achieved by providing a vertical arrangement for the tubes and introducing6

more individual smaller tubes.7

3.3.3 Strain monitoring8Strain measurement was carried out with vibrating wire strain gauges of metallic casing with9

14cm reference length (TES/5.5/T Gage Technique). Past laboratory tests and in-situ10

applications [44,57, 58]have shown that this kind of sensor is robust and adequate for strain11

measurement in concrete at early ages. The strain gauges were placed at the locations identified12

in Figure 4a (VW1 to VW6), dully positioned in order to measure strains in the longitudinal13

direction of the wall. VW7 has distinct intents and shall be specifically addressed later.14

Measurements were taken with the same datalogger and at the same sampling rate as adopted for15

the temperature sensors.16

One important issue to tackle is the zeroing of the measured strains. In fact, before concrete17

sets and has enough stiffness to drive the sensor into the same deformation state, the18

measurements taken by the sensor do not have any relevant physical meaning. It is thus19

necessary to assess the instant of solidarization (i.e. the full bond) between concrete and the20

sensor. In a previous work [57], the zeroing operation has been made by assessing the instant at21

which two sensors with different casing (plastic and metallic), placed under the same conditions22

inside concrete, started yielding the same results. This means that both sensors are solidarized (as23

the plastic sensor is bound to solidarize earlier due to its smaller stiffness). Since the plastic24

cased sensor was not available, an alternative methodology for zeroing the data was25

implemented. By interpretation of the findings reported in [57] it can be considered that the26


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solidarization instant coincides with a progressive change in the derivate of strain variation1

detected by the sensor, that can be obtained by geometrical intersection of tangents of measured2

strains, as shown inFigure 7.It was decided to use such zeroingcriterion in the scope of this3

research work. As a result of the application of such rule, the solidarization instant of each sensor4

VW1 to VW5 was, respectively: 0.17, 0.19, 0.19, 0.20 and 0.17 days (due to malfunctioning of5

the VW6, the strain results from this sensor are not available). The solidarization instants seem6

coherent, since they have a trend to increase with the distance of the sensor from the bottom7

surface of the casting block, due to the natural delay in its involvement by concrete during the8

casting process.9

The measured strains in sensors VW1 to VW5 are shown inFigure 8.Even though the strain10

output is dependent on several factors that interact with each other (thermal deformation,11

restraint, creep), it is possible to find a set of common points and reasoning. Overall, all12

deformations are strongly commanded by the temperature variation, following the same kinetics.13

Sensors VW2, VW3 and VW5, which are located in regions near the core of the walls cross14

section and had similar temperature development histories, also have similar strain15

developments. This is bound to be caused by similar thermal deformations and restraints for16

these locations of measurement. The smallest deformations are recorded in VW1, which is17

located in the bottom of the casting phase (i.e. near the existing concrete) and thus having less18

temperature rise (thus less expansion), while being more restrained by the existing concrete19

below at lower temperatures. Finally VW4, which is near the surface and thus has lower20

temperature rises (maximum temperature of ~33C), also has a smaller deformation variation21

when compared to VW2, VW3 and VW5.22

In order to assess free deformations of the concrete used in the construction (associated to23

unrestrained autogenous shrinkage and thermal deformations), strain was measured in a concrete24

cylinder placed in specially devised conditions. The concrete cylinder (150mm diameter and25

300mm tall), cast simultaneously with the studied construction phase and by using the same26


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concrete, was cast into a special mould internally coated with a soft membrane (with lids also1

coated with such material). A strain gauge was placed inside the mould to measure longitudinal2

strainssee photo of the open mould inFigure 9a. After casting concrete into such mould, it was3

placed horizontally inside the studied construction phase (during its casting procedures) at the4

location that is identified as VW7 inFigure 4a. This kind of procedure, here termed as use of a5

no-stress specimen, has been reported in Choi et al. [59], and it allows to measure free6

deformations of concrete, which can in turn be used to assess the thermal dilation coefficient,7

TDC (provided that the temperature inside the concrete cylinder is relatively uniform, and8

autogeneous shrinkage deformations are known). Unfortunately, due to undetermined causes, the9

output of the sensor could not be read during the first 0.8 days, and thus the reported data only10

starts at such age, as shown inFigure 9b.11

3.3.4 Heat generation and activation energy12In an extensive experimental program for the characterization of the cements marketed in13

Portugal, Azenha [44] has reported a library of heat generation obtained through isothermal14

calorimetry under several temperatures for plain cement pastes with w/c=0.5. The cement used in15

the construction concerned in this paper was also characterized (same supplier and16

manufacturing plant), and the resulting information for calorimetry tests under 20C, 30C, 40C17

and 50C is shown inFigure 10.A reasonable estimate of the heat generated by concrete can be18

obtained by multiplying the heat generation reported inFigure 10by the volumetric content of19

cement, which is of 224kg/m3

. By using the speed method algorithm [44,60], the necessary data20

for the numerical simulation of heat generation according to equation 2 was obtained:21

Ea= 37.31 kJ/mol, A = 4.989109W/m3, Qpot= 8.295107J/m

3, and function f() characterized22

by the following set of data [; f()] = [0.00; 0.00], [0.05; 0.58], [0.10; 0.85], [0.15; 0.98], [0.20;23

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],24

[1.00; 0.00]. Even though this data pertains to CEM I 42.5R of the same company that supplied25

the cement to this construction site, there may be deviations caused by inevitable variations in26


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the characteristics of the cement. Also, the extrapolation procedure mentioned above did not take1

into account the presence of fly ash in the mix (96kg/m3), which may have non-negligible effects2

on the heat generation potential and hydration kinetics. Therefore, in order to assess the potential3

importance of such deviations, a semi-adiabatic test was conducted in-situ (simultaneously with4

the casting operations) in a 30cm edge concrete cube, duly isolated by 2.1cm thick plywood and5

12cm of polystyrene boards. The results of such semi-adiabatic calorimetry test shall be6

addressed in section 4, upon the simulation of its temperature development through the finite7

element method.8

3.3.5 Complementary laboratory characterization9Compressive strength evolution was assessed with concrete cubes (150mm edge) at the ages10

of 1, 3, 7 and 29 days, whereas tensile strength was measured with splitting tests on cylinders11

(150mm diameter and 300mm tall) and at the same ages. It is remarked that testing at the12

reference age of 28 days was not possible due to laboratorial constraints. The evolution of both13

tensile and compressive strength for concrete cured at 20C (saturated conditions) is shown in14

Figure 11a (average results of three specimens at each age). In order to assess the activation15

energy suitable for compressive strength maturity estimations, a set of concrete cubes was cured16

at 40C with the compressive strength measured at the same ages. The corresponding results are17

also shown in Figure 11a. By applying the equivalent age concept [61], together with the18

superposition method [60], it was possible to asses that the activation energy based on19

mechanical testing has the value of 37 kJ/mol, which is rather consistent with the activation20

energy obtained through isothermal calorimetry for the same cement (yet without fly ash),21

37.31 kJ/mol [44]. Such coincidence in activation energy for thermal and mechanical phenomena22

had already been reported by Ulm and Coussy [62].23

Basic creep was assessed in creep rigs on prismatic specimens (sealed) with dimensions24

15cm15cm60cm, loaded at 30%~40% of the concrete compressive strength and internally25


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monitored with vibrating wire strain gages. Such creep tests were conducted at the ages of 1.1,1

3.3 and 7.3 days, and the corresponding specific creep curves are shown inFigure 11b.2

The experimental program also included a single specimen for measurement of autogenous3

shrinkage. Such specimen was a 150mm diameter and 300mm long cylinder, internally4

instrumented with a vibrating wire strain gage, which was kept in its formwork during the5

experiment and sealed with a plastic film on the top surface. Unfortunately, two factors6

contributed to render the results of this specimen unusable for this research: on one hand, the7

monitoring only could be started at the age of 2 days in the laboratory due to unavailability of8

datalogging system; on the other hand, the measurements of autogenous taken since t=2 days9

were disturbed by an inefficient sealing, which promoted undesired drying of the specimen. This10

was not considered a critical problem in view of the low values of autogenous shrinkage that are11

usually expectable in concretes of low cement content and high w/c ratio [63-65].12

3.3.6 Continuous monitoring of concrete stiffness13The evolution of elasticity modulus along time was measured through compressive cyclic14

testing in concrete cylinders (150mm diameter and 300m tall) at the ages of 1, 3, 7, 15 and 2915

days. Concomitantly, E-modulus of concrete was continuously assessed through a methodology16

termed as EMM-ARM (Elasticity Modulus Measurement through Ambient Response Method).17

This methodology has been developed by Azenha et al.[42]and consists in a variant to classic18

resonant frequencies that allows the quantification of E-modulus continuously since the instant19

of casting of the specimen inside the testing mould. The basic principle of EMM-ARM is the20

following: the specimen is cast inside the testing mould, which is in turn placed in simply21

supported conditions and continuously subject to modal identification (using accelerometers)22

without any explicit excitation of the beam, as ambient vibration suffices. As concrete hardens,23

the first resonant frequency of the composite beam evolves, and the stiffness of concrete can be24

inferred by applying the equations of motion. Details about the testing setup and procedure25

applied for the concrete of this spillway application can be found in [56, 66, 67], as it26


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corresponds to an improved version of the originally devised test (a steel mould is used). The1

collected results with EMM-ARM and cyclic compression tests on cylinders are shown inFigure2

12,where the feasibility of EMM-ARM is confirmed in view of the resemblance of results. Also,3

the richness of information that can be obtained through EMM-ARM represents an added value4

for the numerical simulation.5

4 NUMERICAL MODELLING64.1 Geometry, mesh, materials, initial/boundary conditions and time integration74.1.1 Geometry of the model and finite element mesh8

A cross-sectional scheme of the model for simulation is shown in Figure 13a, where the9

construction stages considered in the analysis can be observed. The first stage of the model10

encompasses all concrete until the 7thphase of concreting (inclusive), considered as hardened11

concrete, together with the 8thphase of concrete evaluated as freshly cast concrete. The second12

and third stages correspond to the 9 thand 10thphases of concreting respectively. This strategy13

diminishes the computational cost of the model without relevant effect on the accuracy of results.14

For similar reasons, the underlying subgrade is not modelled, as it is far from the construction15

phases of interest. Due to the geometrical symmetry of the wall, a longitudinal plane of16

symmetry is considered, identifiable by a ZX plane inFigure 13.17

The simulation was made with a 3D finite element model comprising rectangular brick FE of18

8 nodes (222 integration scheme) for concrete in the thermal model, and coincident 20 nodes19

brick FE (333 integration scheme) in the mechanical analysis. Convective boundaries were20

modelled with 4 node planar elements (22 integration scheme), and the cooling ducts were21

considered with linear elements of 2 nodes (2 point integration scheme) [43]. The schematic22

representation of geometry, casting phases and cooling duct location for section A -A is shown23

in Figure 13b. The reader is reminded that phase 8 also had cooling ducts, according to the24


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description of section 3.2.3.The longitudinal layout of all the ducts was considered straight on1

an horizontal plane, in correspondence to simplifications in the vicinity of the extremities of the2

wall.3

It should be remarked that due to the phased analysis, the mesh evolved along time with4

some elements/boundaries being activated/de-activated (e.g. the convective top boundary of a5

given phase is de-activated upon the beginning of the next casting phase). The mesh adopted for6

this simulation is shown in Figure 13c, with a total of 18738 elements and 66037 nodes. As7

consequence of the symmetry simplification, it was considered that the perimeter of the tube8

elements located in the symmetry plane of the model was equal to half the actual perimeter of the9

tubes. As solar radiation does not represent a symmetrical energy input to the structure, the10

symmetry simplification adopted is not truly valid. However, as solar radiation has most of its11

effect near the surface, it was decided to keep the symmetry simplification by considering the12

southeast half of the wall, which is most subject to solar radiation effects.13

4.1.2 Materials (thermal and mechanical properties)14The thermal conductivity and specific heat of concrete were estimated with basis on the15

pondered average of the corresponding thermal properties of the constituent materials of the mix16

[44, 68]. The adopted values for k and c for concrete were, respectively 2.40 W/mK and17

2.4106 J/m3K. Even though it is known that these thermal properties suffer variations during18

early ages [69-75], the adopted modelling approach considers them constant in view of the19

conclusions of the parametric analyses reported by Azenha [44], where a relatively small impact20

of considering evolving kand cwas found on computed temperatures in hardening concrete.21

In regard to the data for heat of hydration, the parameters mentioned in section3.3.4 were used.22

The adequacy of these parameters was evaluated through the semi-adiabatic calorimeter23

described in the same section, whose behaviour was simulated through a FE simulation model24

that explicitly considered the extruded polystyrene (XPS) and wood walls of the calorimeter. The25

material modelling parameters for concrete coincide with those herein described, whereas26


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additional information and mesh representation are shown in Figure 14a. The results of the1

simulation of the calorimeter were quite coherent with those collected experimentally, as seen in2

Figure 14b, leading to the confirmation that the strategy described in section3.3.4 to determine3

the heat generation and activation energy was adequate for defining the heat of hydration4

modelling parameters adopted in the simulations.5

The thermal dilation coefficient (TDC) of concrete was assessed with basis on the no-stress6

specimen described in section 3.3.3. However, in order to obtain the thermal deformation and7

calculate the TDC, it was necessary to subtract the autogenous shrinkage deformation from the8

total deformation. As data on autogenous shrinkage was not available, an estimate of the9

autogenous shrinkage evolution based on Eurocode 2 [54]was used. Another issue to take into10

account is the fact that the TDC of concrete is not constant during the first hours of age. In fact,11

several authors have dealt with this subject, and it generally agreed that the initial TDC tends to12

be larger than that of hardened concrete, and tends to decrease sharply within the first 12 to 2413

hours of age, reaching then the plateau level corresponding to hardened concrete [69,73,76].14

The no-stress specimen cast in the scope of this research cannot be used to estimate TDC at15

concrete early ages, due to the absence of data in the first 8 hours reported in Figure 9b.16

Nonetheless, as the peak temperature occurred later than 24 hours age, and full data is available17

for temperatures and strains occurred in such period, calculations could be made under the18

assumption that TDC was already at its plateau value. The TDC was estimated between instants19

t=2.0d (peak temperature) and t=8.0d (local minimum) as shown in Figure 9b, and the20

autogenous shrinkage strain variation in such period was estimated to be of 8.45, (considering21

fcm=42.3MPa in Eurocode 2 [54]). The estimated constant TDC to be used in the numerical22

simulation was 11.07C. Nonetheless, since it is known that TDC varies during the first 2423

hours of age, an alternative formulation for the evaluation of the TDC was considered, based on24

experimental evidence reported by Laplante and Boulay [77]. Therefore, this alternative25


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formulation considers that during the first 16h the TDC varies according to1

2( ) 0.16 4.88 48.93TDC t t t (tin hours), and remains constant after such age.2

The E-modulus evolution of concrete for the simulation model was directly extracted from3

EMM-ARM data reported in Figure 12, whereas creep modelling was made through the4

adjustment of DPL parameters to the creep data experiments. The best-fit creep parameters and5

their adjustment to the experimental data are shown in Figure 11b. Prestressing ducts were6

modelled with consideration of their inner perimeter of 283mm (90mm inner diameter). Steel7

reinforcement was disregarded in temperature calculations due to its low interference in8

temperature development [44, 52]. Regarding mechanical field simulations, steel was not9

considered because post-cracking behaviour was not sought. In the non-cracked stage, the10

similitude in TDC of steel in regard to that of hardened concrete strongly minimizes the restraint11

to concrete thermal deformation, thus rendering the effect of reinforcement negligible for the12

computation of thermal stresses at early ages [52].13

4.1.3 Initial/boundary conditions and construction phasing14In what concerns the boundary conditions in the thermal problem, it was assumed that an15

average wind speed of 2.5m/s occurred (confirmed during 3 days with an anemometer) and the16

resulting convection/radiation coefficient for concrete surfaces in contact with the environment17

was estimated to be 15.25 2W m Kin accordance to the predictive formula of Branco et al.[78].18

For the particular case of formworks surrounding concrete, an equivalent boundary convection19

coefficient was adopted according to an electrical analogy [46]. Bearing in mind that the20

formworks were made of wood (kwood=0.175W/m2K) and their thickness was 18mm, the21

resulting equivalent boundary coefficient was 6W/m2K. Formwork was applied to the vertical22

surfaces of each casting stage during the first 7 days of age, and removed afterwards. All23

convective boundaries were subjected throughout the analysis to the environmental temperature24


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that was monitored in-situseeFigure 5.The symmetry plane of the model was considered as1

an adiabatic boundary.2

In what regards to solar radiation intake, the several surface directions of the model were3

taken into account, and the absorbed radiation was calculated according to the model described4

in section 2.1. Absorvity of concrete was considered as 0.6 [79], the latitude was 41.77 N, and5

the casting date was 28/03/2011. The Linke turbidity factor was considered as 2.5, and its6

feasibility was confirmed by comparing computed solar radiation on horizontal surfaces with7

solar radiation data from a nearby weather station (Rio Torto Station) in conditions of clear8

skies. The effect of cloudiness was taken into account by normalizing the predictions of the9

adopted solar radiation model according to information obtained from the piranometer of the10

neighbouring weather station. The diminishment of solar absorption caused by shadows cast by11

neighbouring objects was considered negligible throughout the entire day, as the wall was one of12

the tallest elements in the landscape.13

Bearing in mind the construction phases taken into account in this calculation model14

(identified in Figure 13a), the initial temperatures were considered as follows. For the existing15

concrete at the beginning of analysis, it was assumed that concrete was already in thermal16

equilibrium with the environment, and so, the average daily temperature of the preceding week17

(14.5C) was considered for the existing concrete. For the subsequent stages of construction, the18

initial temperature of concrete was obtained from in-situ monitoring (average value), which was19

also of approximately 14.5C.20

In regard to the cooling ducts, due to a limitation of the adopted software, the inlet21

temperature had to be considered constant, equal to the average environmental temperature22

during the time in which the tubes were active. The following temperatures were considered for23

each phase: 8th: 7.13C and 9th: 13.91C. The internal convection coefficient in the ducts was24

obtained with basis on their internal air speed of 8.6m/s, which according to the studies of25

Hedlund and Groth [8] should correspond to a convection coefficient of 30.0W/m2K. The26


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cooling duct elements were activated at each construction phase when the surrounding concrete1

had age of 14 hours.2

Taking into account the directions of the axes of the coordinate system presented in Figure3

13a, the mechanical boundary conditions consisted in placing Z direction supports on the bottom4

surface of the wall and Y supports in all the elements of the symmetry plane.5

The time step strategy adopted for the model consisted in considering the initial time6

coincident with the casting instant of 8th phase. Casting of the 9th and 10th phases were7

considered at the relative instants t=24 days and t=41 days in accordance to the actual8

construction. All analyses were conducted with a constant time step of 1h duration, even though9

some localized adjustments were necessary in view of construction phasing and ventilation10

activation. Nonetheless, all adjustments were carefully made to assure that the duration of all11

time steps remained under 1h.12

4.2 Results and discussion13The presentation and discussion of results is centred in the 9th construction phase, with14

particular emphasis for comparisons between monitored and simulated temperatures/strains. The15

temperature simulations in concrete were quite coherent with the monitored ones, as it can be16

confirmed for a set of representative locations (VW1, VW2, VW4 and VW5), whose results are17

shown inFigure 15.In fact the largest deviations regarding the monitored temperatures during18

the entire calculation always remain under 4C, thus providing confirmation of the feasibility of19

the modelling strategy, particularly in regard to the cooling capacity of the ventilated ducts.20

Based on the confidence gained on the thermal simulation model, a further numerical simulation21

was made, in which the effect of the cooling ducts was disregarded. The corresponding results22

for the hottest region of both models are shown in Figure 15b. It can be confirmed that the23

inclusion of the duct had a twofold effect: not only was the peak temperature diminished by 5C24


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with benefits for cracking safety, but also the return of the internal temperature to thermal1

equilibrium with the environment was accelerated, with advantages for construction phasing.2

The fact that the calculated temperatures matched well the monitored ones is a solid starting3

point for the analysis of results of the mechanical simulation, as any detected deviations are4

bound to be solely attributed to issues in the mechanical simulation itself. The calculated and5

measured strains for the same set of sensors that has just been discussed for temperature6

development are shown in Figure 16. The experimentally measured strains in this figure are7

represented by their value according to the zeroing procedure mentioned in Section 3.3.3 (Figure8

7), but also with a lower and upper bound related to possible uncertainties in the instant for9

zeroing of the sensors output of 2 hours. It can be seen that all the computed strains with10

consideration of constant TDC underestimate the peak strain at 1.96 day, but the post-peak11

kinetics seems to have been well captured. As the constant TDC assumption may lead to12

underestimations of early strain development [57], a further calculation was made using a13

plausible TDC evolution during the first 24 hours, as discussed in 4.1.2. The corresponding14

simulation results are shown inFigure 16,where a better overall fit is seen between experimental15

and calculation data (particularly for core regions). Even though the variable TDC was not based16

on experimental evidence obtained in the scope of this research, it is feasible to assume that a17

significant part of the strain deviations regarding experimental results can be explained by the18

variable TDC at early ages. It has to be kept into consideration that another possible source of19

deviation of results may be related to the instant at which measured strains were zeroed, which20

can be debatable. Nonetheless, the adequate prediction of strains that was attained is a good21

indication of the feasibility of the computed stresses which are to be analysed.22

The discussion of cracking risk is now addressed by comparing the computed principal23

tensile stress at the most unfavourable part of the model (located in the core region: x=8.5m,24

y=0.36m, z=11.7m), as shown in Figure 17. This figure contains the equivalent age-adjusted25

measured evolution of the tensile strength, and the calculation results for the cases of constant26


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and variable TDC, as well as the case of inexistent cooling ducts and constant TDC. The first1

comment that can be made fromFigure 17,is that the consideration of constant or variable TDC2

had very marginal effect on the results. The reason for the very small difference is bound to be3

related with the very low stress level that is induced during the first 24 hours, as the E-modulus4

of concrete is still very small and creep/relaxation is very high. Regardless of the comparison5

between these two models, it can be observed fromFigure 17 that the ratio between the tensile6

stress and the tensile strength of concrete at the most unfavourable instant is of approximately7

0.9, which corresponds to a significant cracking risk. Nonetheless, even though the cracking risk8

was higher than the desirable one, the structure did not present thermal cracks neither later9

through-cracks (evaluations made until 2 years after casting). It should be remarked that this10

point of stress analysis was the most unfavourable one within the structure, and significantly11

lower cracking risks were calculated for distinct regions, resulting in a global scenario of much12

more cracking safety (compared to a single-point analysis).13

Interestingly, a simulation of the same construction situation without consideration of14

cooling ducts would yield to higher cracking risk at the same location (as seen inFigure 17),15

with the ratio between the tensile stress and the tensile strength of concrete reaching 1.2.16

Therefore, if the calculations made here are considered trustworthy, the use of the cooling ducts17

may have been the differentiating factor that avoided a cracking scenario in this concrete lift.18

5 CONCLUSIONS19A case study regarding the assessment of the cracking risk of a thick wall in the entrance of a20

dam spillway, internally cooled with air-filled prestressing ducts, was presented in this paper.21

The use of ventilated prestressing ducts is considered more straightforward than water ducts due22

to the often easy availability of ducts and fans in construction works associated to massive23

concrete structures.24


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The case study has involved a comprehensive experimental part, including laboratory1

characterization of materials and in-situmonitoring for temperatures and strains. A 3D thermo-2

mechanical simulation of the construction phasing was shown, with input data duly based in the3

laboratory characterization program.4

In regard to previous works reported in the literature, the work presented in this paper has its5

main original contributions in the following fields: (i) the EMM-ARM methodology for6

continuous monitoring of concrete E-modulus since casting was applied for the first time as a7

characterization tool for stress simulation in concrete at early ages, thus enhancing the quality of8

input data; (ii) this is the first reported application of horizontally placed air-cooling pipes, with9

its efficiency being assessed and numerically simulated; (iii) the work reported here is relatively10

unprecedented in view of its holistic approach, with the authors being involved in all tasks of11

laboratory characterization (allowing sustainable estimates of material properties and modelling12

strategy for numerical simulations), field monitoring and numerical simulation with thermo-13

mechanical analysis.14

It is nonetheless acknowledged that a significant research effort is still necessary in regard to15

the creep monitoring and modelling at early ages, both in view of the effects of early hydration16

and in view of temperature effects on creep. Even though in-depth analyses of these particular17

issues of creep have not been included in this paper, the authors consider them to be of critical18

importance for adequate stress simulation in hardening concrete, thus demanding further19

research works.20

The numerical simulation results were compared to those collected by the in-situ monitoring21

and good coherences were observed both in terms of temperature and strain, providing good22

prospects in regard to the simulation capabilities of the models and the soundness of the23

experimentally obtained data. The monitoring/simulation results allowed concluding that the24

effectiveness of the air cooling system with horizontally placed pipes is limited in view of the25

significant heating that air suffers along the first meters of tube, thus diminishing its capacity of26


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cooling further regions of concrete. Also, the duct efficiency ends up being quite dependent of1

the environmental temperature, which cannot be easily anticipated during planning. Nonetheless,2

in spite of the acknowledged limitations of air cooling, it may prove quite feasible in relatively3

cool climates and small lengths of embedment (e.g.: 10m or less).4

Furthermore, the risk of cracking on the studied construction phase has resulted acceptable in5

most of its regions, even though a non-negligible risk of internal cracking was observed in some6

regions. The fact that no surface or through cracking was observable in the construction7

corroborates the cracking risk evaluation. It was also concluded that the same construction8

phasing without the use of the cooling ducts had a significantly higher cracking probability, thus9

confirming the usefulness of the cooling system.10

Finally, it is also worth remarking that a parametric analysis regarding the possibility of11

considering variable thermal dilation coefficient of concrete along hydration has been carried12

out. The outcome of such parametric analysis seems to point out a relatively low impact of13

admitting the thermal dilation coefficient as constant along hydration on the corresponding14

computed stresses, thus validating such simplification in this case study.15

6 ACKNOWLEDGEMENTS16Funding provided by the Portuguese Foundation for Science and Technology to the Research17

Unit ISISE, to the second author through the PhD grant SFRH/BD/64415/2009, and to the18

research projects PTDC/ECM/099250/2008 and QREN number 5387, LEGOUSE, is gratefully19

acknowledged. The kind assistance of the contractor (Teixeira Duarte S.A.) and the owner (EDP20

Eletricidade de Portugal) are also deeply appreciated. The contribution of ngelo Costa to the21

experimental work here reported is also gratefully acknowledged.22


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