models for predicting thermal properties and temperature...

MODELS FOR PREDICTING THERMAL PROPERTIES

AND TEMPERATURE IN MASS CONCRETE

CONTAINING GROUND GRANULATED BLAST

FURNACE SLAG

BY

AROSHA DABARERA

A THESIS SUBMITTED IN PARTIAL FULFILLMENT OF THE

REQUIREMENTS FOR THE DEGREE OF MASTER OF SCIENCE

(ENGINEERING AND TECHNOLOGY)

SIRINDHORN INTERNATIONAL INSTITUTE OF TECHNOLOGY

THAMMASAT UNIVERSITY

ACADEMIC YEAR 2015

ii

Acknowledgements

The author wishes to express his profound gratitude to his advisor,

Asst. Prof. Dr. Warangkana Saengsoy, for her precious guidance throughout

the research work. Her continuous advice, encouragement and support made

the authors work to complete within his study period.

Sincere appreciation is contributed to Prof. Dr. Somnuk

Tangtermsirikul for his thorough and invaluable consultation during the study

period. The guidance given by Dr. Krittiya Kaewmanee throughout the study is

acknowledged by the author. Moreover, the author would like to extend his

gratitude for suggestions given by external committee member Asst. Prof.

Saranagon Hemavibool.

The author wishes to extend his gratitude towards Sirindhorn

International Institute of Technology (SIIT), Thammasat University which

enabled him to learn from international professionals and provided scholarship

during the study period. Special appreciation is forwarded to all researchers in

Construction and Maintenance Technology Research Center (CONTEC) for

kind encouragement and suggestions during his study.

The acknowledgement is extended to Taiheiyo Cement

Corporation, Japan which consulted and partially funded for this project. The

author would like to thank project consultants Mr. Ryuichiroh Kuga and Ms.

Kanako Mori for their kind support towards this study.

The author gratefully appreciates his colleagues and friends who

provided a pleasant environment to study and enjoy his stay in Thailand.

Finally sincere gratitude and utmost admiration is dedicated to the author’s

parents and family members who supported all through his life.

iii

Abstract

MODELS FOR PREDICTING THERMAL PROPERTIES AND

TEMPERATURE IN MASS CONCRETE CONTAINING GROUND

GRANULATED BLAST FURNACE SLAG

by

AROSHA DABARERA

Bachelor of Science in Civil and Infrastructure Engineering, AIT, 2014

Ground Granulated Blast Furnace Slag is widely used as a partial

replacement of cement in mass concrete structures. However, the effect of slag

towards thermal properties in mass concrete is not extensively investigated.

Predicting thermal properties and adiabatic temperature rise are essentially

useful for investigating thermal cracking potential especially at early stage of

mass concrete. Existing prediction methods and models have some problems

such as constant thermal properties are mostly utilized for predicting

temperature rise of mass concrete. These assumptions lead to errors and

inaccurate predictions of temperature and thermal cracking potentials

especially at the early stage where thermal properties tend to change rapidly.

Therefore accurate modelling of thermal properties could be beneficial to

provide solid background to accurately predicting thermal cracking of mass

concrete.

This study is aimed to develop time-dependent models for predicting

hydration degrees of cement and slag which are vital parameters in modeling

many properties of concrete. Then the models of free water, specific heat,

thermal conductivity, coefficient of thermal expansion, and total heat

generation of concrete including slag are modified. These models are then

composed to predict the adiabatic temperature rise of mass concrete

iv

incorporating slag. The model simulations can be used to accurately predict the

experimentally measured data from different sources. Moreover, semi-

adiabatic temperature rises of concrete with fly ash and slag are evaluated

experimentally. The 28-day compressive strength model and initial slump

model are verified with the test results of concrete mixtures containing

different replacement levels of fly ash and slag.

Keywords: Adiabatic Temperature Rise, Hydration Degree, Slag, Thermal

Properties, Mass concrete, Model

v

Table of Contents

Chapter Title Page

Signature Page i

Acknowledgements ii

Abstract iii

Table of Contents v

List of Figures ix

List of Tables xiv

1 Introduction 1

1.1 General 1

1.2 Statement of Problems 2

1.3 Objectives and Scope of Study 4

2 Literature Reviews 8

2.1 Ground Granulated Blast Furnace Slag 8

2.2 Hydration Degrees of Cement and Slag 9

2.2.1 Hydration degree of cement 9

2.2.2 Hydration degree of slag 9

2.3 Experiments on Thermal Properties 13

2.3.1 Specific heat 13

2.3.2 Thermal conductivity 14

2.3.3 Coefficient of thermal expansion 16

2.4 Effect of Slag on Heat Evolution of Concrete 17

2.5 Thermal Cracking of Mass Concrete 19

2.5.1 Overview of thermal cracking 19

2.5.2 Models for predicting temperature and thermal cracking 20

3 Hydration Degrees of Cement and Slag 24

3.1 Determination of Hydration Degree of Cement 24

vi

3.2 Determination of Hydration Degree of Slag 28

3.3 Key Factors Affecting Hydration Degree of Slag in Concrete 29

3.4 Effect of Physical Acceleration of Cement by Slag Particles 32

4 Model for Predicting Free Water Content 34

4.1 General 34

4.2 Experimental Program 34

4.2.1 Materials 34

4.2.2 Mix proportions 35

4.2.3 Specimen preparation and test method 36

4.3 Experimental Results 36

4.3.1 Effect of water to binder ratio 38

4.3.2 Effect of slag replacement level 38

4.4 Model for Predicting Free Water Content 38

4.4.1 Free water content 38

4.4.2 Chemically bound water content 39

4.4.3 Gel water content 40

4.5 Verification of Free Water Model 40

5 Model for Predicting Specific Heat 43

5.1 General 43


5.2.1 Materials and mix proportions 43





5.4 Model for Predicting Specific Heat 48

5.5 Verification of Specific Heat Model 51

6 Model for Predicting Thermal Conductivity 53

6.1 General 53

vii







6.4 Model for Predicting Thermal Conductivity 57

6.5 Verification of Thermal Conductivity Model 58

7 Model for Predicting Coefficient of Thermal Expansion 60

7.1 General 60







7.4 Model for Predicting CTE 66

7.5 Verification of CTE Model 67

8 Model for Simulating Adiabatic temperature of Mass Concrete 69

8.1 General 69

8.2 Total Heat Generation of Concrete 69

8.3 Verifications using Proposed Adiabatic Temperature Model 73

9 Semi-Adiabatic Temperature Rise of Mass Concrete 81

9.1 General 81





viii

9.3.1 Effect of w/b 84

9.3.2 Effect of fly ash 84

9.3.3 Effect of slag 86

10 Initial slump 90

10.1 General 90

10.2 Experimental procedure 94


10.4 Verification of the initial slump model for concrete 97

11 Compressive Strength 99

11.1 General 99

11.2 Experimental procedure 100


11.3.1 Effect of w/b 101

11.3.2 Effect of fly ash 101

11.3.3 Effect of slag 104

11.4 Verification of the 28-day Compressive Strength Model 107

12 Conclusions and Recommendations for future studies 110

12.1 Conclusions 110

12.2 Recommendations for Future Study 113

References 114

Appendices 122

Appendix A 123

Appendix B 126

Appendix C 128

Appendix D 129

ix

List of Figures

Figure Page

1.1 Step-by-step procedure to compute adiabatic temperature rise in

concrete 5

1.2 Flowchart to compute semi adiabatic temperature rise and thermal

cracking potential of mass concrete containing slag 7

2.1 Key factors affecting the reactivity of slag in cement 11

2.2 Behavior of some factors towards the tendency of hydration degree of

slag in cement 11

3.1 Average hydration degree of cement pastes with w/b = 0.25 and 0.40 at

28°C and 40°C 27

3.2 Average hydration degree of cement pastes with T = 28°C and 40°C at

w/b = 0.40 27

3.3 Simulations of hydration degree of slag in concrete with w/b = 0.40 and

0.60 at slag replacement ratio (s) = 0.50 29

3.4 Simulations of hydration degree of slag in concrete with s = 0.50 and

0.75 at w/b = 0.40 30

3.5 Simulations of hydration degree of slag in concrete with T = 28°C and

40°C with s= 0.50 and w/b=0.40 31

3.6 Simulations of hydration degree of slag in concrete with Fslag = 4300

cm2/g and 6000 cm

2/g with s= 0.50 and w/b=0.40 32

4.1 Test results of weight ratio of free water to total binder of pastes with

slag replacement levels= 0%, 45%, 60%, and 75% at w/b=0.25 37

4.2 Test results of weight ratio of free water to total binder of pastes with

slag replacement levels= 0%, 45%, 60%, and 75% at w/b=0.40 37

4.3 Comparison between test results and model simulations of weight ratio

of free water to total binder of pastes with slag replacement levels= 0%,

45%, 60%, and 75% at w/b=0.25 41

4.4 Comparison between test results and model simulations of weight ratio

of free water to total binder of pastes with slag replacement levels= 0%,

45%, 60%, and 75% at w/b=0.40 42

x

Figure Page

5.1 Testing specific heat and thermal conductivity (a) setup of the sensor

sandwiched by a paste specimen, (b) Hot Disk Thermal Constants

Analyser apparatus 45

5.2 Test results of specific heat of pastes with slag replacement levels= 0%,

45%, 60%, and 75% at w/b=0.25 47

5.3 Test results of specific heat of pastes with slag replacement levels= 0%,

45%, 60%, and 75% at w/b=0.40 47

5.4 Comparison between test results and model simulations of specific heat

of pastes with slag replacement levels= 0%, 45%, 60%, and 75% at

w/b=0.25 51

5.5 Comparison between test results and model simulations of specific heat

of pastes with slag replacement levels= 0%, 45%, 60%, and 75% at

w/b=0.40 52

6.1 Test results of thermal conductivity of pastes with slag replacement

levels= 0%, 45%, 60%, and 75% at w/b=0.25 56

6.2 Test results of thermal conductivity of pastes with slag replacement

levels= 0%, 45%, 60%, and 75% at w/b=0.40 56

6.3 Comparison between test results and model simulations of thermal

conductivity of pastes with slag replacement levels= 0%, 45%, 60%,

and 75% at w/b=0.25 59

6.4 Comparison between test results and model simulations of thermal

conductivity of pastes with slag replacement levels= 0%, 45%, 60%,

and 75% at w/b=0.40 59

7.1 Temperature changing process to measure CTE of slag-cement pastes 61

7.2 Example of firmly wrapped paste specimens for measuring CTE 62

7.3 Experimental setup for measuring CTE of paste specimens 62

7.4 Test results of CTE of pastes with slag replacement levels= 0%, 45%,

60%, and 75% at w/b=0.25 65

7.5 Test results of CTE of pastes with slag replacement levels= 0%, 45%,

60%, and 75% at w/b=0.40 65

xi

Figure Page

7.6 Comparison between test results and model simulations CTE of pastes

with slag replacement levels= 0%, 45%, 60%, and 75% at w/b=0.25 67

7.7 Comparison between test results and model simulations CTE of pastes

with slag replacement levels= 0%, 45%, 60%, and 75% at w/b=0.40 68

8.1 Comparison of dispersion factors of fly ash and slag which affect the

adherence of cement particles in concrete 72

8.2 Comparison of tested and predicted adiabatic temperature rise for

concrete with 40% slag, w/b=0.392 and initial temperatures of 20°C

and 30°C 74


concrete with 40% slag, w/b=0.493 and initial temperatures of 20°C

and 30°C 74


cement concrete at w/b=0.4 and 0.65 76


concrete with 50% slag and at w/b=0.4 and 0.6 76






concrete with 65% slag and at w/b=0.50 79


concrete with 74% slag and at w/b=0.46 79

9.1 Setup of formwork for measuring semi adiabatic temperature rise of a

mass concrete sample 83

9.2 Tested results of semi-adiabatic temperature of concrete with fly ash

R=0%, 30%, and 50% at w/b=0.4 85


R=0%, 30%, and 50% at w/b=0.5 85

xii

Figure Page


R=0%, 30%, and 50% at w/b=0.6 86

9.5 Tested results of semi-adiabatic temperature of concrete with slag

S=0%, 50%, and 75% at w/b=0.4 87


S=0%, 50%, and 75% at w/b=0.5 87


S=0%, 50%, and 75% at w/b=0.6 88

9.8 Heating up and cooling down slopes computed from temperature

profiles of fly ash and slag concrete with w/b=0.50 89

10.1 Tested results of initial slump of concrete with fly ash R=0%, 30%, and

50% at w/b=0.4, 0.5, and 0.6 96

10.2 Tested results of initial slump of concrete with slag S=0%, 50%, 65%,

and 75% at w/b=0.4, 0.5, and 0.6 96

10.3 Comparison of tested and predicted initial slump values for concrete

containing fly ash R=0%, 30%, and 50% for all tested w/b 98

10.4 Comparison of tested and predicted initial slump values for concrete

containing slag S= 50%, 65%, and 75% for all tested w/b 98

11.1 Tested results of compressive strength of concrete with fly ash R= 0%,

30%, and 50% at w/b=0.40 102


30%, and 50% at w/b=0.50 103


30%, and 50% at w/b=0.60 104

11.4 Tested results of compressive strength of concrete with slag S= 0%,

50%, 65%, and 75% at w/b=0.40 105


50%, 65%, and 75% at w/b=0.50 106


50%, 65%, and 75% at w/b=0.60 106

xiii

Figure Page

11.7 Comparison of strength development ratios of concrete with fly ash R=

0%, 30%, 50%, and concrete with slag S=50%, 65%, 75% at w/b=0.50 107

11.8 Comparison of tested and predicted 28-day compressive strength values

for concrete containing fly ash R=0%, 30%, and 50% for all tested w/b 108

11.9 Comparison of tested and predicted 28-day compressive strength values

for concrete containing slag S=50%, 65%, and 75% for all tested w/b 109

xiv

List of Tables

Tables Page

4.1 Physical properties of cement and slag 34

4.2 Chemical compositions of cement and slag 35

4.3 Mix proportions of the tested paste specimens 35

5.1 Specific heat of the constituents of concrete 46

6.1 Thermal conductivities of the constituents of concrete 55

7.1 CTE values of the constituents of concrete 64

8.1 Maximum heat generation values for major cement compounds 73

9.1 Mix design of the binders used in casting concrete 82

1

Chapter 1

Introduction

1.1 General

Mass concrete can be defined as any volume of concrete in which the

dimensions are large enough for the requirement of considering minimization of

thermal cracking potential. Some examples of massive concrete structures are dams,

mat foundations, pile caps, transfer plates and structural members in which minimum

dimension is large, exceeding 0.9m for example. Hydration of cement generates heat

since it is an exothermic reaction. This generated heat dissipates rapidly in concrete

elements with relatively small dimensions. However, it is not the case for massive

concrete structures. The hydration process causes accumulation of heat inside the

mass concrete. Lower thermal conductivity of concrete makes it difficult to dissipate

the generated heat. Therefore, temperature difference occurs between center and

surface of the mass concrete element resulting in temperature gradients which lead to

thermal cracking especially at early age.

Thermal cracking results in many structural, durability, and aesthetic

problems. It can decrease the service life of concrete structures. Therefore, it is very

important to mitigate early age temperature and thermal cracking of mass concrete

especially at early age of the structure. Many approaches are taken to control

temperature rise of massive concrete during the construction stage. Some examples of

steps taken to alleviate thermal cracking problem are decreasing the placing

temperature of concrete, providing internal and external cooling, providing external

insulation, and using low heat materials, etc.

An effective way to minimize thermal cracking is to use Supplementary

Cementitious Materials (SCM’s) due to their lower reactivity at early stages. For

example, it is proven that fly ash can reduce the thermal cracking potential in mass

concrete when utilized as a partial replacement of cement. Ground Granulated Blast

Furnace Slag (herein after referred to as slag) is also widely used in some countries to

2

partially replace cement in mass concrete. This is a by-product from steel

manufacturing process which has gained attraction due to its engineering benefits

such as low heat generation, low permeability and improvement of some durability

properties. However, modelling of thermal properties and temperature are not

thoroughly investigated in slag replaced concrete. Therefore, understanding about the

behaviour of slag-cement system is essentially beneficial for the use of slag in

concrete construction industry.

To predict the thermal properties and temperature of slag concrete, evaluation

of hydration degree of cement and slag are critical steps. The interaction of slag-

cement system is also important to investigate. Then effect of slag towards the

quantitative change in free water amount and necessary thermal properties including

the heat evolution, heat transfer behaviour in concrete should be investigated. Thus,

thermal properties, temperature of slag concrete can be predicted which, not only can

be used to attain an insight for effectiveness of using slag in mass concrete

construction process, but also can be compared with other SCM’s currently utilized in

mass concrete construction.

1.2 Statement of Problems

The ability to accurately predict hydration degrees of cement and slag is the

major obstacle encountered in modelling many time-dependent properties of cement-

slag concrete. Many efforts have been carried out to find hydration degree of cement

in concrete (Tangtermsirikul & Saengsoy, 2002; Choktaweekarn & Tangtermsirikul,

2010; Kolani et al., 2012). However, the accuracy of the predictions were limited due

to several reasons such as not considering the effect of chemical and physical

properties of cement, errors in predicting stoichiometry of hydrated products, and not

considering all reactive compounds in cement. On the other hand, few studies

attempted to develop models to determine hydration degree of slag in paste or

concrete. Some of these models did not consider the influence of chemical

compositions, and physical properties of slag, time-dependent effect of free water and

specific heat, total heat generation in theoretical concepts of their modelling.

3

Therefore, accurate modelling of hydration degrees of cement and slag is essentially

beneficial for evaluating thermal cracking potential of slag concrete.

Predicting thermal properties such as specific heat, thermal conductivity,

coefficient of thermal expansion, heat transfer coefficient at exposed surfaces, and

mechanical properties such as early age tensile strength, tensile strain capacity and

modulus of elasticity are critical steps for modelling of thermal cracking potential in

concrete. It is a well-known fact that the hydration degrees of cement and slag

increase with time, which results in altering of the properties of concrete. Hence, the

thermal properties change with time especially at the early age of concrete. Therefore,

it is crucial to understand the effect of slag towards the behaviour of these properties

in order to model thermal properties of slag concrete.

Many researchers were able to experimentally investigate those thermal

properties of paste, mortar and concrete. However, most of the experiments were

conducted at 28 days or longer ages after casting (Bentz et al., 2011; Damdelen et al.,

2014; Demirboga, 2007). Some studies used oven-dried specimens thus omitting the

moisture condition inside the specimens, which supposed to be considered in accurate

evaluation to measure thermal properties. Few efforts have been made to find the

effect of fly ash towards altering time-dependent thermal properties in concrete

(Saengsoy, 2003; Choktaweekarn, 2008). In these studies, specific heat, thermal

conductivity and coefficient of thermal expansion were modelled based on the

thermal properties of each constituent and fraction of constituents in concrete at

considered age. However, the effect of slag was not modelled or experimentally

evaluated in these studies. The specific heat and thermal conductivity of mortar

containing different replacements of slag were experimentally measured by Viet et al.

(Viet et al., 2014). The measurements were obtained at 3 and 7 days and it was found

that the behaviour is similar to that of fly ash. However, modelling of thermal

properties for paste, mortar and concrete was not developed in his study.

Mostly, time-dependent modelling of thermal properties is limited due to

conflicts among results (Guo et al., 2001). Therefore, constant thermal properties are

4

mostly utilized for thermal cracking modelling. These assumptions lead to errors and

inaccurate predictions of temperature and thermal cracking potentials especially at the

early stage where thermal properties tend to change rapidly. Therefore accurate

modelling of thermal properties could be beneficial to provide solid background to

accurate thermal modelling of concrete.

In the case of mass concrete, adiabatic temperature rise is vital since it is the

most influential factor for thermal cracking potential. Several investigations have

been done to experimentally measure adiabatic temperature rise in mass concrete

incorporating slag (Attari et al., 2012; Tada et al., 2014). Models have been

developed for predicting temperature rise and evaluating thermal cracking potential of

mass concrete as well (Wang & Linger, 2010; Santhikumar et al., 1993; Faria et

al.2006). However, these models assumed constant specific heat values in their

analysis to predict temperature in concrete. Therefore the accuracy of these

simulations is limited.

A time-dependent model was developed (Choktaweekarn, 2008) to predict the

adiabatic temperature rise and thermal cracking potential of mass concrete containing

fly ash. It consists of various sub-models such as hydration degree of cement,

hydration degree of fly ash, thermal properties models and heat generation model.

Thereby, adiabatic temperature rise in the center of mass concrete can be computed

based on the criteria that no heat loss to or gain from surrounding. If any interaction

occurs between the concrete to or from surrounding, thermal conductivity involving

rate of heat flow in concrete, heat transfer coefficients and other external effects are

needed, hence semi adiabatic temperature rise can be predicted as well. Therefore,

modelling of thermal properties, temperature of mass concrete containing slag could

be beneficial for bridging a knowledge-gap for construction industry.

1.3 Objectives and Scope of Study

The key objective of this study is to develop models that can accurately

predict thermal properties and adiabatic temperature in concrete containing slag. For

5

this process, the initial step is to model the hydration degrees of cement and slag.

When slag is included as a binder, it not only reacts but also can affect the cement

hydration phases as well. Therefore, thorough investigation is carried out on the

interaction of slag-cement system and the effect of slag towards hydration process of

cement. Then, a time-dependent model for free water is developed based on

experiments for slag-cement paste. It is followed by modelling time-dependent

thermal properties which are specific heat, thermal conductivity, and coefficient of

thermal expansion based on experiments of slag-cement pastes. Then, these models

are linked to compute the adiabatic temperature rise at the center of mass concrete

containing slag. The developed model is to be verified for its accuracy using previous

experimental results on adiabatic temperature rise of mass concrete samples including

slag from different sources. The first part of this study can be illustrated through the

flowchart shown in Fig. 1.1.

Fig. 1.1 Step-by-step procedure to compute adiabatic temperature rise in concrete

Inputs: Initial temperature, Mix proportions, Necessary properties of cement, slag, aggregates and water

Total heat generation (Q)

Hydration degrees of cement and slag

Free water content

Specific heat (c)

Q=mcΔT

Adiabatic temperature rise (ΔT)

6

After efficaciously attaining the first objective, the next is to obtain semi-

adiabatic temperature rise and to evaluate thermal cracking potential in mass concrete.

Adiabatic conditions interpret no heat loss or gain from the surrounding environment.

However, when practical application is considered, semi-adiabatic condition is

observed due to some heat loss to or gain from environment. Therefore, this model is

to be extended in such a way that it could predict not only the semi-adiabatic

temperature rise in concrete, but also the risk of thermal cracking of the structural

member due to restraint of concrete (Choktaweekarn, 2008). In order to obtain semi-

adiabatic temperature and thermal cracking, the dimension of element, environmental

and boundary conditions must be input other than the parameters mentioned in the

original model. Heat of hydration of binder obtained from the previous adiabatic

temperature rise model, is used as an input for a commercial FEM program to

evaluate semi-adiabatic temperature rise. Then, thermal cracking potential is

evaluated based on the criteria comparing restraining strain to that of the tensile strain

capacity of any considered location in the mass concrete structure. The second part of

this study can be illustrated through the flowchart shown in Fig. 1.2. The prediction of

thermal cracking is not included in the scope of this study, thus, it can be investigated

as a further step.

Experimental process is carried out to investigate the semi-adiabatic

temperature rise of mass concrete specimens with different fly ash and slag

replacements. These results are to be verified by the model predictions. Initial slump

model is discussed and verified using the experimental results obtained from this

study. Moreover, compressive strength of slag concrete and fly ash concrete is

experimentally investigated as well. The 28-day compressive strength model is

described and verified using the test results in this study.

7

Fig. 1.2 Flowchart to compute semi-adiabatic temperature rise and thermal cracking

potential of mass concrete containing slag

Finite element heat transfer analysis

Thermal conductivity Total heat generation

Input: Convection heat transfer coefficient at the

surface of concrete

Input: Dimension of structure and ambient and boundary

conditions

Semi-adiabatic temperature

Temperature gradient between center and surface of concrete

Finite element structural analysis

Restraint due to different thermal deformation

Coefficient of thermal expansion

Input: Modulus

of elasticity, Poisson’s

ratio

Restraint due to differential thermal deformation

Check: Restrained strain > Tensile strain capacity

No Yes

No crack Crack

8

Chapter 2

Literature Reviews

2.1 Ground Granulated Blast Furnace Slag

Slag is a by-product obtained from steel and iron manufacturing process

(Beushausen et al., 2012). It is produced in the blast furnace as a by-product when the

iron ore is reduced by coke at 1350°C-1550°C. The main product which is molten

iron is formed in the bottom of the blast furnace while the liquid slag forms the layer

above it due to lower density of slag. Subsequently the liquid slag layer is separated

and it is cooled down using air or water flow. It is estimated that roughly 300 kg of

slag is produced per metric ton of pig iron (Chen, 2007).

There are three main types of slag which is categorized by the cooling process

(Chen, 2007). First is the air-cooled slag which is produced by solidifying the liquid

layer slowly in the air followed by accelerated cooling by spraying water. This is

normally used in asphalt paving, road bases, and as a concrete aggregate due to its

hard and dense structure. The second type is named as pelletized slag which is

produced by cooling slag with water and flinging it into air. This is mostly used in

cement production as a raw material. It has lower glass content than other types. The

third and mostly common form of slag is Ground Granulated Blast Furnace Slag. It is

produced by quenching the liquid slag with water in order to obtain granulates. These

granulates consist more than 95% of glass content. Grinding this material will result

in GGBS which is used as a mineral admixture in concrete (Oner & Akyuz, 2007).

The chemical composition of GGBS generally consists of 27-40% SiO2, 30%-

50% CaO, 5%-15% Al2O3 and 1%-10% MgO (Zhu et al., 2012). It is stated that the

performance of slag as a cementitious material depends highly on the chemical

composition, glass content and fineness of particles. Major positive effects of GGBS

could be mentioned as lower heat of hydration, increase durability and long term

strength and suitability in normal and severe conditions such as marine environment

9

as well. (Ground Granulated Blast Furnace Slag is referred to as slag here onwards, in

the forthcoming chapters).

2.2 Hydration Degrees of Cement and Slag

2.2.1 Hydration degree of cement

Hydration degree of cement has been modelled in many studies since it is the

governing process for prediction of various properties of concrete (Tantermsirikul &

Saengsoy, 2002; Choktaweekarn & Tangtermsirikul, 2010;Kolani et al., 2012).

Kolani et al. (2012) developed a hydration model for cement based on stoichiometric

calculations by relating the chemical composition of the cement to that of their

hydrated products. However, it was concluded that errors in the stoichiometry of

hydrated products reduced the accuracy of this model. Wang and Lee (2010)

developed a kinetic model for predicting hydration degree of cement based on

production and consumption of calcium hydroxide in concrete. However, the

accuracy was limited as this model did not consider the effect of hydration of each

reactive compound and influence of chemical and physical properties of binders.

Saengsoy and Tangtermsirikul (2002) developed a model for estimating hydration

degree of cement by considering reactivity of each reactive compound in cement

(C3A, C3S, C2S, and C4AF). An average hydration degree can also be estimated as a

weight average of the hydration degrees of all reactive compounds in the cement. In

the current study, this model is adopted to compute time-dependent hydration degree

of cement.

2.2.2 Hydration degree of slag

Hydration degree of slag in paste may be experimentally obtained by several

methods such as selective dissolution, recrystallization of slag from differential

scanning calorimetry, cumulative heat evolution curves from isothermal calorimetry,

and chemical shrinkage curves (Kochaba et al., 2011). Among these tests, selective

dissolution method is a comparatively simple, effective and accurate method to obtain

hydration degree of slag. It is based on a preferential chemical dissolution of all other

10

products except unreacted slag (Kochaba et al., 2011; Lumley et al., 1996; Hanehara

et al., 2014; Escalante et al., 2001). The hydration degree can then be calculated

based on the ratio between the reacted amounts of slag per total amount of slag in the

paste mixture. Lumley et al. (1996) observed that hydration degree of slag

significantly changes with age where at water to binder ratios (w/b) of 0.4-0.6 and

20°C, 30-55% of slag reacts in 28 days and 45-75% in 1-2 years. Escalante et al.

(2001) investigated hydration degree of slag by selective dissolution method under

different conditions. It was found that hydration degree of slag increases with

temperature and w/b ratio, whereas, it decreases with increasing replacement level.

Previous studies carried out in Japan (Hanehara et al., 2014; Sagawa & Nawa, 2014)

observed that hydration degree increases as the fineness of slag increases indicating

the importance of physical properties of the slag particles towards hydration.

Analysing experimental results from selective dissolution method from different

sources could be beneficial to develop a generalized model to quantitatively

determine hydration degree of slag.

The behaviour of slag when mixed with water is somewhat different to that of

other minerals admixtures such as fly ash. Slag has a slight self-cementing ability

which does not require calcium hydroxide to form cementitious products (Wang &

Linger, 2010). Although slag show self-hydrating characteristics, the amount of

cementitious products formed and the rates of formation are insufficient for use in

structural applications. Slag reacts remarkably when the environment is suitable, an

alkaline environment in case of concrete (Chen, 2007). The reaction of slag with

hydroxides could be explained by two stages. Initially, slag reacts with alkali metal

hydroxides which are dissolved immediately when cement is mixed with water. In the

second stage, once the concentration of calcium hydroxide reaches a certain amount,

the reaction between slag and CH becomes dominant (Merzouki et al., 2013).

Many factors affect the rate and the reactivity of slag. The key factors

affecting reactivity of slag in cement summarized by Chen (2007) is illustrated in Fig.

2.1. These factors result in either increase or decrease of the hydration degree of slag.

Higher hydration temperature, water to binder ratio, hydraulic index, vitreous fraction

11

and specific surface area increase the reactivity of slag (Escalante et al., 2001).

However, reactivity of slag decreases with the increase of percentage of slag

replacement. The representations of behaviour of few such factors were illustrated

previously by Escalante (2001) as indicated in Fig. 2.2.

Fig. 2.1 Key factors affecting the reactivity of slag in cement (Chen, 2007)

Fig. 2.2 Behaviour of some factors towards the tendency of hydration degree of slag

in cement (Escalante, 2001)

12

Several models have been developed to determine hydration degree of slag

(Kolani et al., 2012; Wang & Linger, 2010; Merzouki et al., 2013; Luan et al., 2012).

Kolani et al. (2012) developed a model to quantify the hydration degree of slag in

blended cement through the measurements of portlandite content, free water content

and hydration heat. Wang and Linger (2010) introduced a kinetic model to predict

hydration degree of slag by considering the portlandite production by cement

hydration and its consumption by slag. These models did not consider the influence

of chemical compositions, and physical properties of cement and slag. Merzouki et

al. (2013) proposed another kinetic model based on stoichiometry of the reaction

between slag and portlandite in which chemical composition of cement and slag,

fineness, curing temperature and w/b were taken into account. However, the model

did not consider the time-dependent effect of free water and specific heat and also

underestimated the portandite for high replacements of slag. Luan et al. (2012)

proposed a model to predict hydration degree considering the role of portlandite as the

activator and the Ca/Si ratio of C-S-H. Time-dependent variation in total heat

generation and specific heat were not considered in this model. Therefore,

generalized model for quantification of reaction degree of slag by considering all

main factors and time-dependent variables is critical to determine the behaviour of

slag in concrete.

13

2.3 Experiments on Thermal Properties

2.3.1 Specific heat

Specific heat can be defined as the amount of heat required to change the

temperature of a unit mass of a certain substance by a unit degree of temperature

change. This parameter is used to relate heat of hydration generated in concrete and

its temperature rise is computed by Eq. (2.1) where Q is the cumulated heat of

hydration (kcal), m is mass of concrete (kg), c is the specific heat of concrete

(kcal/kg/oC) and ∆T is temperature rise at considered time t (

oC).

TmcQ (2.1)

The specific heat of concrete is a time-dependent parameter. As the hydration

proceeds, amount of free water reduces rapidly with time hence increasing the degree

of hydration. Since the specific heat of water is the highest among all other

constituents in concrete, specific heat decreases rapidly with time at the early age

(Choktaweekarn, 2008).

Few experimental investigations have been done to observe the tendency of

specific heat of slag-cement paste. A previous study (Schutter & Taewe, 1995)

measured specific heat of slag-cement paste by obtaining the temperature rise in a

calorimeter when given certain energy supply. Tests were done for paste specimens

with w/b of 0.5 and it was concluded that specific heat decreases linearly with the

increase of hydration degree. Another study (Viet T. , 2013) measured specific heat of

slag-cement paste at w/b of 0.4. Two slag replacements of 25% and 60% were utilized

and it was found that, specific heat increases as the slag level increases. This is due to

higher free water amount available in the initial stage of slag-cement paste than that of

the cement only paste. However, this study only concerned results at 3 and 7 days.

Bentz (2007) and He (2005) used transient plane source technique to measure

specific heat of cement paste. This method is based on a use of a transiently heated

plane sensor and it is typically stated as Hot Disk Thermal Constants Analyzer. The

14

hot disk technique is based on using a thin metal strip or a metal disk as a continuous

plane heat source (He, 2005; Gustafsson, 1991; Log & Gustafsson, 1995). The metal

disk or strip is sealed between two thin polyimide films for electrical insulation.

During the experiment, the hot disk sensor is sandwiched between two pieces of

samples to be investigated, and a small constant current is supplied to the sensor. The

sensor also serves as a temperature monitor, so that the temperature increase in the

sensor is accurately determined through resistance measurement as a function of time.

For the samples, w/b ratios of 0.3 and 0.4 were used and it was concluded that

specific heat increases as w/b of the sample increases (Bentz, 2007). Moreover, it

decreases as hydration degree increases.

2.3.2 Thermal conductivity

Thermal conductivity is known as the rate of heat transfer through a unit cross

sectional area of material for a specific temperature gradient. Heat transfer through

conduction in a specified direction is proportional to temperature gradient in that

direction and the area perpendicular to the direction of heat flow. The general

equation of rate of heat flow in specified direction x, qx (kcal/day) is shown in Eq.

(2.2) as follows.

dx

dTkAqx (2.2)

where k is the thermal conductivity (kcal/ m day °C), A is the area of the cross

section perpendicular to the direction of heat flow (m2), dT/dx is the temperature

gradient in the direction of heat flow (°C/m).

Thermal conductivities of aggregates are higher than that of the paste, thus the

mineralogical character of aggregate greatly influences thermal conductivity of

concrete (Choktaweekarn, 2008). Moreover, degree of crystallization of aggregate has

significant effect towards the thermal conductivity (Demirboga, 2007). Heat

conduction is higher in aggregate with crystalline structure than amorphous and

vitreous aggregate. Porosity and moisture content are other factors that affect this

15

property in concrete. Moreover, as the hydration process continues, amount of free

water in concrete reduces with an increase of hydrated products resulting in increasing

thermal conductivity with time especially in early stages.

There are two main techniques for measuring thermal conductivity which are

steady state and transient methods. Steady state method is useful when the material

under examination is rigid and dry or conditioned to the ambient condition

(Choktaweekarn, 2008). The method is not suitable when moisture redistribution can

occur during the period of the test. The transient method is convenient to use with

rigid and semi-rigid materials and has particular advantages when thermal

conductivity of moist materials is to be measured. Transient measurement technique is

appropriate for low conductivity porous materials. The rapidity of the determination

does not allow sufficient time for any moisture movement to occur within the sample

during testing. Moisture has great effect on thermal conductivity of concrete, then

transient method is preferable.

Steady state method was applied in a previous study to find the effect of slag

on thermal conductivity by using 15% and 30% slag replacements by weight

(Demirboga, 2007). Paste and mortar specimens were tested at 28 days in oven dried

conditions. The results showed that thermal conductivity decreased with the increase

of slag replacements levels. Demirboga further analyzed the effect on thermal

conductivity with high replacement levels of slag (Demirboga et al., 2007). For 50%,

60% and 70% replacement levels of slag, thermal conductivity decreased 15%, 18%

and 17% respectively. It is stated that the density reduces as the slag replacement is

higher, thus affecting lower thermal conductivities of concrete.

Transient method was used by Bentz (2007) to measure thermal conductivity

of cement pastes since casting until 28 days. Hot disk thermal constants analyzer was

used for the measurements. The variables investigated included w/b and different

curing conditions of saturated and sealed. It was observed that hydration did not

significantly affect the thermal conductivity of pastes. The similar method was used in

another study (Choktaweekarn, 2008) to measure thermal conductivity of paste and

16

mortar with partial fly ash replacements. It was concluded that the values increase just

after casting up to 3 days of age, then followed by constant tendencies. Moreover, it

was found that the use of fly ash reduces thermal conductivity due to decreased

density of specimens.

2.3.3 Coefficient of thermal expansion

A study stated that concrete can slightly expand or contract depending on the

temperature rise or fall (Neville & Brooks, 1987). The coefficient of thermal

expansion (CTE) is known as the length change of a unit length per unit degree of

temperature change. The expression for CTE for concrete is mentioned from Eq.

(2.3).

T

LLCTEc

)/( 0 (2.3)

Where CTE is the coefficient of thermal expansion of concrete (micron/oC),

∆L is the length change of the specimen (mm), Lo is the initial length of the specimen

(mm) and ∆T is the temperature change (oC).

CTE of concrete is a result of CTE of cement paste and aggregate since they

are the main constituents. If the CTE of aggregate and cement paste differs too much,

a large temperature change which cause differential movement may break the bond

between them easily (Choktaweekarn, 2008). Therefore utilizing a proper aggregate

which has a similar CTE value to the paste could reduce the internal stresses.

However, due to the availability and cost of aggregate, use of mineral admixtures is

considered as a more feasible way of reducing CTE in concrete (Shui et al., 2010).

Shui et al. (2010) investigated CTE of hardened cement paste with slag

replacements of 15% and 30%. The thermal dilation rate was measured with changing

temperature from 20oC to 85

oC at a rate of 0.5

oC per minute. For every 5

oC

temperature was kept constant for two minutes to allow sufficient temperature

equilibrium of specimens. Then length change of specimen was recorded. Afterwards

17

using the following Equation, thermal dilation rates were converted into CTE.

Important conclusions were obtained by measuring CTE together with porosity and

CH content of cement paste. It was observed that slag increased porosity of the

cement paste and decreased CH content leading to lowering CTE values. Voids could

accommodate internal thermal expansion of material, thus increased porosity resulted

in decreasing CTE.

Kada et al. (2002) developed a simple method to find the CTE at early age of

concrete. This method is based on applying a temperature shock in a range of 10°C to

50°C in a short period not longer than one hour. Since the duration was short, the

effect of autogeneous shrinkage was not considered in this experiment. It was

observed that the CTE of concrete with w/b of 0.45 decreased rapidly within few

hours after casting, and then remained almost constant. The similar method was

adopted in a previous study (Choktaweekarn, 2008) for testing paste, mortar and

concrete by cooling them down in a refrigerator to reduce temperature until 10°C,

then followed by moving out to return the temperature back to room temperature

(about 30±2°C). The specimens were tested for the change in length for every 5°C

change of temperature, thereby CTE values were computed. It was concluded that

CTE of concrete is a time-dependent property which increases with age due to

increased continuity of the structure.

2.4 Effect of Slag on Heat Evolution of Concrete

A laboratory program (Attari et al., 2012) measured the heat generated from

concrete mixes including slag replacements from 0% to 70%. Thermocouples were

used to record internal temperature of concrete specimens with dimensions of

30×30×15 cm. The specimens were insulated during the curing process to prevent

excessive heat loss to or gain from surrounding. Then the total amount of heat

released was computed based on the laws of heat transfer and measured temperature

profiles until one week. It was concluded that the total heat released during the

measurement period reduced as the slag replacement level increased, however,

18

continuous increase was observed at the later stage when compared to that of the

cement only concrete.

Hydration temperatures of high-strength concrete columns were

experimentally investigated to assess the influence of slag replacement level ranging

from 0% to 70% (Sioulas & Sanjayan, 2000). The large concrete columns of

dimensions 80×80×120 cm were cast with top and bottom surfaces insulated using

50mm thick expanded polystyrene form. The results indicated that the peak

temperature rise encountered at the center of the columns significantly reduced with

the inclusion of slag into concrete. A delay in the time to attain the maximum

temperature is observed for the specimens with slag and it increased as the slag

replacement increased. The thermal gradients observed were 128, 109, 86, and 64

°C/m for the specimens with slag replacements of 0%, 30%, 50%, and 70%,

respectively. Moreover, the temperature difference between center and surface was

mitigated by including slag in concrete.

Semi-adiabatic calorimetry technique was used previously (Ruiz et al., 2001)

to investigate the heat evolution of concrete containing 30% and 50% of slag. It was

reported that the rate of temperature rise was lower when the slag replacement is

higher. However, the time to attain peak temperature was shortened for the 30% slag

specimen when compared to that of the 50% slag replaced specimen. Moreover, it

was concluded that lower water to binder ratio may result in similar or slightly higher

heat evolution of concrete containing lower replacements of slag.

A previous study analyzed the heat evolution of high performance concrete

using normal and ultra-fine slags (Divsholi & Lim, 2012). Two types of slag are used

which have different particle sizes of 4100 cm2/g and 8700 cm

2/g, respectively. The

slag replacements used in this study were 45% and 60%. It was observed that the

temperature rise was higher in concrete containing ultra-fine slag when compared to

that of the concrete containing normal slag. This was due to increase of total surface

area for reaction which increases the rate of hydration and pozzolanic reactions.

19

An experimental study (Tada et al., 2014) comparatively investigated the

effect of fly ash and slag on heat evolution of concrete. Adiabatic temperature rise

was tested using air circular type measurement equipment containing a sample

capacity of 50 liters. Practical fly ash and slag replacements levels which are mostly

used in mass concrete were used. Fly ash was varied from 20% to 50% and slag was

varied from 50% to 75%. It was reported that peak temperature significantly reduces

in fly ash concrete when compared to that of slag concrete. Moreover, lower thermal

gradients were observed in fly ash concrete as well.

2.5 Thermal Cracking of Mass Concrete

2.5.1 Overview of thermal cracking

Hydration of cement generates heat since it is an exothermic reaction. This

generated heat dissipates rapidly in concrete elements with relatively small dimension

(Alhozaimy et al., 2015). But, for massive concrete structures, the hydration process

causes accumulation of heat inside the structures. The difference of temperature

between center and surface of the mass concrete results in temperature gradient which

leads to thermal cracking especially at early age (Ballim, 2003; Nili & Salehi, 2010;

Saengsoy & Tangtermsirikul, 2003). Thermal cracking results in deterioration and

reduces service life of concrete structures.

There are several reasons for the propagation of thermal cracking. One of the

main reasons is utilizing concrete mixes that generate significant amount of heat such

as high-strength and self-consolidating concrete mixes (Ng, 2014). Construction of

massive concrete structures without appropriate measures to mitigate heat generation

is another reason. Disregarding the consequences of thermal movements during both

design and construction stage may lead to occurrence of thermal cracking as well.

It is required to control temperature especially in massive concrete structures,

hence many methods are practically utilized in order to alleviate thermal cracking

problem. There are three main methods used to mitigate thermal cracking in the

design and construction phases of massive concrete structures (Ng, 2014). The first

20

method is restraint analysis and optimization. It is done by reducing adverse effects of

movement restraint. For example, using different construction sequence or cycles

such as layer casting method can be utilized to reduce thermal cracking

(Choktaweekarn & Tangtermsirikul, 2010). The second method is taking thermal

control measures. Using in-situ cooling pipes by circulating air or water in the curing

process, reducing placing temperature, and external insulation are some of examples

for thermal control measures. The third method is to use appropriate mix design such

that the thermal cracking potential is lowered. Replacing part of mixing water with

ice, use of aggragates with lower CTE, and use of supplementary cementitious

materials (SCM’s) as partial replacement of cement are some of the examples under

this method.

Use of SCM’s is the easiest and economically and environmentally viable

method among all these methods to reduce thermal cracking of massive concrete

structures. However, rheological, thermal and mechanical properties may change

when cement is partially replaced by mineral admixtures. Therefore, it is essentially

beneficial to develop modeling for the concrete to predict the most appropriate mix

design of concrete with mineral admixtures, in order to minimize thermal cracking

potential of concrete.

2.5.2 Models for predicting temperature and thermal cracking

A numerical model was previously developed (Attari et al., 2012) in order to

simulate adiabatic temperature rise of concrete containing slag. Firstly, hydration

degree of concrete was quantified based on the heat released. Then heat of hydration

was experimentally verified from testing semi- adiabatic temperature rise of concrete

containing slag replacements from 0% to 70%. Best fit mathematical model was

proposed to compute adiabatic temperature rise based on laws of heat transfer.

However, it was reported that the prediction of hydration degree is not accurate since

it was computed based on the heat released. Moreover, the effect of slag on free

water, specific heat was not considered as well.

21

Ballim (2003) proposed a model for predicting time based temperature profiles

in mass concrete elements. The rate of heat generation was determined using an

adiabatic calorimeter in which a concrete sample was placed and temperature rise was

monitored up to 7 days. The results were then converted into maturity rate of heat

evolution. A finite difference model was used to compute time-dependent temperature

at any location of the concrete by using the results obtained from heat rate-maturity

relationship. Laboratory verifications were carried out in order to validate the model.

It was reported that the accuracy of the model predictions was limited due to errors in

assumptions of constant thermal properties.

Previous study (Luan et al., 2012) composed a model to predict adiabatic

temperature rise of concrete containing slag. The model was developed based on

hydration degrees of cement and slag. Heat generation and specific heat of concrete

were assumed as constants. Adiabatic temperature rise in concrete at different initial

temperatures and mix proportions were simulated. The accuracy of the predictions

was limited especially at early age due to assumptions of constant thermal properties

and heat rates.

Wang and Linger (2010) developed a kinetic model for predicting adiabatic

temperature rise of concrete containing fly ash and slag. The hydration degrees of

cement and mineral admixtures were computed based on production and consumption

of calcium hydroxide in concrete. Adiabatic temperature rise predictions were done

and it was reported that limitations exist as this model did not consider the effect of

hydration of each reactive compound and influence of chemical and physical

properties of binders.

A multi-component model for hydration heat of blended cement was proposed

in a previous study (Kishi & Maekawa, 1996). Reactive components considered were

Portland cement, fly ash, and slag. The heat rate and the thermal activity of each

mineral compound were computed based on material properties. The hydration

degrees of cement and mineral admixtures were computed step by step by using

modified Arrhenius’s low of chemical reactions. The retardation of fly ash on cement

22

hydration was considered when computing the hydration degrees. The model was

verified by analysis of adiabatic and semi-adiabatic temperature rise simulations.

However, the hydration of cement was affected by the addition of slag as well, which

was not considered in this model. Moreover, the effect of slag for thermal properties

was not considered as well.

A model to predict thermal cracking was previously developed (Schutter,

2002). It was based on hydration degree of blended cement containing slag. The

cracking tendency was implemented by a smeared cracking approach having non-

linear softening behavior. In this study, it was assumed that the specific heat and

thermal diffusivity of hardening concrete decrease linearly with increasing hydration

degree. However, constant value was used for the coefficient of thermal expansion

due to lack of experimental results. The proposed model was verified with

experimental results and showed good agreement with the tested tendencies of

thermal cracking.

A computerized program was developed for simulating temperature in mass

concrete in a previous study (Saengsoy, 2003). It is capable of simulating the

temperature in fly ash concrete. The required input parameters for this program are

initial temperature, mix proportions of concrete and properties of cement and fly ash.

When these parameters are given, the program initially computes the degree of

hydration in cement compounds and degree of pozzolanic reaction of fly ash in

concrete. It was followed by computing the cumulative heat based on summation of

all the heat liberated from the reactions of each cement compound including

formation of ettringite and monosulphate, and the reaction of fly ash. A time-

dependent model for specific heat was introduced, enabling more realistic temperature

simulation especially at early age of mass concrete. It was reported that the

formulated model could accurately predict the test results of adiabatic temperature

rise.

23

When practical applications are considered in real mass concrete construction,

semi-adiabatic condition is observed due to heat loss to environment. Therefore, an

enhanced model to compute not only semi-adiabatic temperature rise in concrete, but

also the risk of thermal cracking of structural members due to restraint of concrete is

developed in a previous study (Choktaweekarn, 2008). In order to obtain semi-

adiabatic temperature and thermal cracking potential, the dimension of element,

environmental and boundary conditions must be input other than the parameters

mentioned in the original model by Saengsoy (2003). Heat of hydration and

pozzolanic reactions were obtained by from the previous adiabatic rise model and

they were used as inputs for a commercial FEM program. Moreover, models were

proposed to obtain thermal properties such as specific heat, thermal conductivity and

coefficient of thermal expansion. FEM program was used to analyze semi-adiabatic

temperature and restrained strain of concrete. In this study, restrained strain was

compared to that of the tensile strain capacity of specific point of concrete to evaluate

the thermal cracking potential of fly ash concrete. If the restrained strain is higher

than that of the tensile strain capacity of concrete at any point, it results in thermal

cracking.

24

Chapter 3

Hydration Degrees of Cement and Slag

3.1 Determination of Hydration Degree of Cement

The mechanism of cement hydration process is complex, thus reaction

behavior has been studied considering each individual reactive compound (Young et

al., 1998; Saengsoy, 2003). The basic reactions of the cement hydration process is

described below. The abbreviated notations are used for the principal oxides in clinker

as follows; C for CaO, S for SiO2, A for Al2O3, F for Fe2O3, S for SO3, and H for

H2O.

Eq. (3.1) and Eq. (3.2) indicate examples of hydration reactions of Calcium

Silicates which are C3S and C2S. These reactions produce calcium silicate hydrate

gels and calcium hydroxide.

CH3 HSC H6 SC2 3233 (3.1)

tricalcium silicate Water C-S-H Gel Calcium hydroxide

CH HSC H4 SC2 3232 (3.2)

dicalcium silicate Water C-S-H Gel Calcium hydroxide

As seen from Eq. (3.3), C3A reacts with calcium and sulfate ions that are

available due to dissolution of gypsum.

323623 HASC H26 CSH3 AC (3.3)

tricalcium aluminate gypsum water ettringite

Ettringite is stable only if sulphate is available. If all sulphate was consumed

before C3A hydration complete, ettringite transform into monosulfoaluminate as

indicated in Eq. (3.4).

25

12432363 ASHC3 H4 HASC AC2 (3.4)

tricalcium aluminate ettringite water monosulfoaluminate

If monosulfoaluminate is again contacted with a new source of sulphate, it

reacts and forms ettringite as indicated in Eq. (3.5).

32362124 HASC H16 CSH2 ASHC

(3.5)

Monosulfoaluminate gypsum water ettringite

If there are no sulphate ions present, C3A will form calclium aluminate

hydrates as indicated in Eq. (3.6).

821343 AHC AHC H21 AC (3.6)

C4AF reactions are slower than C3A reactions. Due to insufficient calcium to

form calcium sulfoaluminates, amorphous hydrous oxides of aluminium or iron will

form as indicated in Eq. (3.7) and Eq. (3.8).

3323624 H)F,A( HS)F,A(C H21 CSH3 AFC (3.7)

312432364 H)F,A(2 SH)F,A(C3 H7 HS)F,A(C AFC2 (3.8)

In this study, the equations for predicting the hydration degrees of C3A, C3S,

C2S, and C4AF are adopted from a previous study carried out by Saengsoy and

Tangtermsirikul (2003). The hydration degree which represents reaction rate was

modeled previously as a function of age, water to cement ratio and concrete

temperature. This proposed model computed the hydration degree at each 10°C

increment of concrete temperature. Therefore, if the temperature is in between a

multiplier of 10°C, linear interpolation was done to obtain the hydration rates. The

original model was modified by Choktaweekarn (2008) to increase the accuracy at

26

early age of concrete. The modified model is used in this study to simulate the

hydration degree of cement.

Since, the model predict hydration degree of each reactive compound, the

average hydration degree is computed based on the assumption that each cement

compound react independently. The average hydration degree of cement in concrete is

defined as the weight fraction average of hydration degree of all cement compounds

in the concrete mix. It is computed by using Eq. (3.9).

4

1i

i

4

1i

ii

hy

m

)t(m

)t( (3.9)

Where αhy(t) is the average hydration degree of cement (%), i denotes each

mineral compound of cement (C3A, C3S, C2S, and C4AF), mi denotes the mass of

each compound per unit cubic meter of cement (kg/m3), and αi(t) denotes respective

hydration degree of reactive cement compound i (%).

Some examples of the average hydration degree of cement in pastes which

illustrate the effect of w/c ratio and temperature are shown in Fig. 3.1 and Fig. 3.2,

respectively.

27

Fig. 3.1 Average hydration degree of cement pastes with w/b = 0.25 and 0.40 at 28°C

Fig. 3.2 Average hydration degree of cement pastes with T = 28°C and 40°C at w/b =

0.40

28

3.2 Determination of Hydration Degree of Slag

The hydration degree of slag in this study is defined as the weight fraction of

already reacted slag per total slag in the concrete mix. As mentioned in the literature

review, this can be obtained experimentally for paste by selective dissolution method.

In this study, the main objective is to develop a time-dependent model for

predicting hydration degree of slag in concrete. It is difficult to develop a generalized

model to compute hydration degree of slag based on previous experimental results

obtained from different sources due to dissimilar values, possibly caused by different

methods of measurements. Moreover, analysis of selective dissolution data done only

in paste samples is considered not possible to include the effect of fine and coarse

aggregate in concrete. It is known that the presence of aggregate enhances the degree

of reaction by providing a better mixing efficiency (Chang & Peng, 2001). Therefore,

the model equation is obtained using the method of back analysis from adiabatic

temperature results of slag concrete from Tada et al. (2014) since the slag and cement

used in this previous study and our study are obtained from the same source of

manufacturer.

However, the tendencies of previously obtained results from selective

dissolution method by Lumley et al., 1996; Copeland & Kantro, 1969; Shinwa et al.,

2009; Escalante et al., 2001; Hinrichs & Odler, 1989; Sagawa & Nawa, 2009 are also

studied to investigate the key parameters affecting the hydration degree of slag. The

key parameters of the model for predicting hydration degree of slag are age, water to

binder ratio, concrete temperature, slag replacement ratio and fineness of the slag.

The equation for computing hydration degree of slag is shown in Eq. (3.10).

)b/wj1(

slag

slag

cb

slag

k

Ti)3500F(h1

T)sgb/wf(*Expe1)]1t(d[Expsta

(3.10)

where αslag is the hydration degree of slag in concrete at the considered age. t,

s, w/b, T, and Fslag are the considered age (days), slag replacement ratio, water to

29

binder ratio, initial temperature of concrete (°C) and Blaine’s fineness (cm2/g) of the

slag, respectively. Coefficients a, b, c, d, e, f, g, h, i, j, k and h are constants obtained

from regression analysis (a=11.35, b=0.1, c=0.45, d=-0.088, e=0.018, f=1.99, g=-2.3,

h=0.00017, i=0.043, j=3.18, k=2.1). The model is capable of predicting the hydration

degree within the limits of slag replacement from 0.40 to 0.75, water to binder ratio

from 0.35 to 0.65, and slag fineness from 3500 cm2/g to 6000 cm

2/g. The effect of

chemical compositions of slag is not included in the slag hydration degree model.

3.3 Key Factors Affecting Hydration Degree of Slag in Concrete

The hydration degree increases as the w/b of the paste increases. Addition of

water results in higher contact between slag and water which leads to form increased

amount of hydrated products. This results in enhancing reactivity of slag (Lumley et

al., 1996; Copeland & Kantro, 1969). The tendency of model simulations follows the

similar pattern as shown in Fig. 3.3 which illustrates the hydration degree of slag in

concrete at w/b ratios of 0.4 and 0.6.

Fig. 3.3 Simulations of hydration degree of slag in concrete with w/b = 0.40 and 0.60

at slag replacement ratio (s) = 0.50

30

Decrease in slag replacement percentage results in increasing the hydration

degree of slag in cement paste. This occurs due to higher alkaline activating effect of

the cement at lower concentrations of slag as previously reported by Escalante et al.

(2001) and Hinrichs & Odler (1989). The effect of slag replacement level for

hydration degree of slag in current model is shown in Fig. 3.4 which illustrates the

hydration degree of slag in concrete at slag replacements of 0.50 and 0.75.

Fig. 3.4 Simulations of hydration degree of slag in concrete with s = 0.50 and 0.75 at

w/b = 0.40

Effect of temperature on hydration degree of slag is investigated and described

previously by Shinwa et al. (2009), Escalante et al. (2001) and Luke & Glasser,

1988). It is reported that the hydration degree of slag is more sensitive to the curing

temperature than that of the hydration degree of cement. It is well known that as the

temperature increases, the molecules move faster and collide more vigorously

resulting in higher enhancing the hydration degree of slag. Fig. 3.5 illustrates the

hydration degree of slag in concrete at 28°C and 40°C in which the hydration degree

increases as the temperature increases.

31

Fig. 3.5 Simulations of hydration degree of slag in concrete with T = 28°C and 40°C

with s= 0.50 and w/b=0.40

Fineness of the slag particles is another vital factor affecting the hydration

degree of slag. Hydration degree of slag is proportional to the total surface area

(Chen, 2007). Thus, it is reported that as the fineness increases, the hydration degree

of slag increases due to higher contact area for the reaction to progress (Hinrichs &

Odler, 1989; Sagawa & Nawa, 2009). The effect of slag fineness from the model

simulations in this study is illustrated in Fig. 3.6.

It is clearly seen that the model simulation patterns follow similar theoritical

background as described in previous studies. However, the comparison between

magnitudes of results of hydration degree from selective dissolution method

conducted on pastes containing slag and hydration degree of slag computed in this

study may not be reasonable at this stage. This is because the model in this study

concerns about hydration degree of slag in concrete whereas the experimental results

by selective dissolution methods concerns about pastes.

32

Fig. 3.6 Simulations of hydration degree of slag in concrete with Fslag = 4300 cm2/g

and 6000 cm2/g with s= 0.50 and w/b=0.40

3.4 Effect of Physical Acceleration of Cement by Slag Particles

The addition of slag can affect the cement hydration process. The reduction of

cement content decreases the total heat generation in the presence of mineral

admixtures (Alhozaimy et al., 2015; Nili & Salehi, 2010). However, it does not

proportionally reduce the initial rate of heat evolution. This is due to the ability of the

secondary fine particles to disperse cement particles denoted as a kind of physical

acceleration (Saengsoy et al., 2003; Choktaweekarn & Tangtermsirikul, 2011).

This physical acceleration effect was considered as an important factor in

previous studies for evaluating the effect of limestone powder and fly ash on cement

hydration (Choktaweekarn & Tangtermsirikul, 2011; Poppe & Schutter, 2005).

In the case of slag, Ogawa et al. (1980) indicated that slag can accelerate the

hydration of C3S based on analysis from C3S-slag system using synthesized C3S from

reagents. Hoshino et al. (2006) reported that slag accelerates the early age hydration

33

of C3S and C3A from analysis of hydration degree of cement paste using combined X-

ray diffraction and Rietveld analysis method.

In this study, dispersion effect is considered to affect hydration degrees of C3S

and C3A in cement at very early age. However, its effect on hydration reactions at

later age is insignificant. Equations for dispersion effect are proposed for acceleration

of C3S and C3A hydration, by adopting the concept of a previous study by

Choktaweekarn (2011), as shown in Eq. (3.11) and Eq. (3.12), respectively. Hence,

the existing hydration degrees of C3S and C3A are modified using Eq. (3.13) and Eq.

(3.14).

rslagcslag

SC

SCSCSC

SC

wwFt

ttt

1

234

9420100

120450

100

148642

100

189913

100

166121

3

333

3

tan.)(

.

)(.

)(.

)(.

(3.11)

r

slagcslag

1AC

2

AC

3

AC

4

AC

AC

wwFtan78.0100

)1t(1354.0

100

)1t(4181.0

100

)1t(5095.1

100

)1t(0281.1

3

333

3

(3.12)

SCSCifiedmod,SC 333 (3.13)

ACACifiedmod,AC 333 (3.14)

Where ΦC3S and ΦC3A are the dispersion factors for C3S and C3A acceleration,

respectively. Fslag is Blaine’s fineness of slag powder (cm2/g). wc and wslag are weight

ratios of cement and slag per unit weight of concrete. r is the replacement ratio of

slag. αC3S and αC3A are hydration degrees of C3S and C3A, respectively. αC3S,modified,

αC3A,modified are the modified hydration degrees of C3S and C3A after the physical

acceleration, respectively and t is the considered age in days.

34

Chapter 4

Model for Predicting Free Water Content

4.1 General

Water inside hardened paste system can be classified into two main parts

which are evaporable and non-evaporable water. Free water is defined as evaporable

water which is freely accessible for hydration process (Saengsoy & Tangtermsirikul,

2003). Many studies have reported that change of free water in concrete is significant

especially at early age, and it affects the thermal properties as well (RILEM

Commiision, 1981; Neville, 1995; Kolani et al., 2012). The amount of free water

reduces as the hydration progresses in cement and binders (Choktaweekarn, 2008).

Thus the thermal properties change as the hydration of cement and slag proceeds.

Therefore, it is essentially beneficial to model the behaviour of free water in paste in

order to predict thermal properties of concrete.

4.2 Experimental Program

4.2.1 Materials

The binders used throughout this study were Ordinary Portland Cement type I

and Ground Granulated Blast Furnace Slag provided from Taiheiyo Cement

Corporation, Japan. The physical and chemical compositions of cement and slag are

shown in Table 4.1 and Table 4.2, respectively. Normal tap water was used in all

mixes. Small plastic boxes were used to cast and cure the paste specimens in sealed

condition.

Table 4.1 Physical properties of cement and slag

Physical properties Cement type I Slag

Specific gravity 3.16 2.89

Fineness (cm2/g) 3570 4330

Loss on ignition (%) 2.06 0.96

35

Table 4.2 Chemical compositions of cement and slag

Binder

type

Chemical composition (%)

SiO2 Al2O3 Fe2O3 CaO MgO SO3

Cement 20 5.4 3 63.4 2.7 2

Slag 33.47 14.35 0.26 43.26 5.14 2.1

4.2.2 Mix proportions

This experiment was conducted to obtain free water content inside paste

specimens with time. A total of eight mixtures were cast to test free water of paste

specimens. Water to binder ratios (w/b) of 0.25 and 0.40 were used. Slag

replacements of 0%, 45%, 60%, and 75% were utilized. Mixture designation is

denoted as “W40 S45” which indicates a paste mixture having w/b of 0.40 and slag

replacement of 45%. All mix proportions of the tested specimens are mentioned in

Table 4.3.

Table 4.3 Mix proportions of the tested paste specimens

Mix designation w/b c/b s/b

W25 S00 0.25 1 0

W25 S45 0.25 0.55 0.45

W25 S60 0.25 0.40 0.60

W25 S75 0.25 0.25 0.75

W40 S00 0.40 1 0

W40 S45 0.40 0.55 0.45

W40 S60 0.40 0.40 0.60

W40 S75 0.40 0.25 0.75

Remarks: w: water, c:cement, s:slag, and b:binders (c+s).

36

4.2.3 Specimen preparation and test method

This experiment was conducted by oven- drying the paste specimens in order

to measure the amount of free water removed from the specimens. This method was

used in previous studies to measure the free water amount in paste and mortars as well

(Saengsoy, 2003; Choktaweekarn, 2008). To make sure that all free water is removed,

specimens with small dimensions were cast.

Specimens for testing free water content were cast in plastic cube moulds with

dimensions 20×20×40mm and were kept sealed until tested. All specimens were

cured in 28±2°C and 50-70% relative humidity conditions until tested at 3, 7, and 28

days. Free water content was obtained by measuring the weight loss upon drying the

specimens at 105°C for 24±2 hours until the weight loss rate was less than 1% of the

total weight.

Free water was computed from the amount of weight loss of paste specimens

which were subjected to drying at 105°C for one day. It is quantified in kg/m3 of

paste.

4.3 Experimental Results

From this experiment, it is observed that the free water contents of paste

specimens clearly reduce with age due to the consumption of water by cement and

slag hydration. From the results, the weight ratio of free water to total binder content

is used to show the effect of water to binder ratio and slag replacement level. The

results of the tested specimens at various ages are shown in Fig 4.1 and Fig 4.2.

37

Fig. 4.1 Test results of weight ratio of free water to total binder of pastes with slag

replacement levels= 0%, 45%, 60%, and 75% at w/b=0.25

Fig. 4.2 Test results of weight ratio of free water to total binder of pastes with slag

replacement levels= 0%, 45%, 60%, and 75% at w/b=0.40

38

4.3.1 Effect of water to binder ratio

When comparing Fig. 4.1 with Fig. 4.2, it is clearly seen that when water to

binder ratio is higher, the free water amount increases. Specimens with w/b of 0.25

have lower free water amount than specimens with w/b of 0.4, regardless of the age.

The tendency is similar for both cement-only and slag-cement paste specimens.

4.3.2 Effect of slag replacement level

The effect of slag is clearly shown for both w/b cases where free water

increases when the slag substitution level is higher. Similar results were previously

reported by Kolani et al. (2012). This is due to lower reactivity of slag at early age

when compared to cement. However, the decreasing rate of free water is higher at

later age for paste with slag. At later age, the reaction of slag in the paste increases.

This is due to the activation of pozzolanic reaction of slag by Ca(OH)2, which is a

product of hydration of cement. Therefore, as hydration of slag continues at later age,

free water is reduced continuously as well.

4.4 Model for Predicting Free Water Content

4.4.1 Free water content

It is clearly seen from the results obtained from this study that free water is a

time-dependent property which significantly reduces as hydration progresses

especially at early age. Therefore, simulating free water content in pastes with slag is

archieved by adopting a time-dependent model which computes free water for paste

containing fly ash. This computation is done based on Eq. (4.1).

)t(W)t(WW)t(W wgelwhpwowfree (4.1)

Free water is computed from Eq. (4.1), where Wfw(t), Wfw0, Wwhp(t), and

Wwgel(t) are the weights of free water (kg/m3), unit water content of the mix (kg/m

3),

weight of water consumed by hydration of cement and reaction of slag (kg/m3), and

39

gel water content (kg/m3), respectively. Free water content in paste is reduced due to

increase of water consumed by hydrations of cement and slag and increase of gel

water content.

4.4.2 Chemically bound water content

In the hardening process of paste, water is consumed by cement and slag

hydrations as well as produces hydrated products. The amount of water involved in

producing hydrated products is computed as the chemically bound water. It is

modelled in this study assuming that the cement and slag hydrations are independent,

as shown in Eq. (4.2).

100

)t(W

100

)t(W)t(W

slag

0uslagslag,hp

hy

0ucc,hpwhp

(4.2)

where Wwhp(t) is the weight of water consumed by hydration of cement and

reaction of slag which is represented in kg/m3. Wuc0, Wuslag0, αhy(t) and αslag(t) are the

initial weight of cement (kg/m3), initial weight of slag (kg/m

3), average hydration

degree of cement (%) and hydration degree of slag (%), respectively.

θhp,c and θhp,slag are the ratios of minimum water to binder ratio required to

obtain maximum reactions of cement and slag, respectively. The value for θhp,c is

mentioned in Eq. (4.3) which was previously used in many studies for representing

chemically bound water of cement (Powers, 1960; Neville, 1995; Lam, Wong, &

Poon, 2000; Saengsoy, 2003). The value for θhp,slag is shown in Eq. (4.4) which was

used to estimate the amount of chemically bound water previously by Maekawa &

Ishida (2002) and Wang & Linger (2010).

21.0c,hp (4.3)

30.0slag,hp (4.4)

40

4.4.3 Gel water content

Gel water is defined as the water which is entrapped within the hydrated

products. Thus, gel water increases as the hydration process of cement and slag

continues. In this study the amount of gel water content is obtained by back analysis

from the test results of free water of paste. It is computed in this study from Eq. (4.5)

assuming that the cement and slag hydrations are independent. In the case of cement

hydration, the gel water is affected by water to binder ratio as previously reported by

Saengsoy (2003). In the case of slag reaction, the gel water is affected by water to

binder ratio and slag replacement level.

100

)t(W

)s1914.3exp(83703.1

08812.3b/w01.0

100

)t(W

b/w1414.0exp(009.1

0026.00126.0)t(W

slag

0uslag

hy

0ucwgel

(4.5)

` where Wwgel(t) is the weight of gel water content (kg/m3), w/b is the water to

binder ratio of the mix and s is the slag replacement level (%). Wuc0, Wuslag0, αhy(t) and

αslag(t) are the initial weight of cement (kg/m3), initial weight of slag (kg/m

3), average

hydration degree of cement (%) and hydration degree of slag (%), respectively.

4.5 Verification of Free Water Model

The proposed model is used to simulate the free water content at various ages

of the tested specimens of paste with different water to binder ratios and slag

replacement levels. The comparison between the model simulations and test results

are shown in Fig. 4.3 and Fig. 4.4. Fig. 4.3 indicates the comparison for paste

specimens with different slag replacement levels at a water to binder ratio of 0.25.

Fig. 4.4 indicates the comparison for paste specimens with different slag replacement

levels at a water to binder ratio of 0.40.

41

The model simulations show sufficient accuracy in predicting the test results

of paste specimens at different water to binder ratios and slag replacement levels. In

the case of slag-cement paste, it is seen that the model predictions show continuous

decrease in long term when compared to that of the cement paste due to slag

hydration.

Fig. 4.3 Comparison between test results and model simulations of weight ratio of

free water to total binder of pastes with slag replacement levels= 0%, 45%, 60%, and

75% at w/b=0.25

42

Fig. 4.4 Comparison between test results and model simulations of weight ratio of

free water to total binder of pastes with slag replacement levels= 0%, 45%, 60%, and

75% at w/b=0.40

43

Chapter 5

Model for Predicting Specific Heat

5.1 General

Specific heat is a time-dependent property which is changed rapidly especially

at the early age after casting of concrete (Choktaweekarn et al., 2009). However,

many studies have used constant specific heat values for analyzing thermal cracking

problems (Kwak & Ha, 2006; Guo et al., 2001). Modeling specific heat as a time-

dependent variable is essentially useful for more accurate and precise predictions of

temperature profiles.

Choktaweekarn et al. (2009) reported that specific heat of concrete at a given

time can be calculated based on weight fraction and specific heat of each constituent.

A time-dependent model was proposed in this previous study for fly ash concrete. In

this study, model proposed by Choktaweekarn et al. (2009) is modified for the effect

of slag in order to simulate the specific heat of slag concrete.


5.2.1 Materials and mix proportions


and Ground Granulated Blast Furnace Slag provided by Taiheiyo Cement

Corporation, Japan. The physical and chemical compositions of the cement and slag

are shown previously in Chapter 4 in Table 4.1 and Table 4.2, respectively. Normal

tap water was used as mixing water in all mixes.

This experiment was conducted to obtain specific heat of paste specimens with

time. A total of eight mixtures were cast to test free water of paste specimens. Water

to binder ratios (w/b) of 0.25 and 0.40 were used. Slag replacements of 0%, 45%,

60%, and 75% were utilized. All mix proportions of the tested specimens are

mentioned previously in Chapter 4 in Table 4.3.

44


For testing specific heat, specimens were cast with dimensions 50×50×50mm.

These cubes were removed 24 hours after casting and immediately wrapped firstly by

a layer of plastic sheet followed by a layer of aluminum foil to prevent moisture loss

to the environment. The firm wrapping was done in order to prevent evaporation of

water and to simulate no moisture loss to or gain from environment, similar to that of

the condition inside the mass concrete. All specimens were cured in 28±2°C and 50-

70% relative humidity conditions until tested at 3, 7, and 28 days. Each specimen was

smoothly divided into two pieces just before the test was done.

This experiment was conducted by a method based on transiently heated plane

sensor using the instrument named Hot Disk Thermal Constants Analyser (Model:

TPS 2500S). In this test, a hot disk sensor was fitted between the two pieces of paste

specimens with same mix proportions. Then the setup was covered with a sealed

container to prevent moisture movement. An electrical current was passed high

enough to increase temperature of the sensor and the resistance of the specimen was

recorded simultaneously as a function of time. A Ni foil probe which was wrapped in

Kapton was used in this experiment. Figs. 5.1a and 5.1b indicate the setup of the hot

disk sensor in between two pieces of the paste specimen, and the Hot Disk Thermal

Constants Analyser apparatus, respectively.

45

(a) Setup of the sensor sandwiched by a paste specimen

(b) Hot Disk Thermal Constants Analyser apparatus

Fig. 5.1 Testing specific heat using transient method

46


The specific heat values of cement and slag were experimentally obtained

using Differential Scanning Calorimetry (DSC) technique. These values were similar

to that of the values obtained previously by Krishnaiah & Singh (2006) and Bentz

(2007). The values for other constituents of concrete were obtained from previous

studies (Schutter & Taewe, 1995; ASHRAE, 1993; Klieger & Lamond, 1994; Roller,

2000). The values for specific heat of the constituents of concrete are mentioned in

Table 5.1.

The experimental results for specific heat of paste specimens with different

water to binder ratios and slag replacement levels are indicated in Fig. 5.2 and Fig.

5.3. From this experiment, it is observed that the specific heat of paste specimens

clearly reduce with age. The results of specific heat have higher correlation with that

of the results of free water of paste specimens with time. This can be explained by

considering the specific heat of each constituent of concrete. The specific heat of

water is the highest among all constituents in concrete as shown in Table 5.1.

Therefore, it is evident that free water content is the major factor which affects the

values of specific heat of concrete.

Table 5.1 Specific heat of the constituents of concrete

Constituents in concrete Specific heat (kcal/kg °C)

Cement 0.180*

Slag 0.167*

Water 1.000 (Schutter & Taewe, 1995)

Quartz sand 0.190 (ASHRAE, 1993)

Limestone 0.200 (Klieger & Lamond, 1994)

Air 0.240 (Roller, 2000)

Hydrated products Eq. (5.6)**

*The values were obtained from differential scanning calorimetry in this study

**The relationship was obtained from regression analysis in this study

47

Fig. 5.2 Test results of specific heat of pastes with slag replacement levels= 0%, 45%,

60%, and 75% at w/b=0.25

Fig. 5.3 Test results of specific heat of pastes with slag replacement levels= 0%, 45%,

60%, and 75% at w/b=0.40

48

` 5.3.1 Effect of water to binder ratio

When comparing Fig. 5.2 and Fig. 5.3, it is clearly seen that when water to

binder ratio is higher, the specific heat increases. Specimens with w/b of 0.25 have

lower specific heat than specimens with w/b of 0.4, regardless of the age. The

tendency is similar for both cement-only and slag-cement paste specimens. This is

due to the fact that free water governs the behavior of specific heat as mentioned

previously.


The effect of slag is clearly shown for both w/b cases from Fig. 5.2 and Fig.

5.3 where specific heat increases when the slag substitution level is higher. Similar

results were previously reported by Viet (2013) for mortar containing different

replacements of slag. This is due to lower reactivity of slag at early age when

compared to cement. However, the decreasing rate of specific heat is higher at later

age for pastes with slag. This is due to the activation of pozzolanic reaction of slag by

Ca(OH)2 which is a product of hydration of cement at later age. As hydration

progresses, free water decreases, resulting in continous decrease of specific heat of

paste containing slag.

5.4 Model for Predicting Specific Heat

It is clearly seen from the results obtained from this study, that specific heat is

a time-dependent property which is significantly affected by the availability of free

water at the considered age. Moreover, constituents of paste involve in hydration

process which lead to change in their reacted and un reacted compositions. Hence,

time-dependent change of each constituent of paste is vital to simulate specific heat of

concrete.

Model for simulating specific heat of concrete with slag is developed by

adopting a time-dependent model which computes specific heat of concrete

49

containing fly ash (Sarker et al, 1999; Saengsoy, 2003; Choktaweekarn, 2008). In this

model, specific heat of concrete at a given time can be calculated based on weight

fraction and specific heat of each constituent. Few assumptions are made in proposing

this model for specific heat. Specific heat of air is neglected due to smaller portion of

air compared to that of the other constituents. The changes of unit weight of concrete

during hydration process are assumed to have insignificant effect on specific heat.

The modified equations in this study are shown in Eq. (5.1), Eq. (5.2), Eq.

(5.3), Eq. (5.4) and Eq. (5.5). The specific heat of concrete can be computed at any

time from Eq. (5.1). The specific heat of coarse aggregate and fine aggregate is an

intristic property which does not depend on time. Therefore, fine and coarse aggregate

weights remain constant throughout the reaction. However, cement and slag react

with water and the amount of unreacted portions of those binders are time-dependent.

The weight fraction of un-reacted cement and slag can be calculated from Eq. (5.2)

and Eq. (5.3), respectively. The total weight fraction of concrete is kept constant and

can be computed using Eq. (5.4). The hydration process results in reducing un-reaced

binders but producing hydrated products. The time-dependent production of hydrated

products can be computed by Eq. (5.5).

hphpslaguslagcucwfwssgg c)t(wcwc)t(wc)t(wcwcw)t(c (5.1)

0c

hy

uc w100

)t(1)t(w

(5.2)

0slag

slag

uslag w100

)t(1)t(w

(5.3)

At specified time, t,

0.1wwwww 0slag0c0wsg (5.4)

)]t(w)t(w)t(www[0.1)t(w uslagucfwsghp (5.5)

50

where c(t) is the specific heat of concrete at the considered age t (kcal/kg/°C).

wg and ws are the weight ratios of coarse and fine aggregates per unit weight of

concrete, respectively. wfw(t), wuc(t), wuslag(t) and whp(t) are the weight ratios of free

water, unhydrated cement, non-reacted slag, and hydrated products, respectively, at

the considered age t. cg, cs, cw, cc, cslag, and chp are specific heat values of coarse

aggregate, fine aggregate, water, cement, slag, and hydrated products, respectively

(kcal/kg/°C). The values of cslag and cc are obtained from differential scanning

calarimetry. The specific heat values of all constituents in concrete are shown

previously in Table 5.1.

The value of chp is determined from regression analysis from the test results of

specific heat. The hydrated product properties can change with different water to

binder ratios and slag replacement levels. As the w/b increases, the specific heat of

hydrated products increases due to addition of water. As slag replacement increases,

specific heat of hydrated product decreases due to lower specific heat of slag

compared to that of cement. A relationship was obtained to compute specific heat of

hydrated products as a function of w/b and slag replacement which is shown in Eq.

(5.1).

)s168.0b/w357.0exp(0.094 chp (5.1)

where chp is the specific heat of hydrated product (kcal/kg °C), w/b is the water to

binder ratio, and s is the slag replacement ratio in paste.

51

5.5 Verification of Specific Heat Model

The proposed model is used to simulate specific heat at various ages of the

tested specimens of paste with different water to binder ratios and slag replacement

levels. The comparison between the model simulations and test results are shown in

Fig. 5.4 and Fig. 5.5. Fig. 5.4 indicates the comparison for paste specimens with

different slag replacement levels at a water to binder ratio of 0.25. Fig. 5.5 indicates

the comparison for paste specimens with different slag replacement levels at a water

to binder ratio of 0.40.


of specific heat in paste specimens at different water to binder ratios and slag

replacement levels. In the case of slag-cement paste, it is seen that the model

predictions show continuous decrease in long term when compared to that of the

cement paste due to slag hydration.

Fig. 5.4 Comparison between test results and model simulations of specific heat of

pastes with slag replacement levels= 0%, 45%, 60%, and 75% at w/b=0.25

52

Fig. 5.5 Comparison between test results and model simulations of specific heat of

pastes with slag replacement levels= 0%, 45%, 60%, and 75% at w/b=0.40

53

Chapter 6

Model for Predicting Thermal Conductivity

6.1 General

Thermal conductivity is known as the rate of heat transfer through a unit cross

sectional area of a material for a specific temperature gradient. As the hydration

process continues, amount of free water in concrete reduces with an increase in

hydrated products resulting in increasing thermal conductivity with time especially at

early stage. The thermal conductivity of concrete is affected by the mix proportions

and constituent types in the concrete as well.

There are two main techniques for measuring thermal conductivity which are

steady state and transient methods. Steady state method is useful when the material

under examination is rigid and dry or conditioned to the ambient condition

(Choktaweekarn, 2008). The method is not suitable when moisture redistribution can

occur during the period of the test. Transient measurement technique is appropriate

for low conductivity porous materials. The rapidity of the determination does not

allow sufficient time for any moisture movement to occur within the sample during

testing. Moisture has great effect on thermal conductivity of concrete then transient

method is preferable.








54

An experiment was conducted to obtain thermal conductivity of paste

specimens with time. A total of eight mixtures were cast to test thermal conductivity

of paste specimens. Water to binder ratios (w/b) of 0.25 and 0.40 were used. Slag

replacements of 0%, 45%, 60%, and 75% were utilized. All mix proportions of the

tested specimens are mentioned previously in Chapter 4 in Table 4.3.


Transient method which was previously used in this study to measure specific

heat, provides the output of thermal conductivity as well. Therefore, all the specimen

preparation and test procedure is similar to that described in Chapter 5.


The experimental results for thermal conductivity of paste specimens with

different water to binder ratios and slag replacement levels are indicated in Fig. 6.1

and Fig. 6.2. From this experiment, it is observed that the thermal conductivities of

paste specimens increase up to three days and after that slightly decrease. However,

some of 28 day results show a slight decrease which is probably due to the self-

desiccation of the specimens. If there is no loss of water from the specimens, thermal

conductivity is supposed to increase slightly with age due to increase of hydrated

products and continuity of paste structure.


Fig. 6.1 and Fig. 6.2 indicate that thermal conductivity values of paste

specimens with w/b 0.25 were slightly higher than that of the specimens with w/b

0.40. This is due to higher thermal conductivity of cement than that of the water as

shown in Table 6.1.

55

Table 6.1 Thermal conductivities of the constituents of concrete

Constituents in concrete Thermal conductivity (kcal/m hr °C)

Cement 1.33 (Bentz, 2007)

Slag 0.695*

Water 0.51 (Klieger & Lamond, 1994)

Quartz sand 3.00 (ASHRAE, 1993)


Air 0.026 (Roller, 2000)

Hydrated products 1.03*

* The value was obtained from regression analysis in this study


The effect of slag is shown for both w/b cases in Fig. 6.1 and Fig. 6.2 where

increase of slag replacement level, result in a slight decrease in the values. This is due

to the lower thermal conductivity of slag when compared to cement as shown in Table

6.1. Slag decreases the density hence increase porosity of paste which result in lower

thermal conductivities as mentioned previously in many studies (Demirboga, 2003;

Demirboga, 2007; Choktaweekarn, 2008).

56

Fig. 6.1 Test results of thermal conductivity of pastes with slag replacement levels=

0%, 45%, 60%, and 75% at w/b=0.25

Fig. 6.2 Test results of thermal conductivity of pastes with slag replacement levels=

0%, 45%, 60%, and 75% at w/b=0.40

57

6.4 Model for Predicting Thermal Conductivity

Model for simulating thermal conductivity of concrete with slag is developed

by adopting a time-dependent model which computes thermal conductivity of

concrete containing fly ash (Sarker et al, 1999; Choktaweekarn, 2008). In this model,

thermal conductivity of concrete at a given time can be calculated based on volume

fraction and thermal conductivity of each constituent. Few assumptions are made in

proposing this model. The changes of total volume of concrete during hydration

process are assumed to have insignificant effect on the simulations. Moreover, the

volume decrease of the hydrated products during hydration process is assumed to be

insignificant as well.

The modified equations in this study are shown in Eq. (6.1), Eq. (6.2), Eq.

(6.3), and Eq. (6.4). The thermal conductivity of concrete at any age can be computed

from Eq. (6.1). Similar to that of the specific heat model, fine and coarse aggregate

weights remain constant throughout the reaction process. The volume fractions of un-

reacted cement and slag can be calculated from Eq. (6.2) and Eq. (6.3), respectively.

The volumetric ratio of hydrated products can be computed by Eq. (6.4).

hphpraraslaguslagcucwfwssgg k)t(nknk)t(nk)t(nk)t(nknkn)t(k (6.1)

0c

hy

uc n100

)t(1)t(n

(6.2)

0slag

slag

uslag n100

)t(1)t(n

(6.3)

)]t(n)t(n)t(nnnn[0.1)t(n uslagucfwrasghp (6.4)

58

where k(t) is thermal conductivity of concrete at considered age(kcal/m hr oC),

kg, ks, kw, kc, kslag, kra, khp are thermal conductivities of coarse aggregate, fine

aggregate, free water, cement, slag, air and hydrated products, respectively. The

values are shown in Table 6.1. ng, ns and nra, are volumetric ratios of coarse

aggregate, fine aggregate and air, respectively. nfw(t), nuc(t), nuslag(t) and nhp(t) are

volumetric ratios of free water, unhydrated cement, non-reacted slag and hydrated

product at the considered age, respectively. nc0 and nslag0 are the initial volume

fractions of cement and slag, respectively.

6.5 Verification of Thermal Conductivity Model

The proposed model is used to simulate thermal conductivity at various ages

of the tested specimens of paste with different water to binder ratios and slag

replacement levels. The comparisons between the model simulations and test results

are shown in Fig. 6.3 and Fig. 6.4. Fig. 6.3 indicates the comparison for paste

specimens with different slag replacement levels at a water to binder ratio of 0.25.

Fig. 6.4 indicates the comparison for paste specimens with different slag replacement

levels at a water to binder ratio of 0.40.


of thermal conductivity of paste specimens at different water to binder ratios and slag

replacement levels. The experimental results at 28 days slightly decreased due to self

dessication especially in specimens with w/b 0.25. However, the model predictions

increase slightly at later age since the continuity of the structure increases if no

moisture movement occurs.

59

Fig. 6.3 Comparison between test results and model simulations of thermal

conductivity of pastes with slag replacement levels= 0%, 45%, 60%, and 75% at

w/b=0.25

Fig. 6.4 Comparison between test results and model simulations of thermal

conductivity of pastes with slag replacement levels= 0%, 45%, 60%, and 75% at

w/b=0.40

60

Chapter 7

Model for Predicting Coefficient of Thermal Expansion

7.1 General

Temperature difference in concrete causes thermal stress which leads to

thermal cracking. One of the key functions in computing thermal strain due to the

occurrence of thermal gradient in concrete is the coefficient of thermal expansion

(CTE). It is known as the unit length change per unit degree of temperature change.

The CTE of concrete depends on the CTE of the constituents which are cementitious

materials, water, hydrated products and aggregates. In this study, CTE of paste

containing slag is experimentally evaluated and time-dependent model is proposed

based on volumetric fractions and CTE of constituents of pastes.








This experiment was conducted to obtain CTE of paste specimens with time.

A total of eight mixtures were cast to test CTE of paste specimens. Water to binder

ratios (w/b) of 0.25 and 0.40 were used. Slag replacements of 0%, 45%, 60%, and

75% were utilized. All mix proportions of the tested specimens are mentioned

previously in Chapter 4 in Table 4.3.

61


Kada et al. (2002) developed a simple method to find the CTE at early age of

concrete. This method was based on applying a temperature shock in a range of 10°C

to 50°C in a short period, not longer than one hour. Since the duration for each

measurement step was short, the effect of autogeneous shrinkage was not considered

in this experiment. The similar method was adopted in a previous study

(Choktaweekarn, 2008) for testing paste, mortar and concrete.

In this study, CTE was tested by the method used in Choktaweekarn’s study

(Choktaweekarn, 2008). The temperature of the paste specimens was reduced by

cooling them down in a refrigerator until 10°C, then followed by moving out to return

the temperature back to room temperature (about 30±2°C).The specimens were tested

for the change in length at every 5°C change of temperature, then CTE values were

computed. The step by step of temperature changing process is illustrated in Fig. 7.1.

Step 1 Step 2 Step 3 Step 4

Step 8 Step 7 Step 6 Step 5

Fig. 7.1 Temperature changing process to measure CTE of slag-cement pastes

Prism specimens with dimensions 25×25×285mm were cast for slag-cement

paste. The specimens were firmly sealed first by paraffin layer, then plastic and

followed by aluminium foil, immediately after casting in order to prevent evaporation

of water. All specimens were cured in 28±2°C and 50-70% relative humidity

conditions until tested at 1, 3, 7, and 28 days. To obtain data of the temperature,

thermocouples were inserted. Fig. 7.2 shows the sealed specimens which were used to

measure CTE of the pastes. Fig. 7.3 illustrates the setup of the experimental process

where the length change was measured by a length comparator at each step of

temperature change as described in Fig. 7.1.

Room temperature, 30±2°C 25°C 20°C

10°C

15°C

62

Fig. 7.2 Example of firmly wrapped paste specimens for measuring CTE

Fig. 7.3 Experimental setup for measuring CTE of paste specimens

It was reported that autogenous shrinkage does not signifantly affect the

measurement of CTE in fly ash-cement paste (Choktaweekarn, 2008). However,

many studies reported that autogenous shrinkage of paste containing slag was

considerably higher than that of the paste with fly ash (Johari, 2000; Matsuka et al.,

2010). Therefore, autogenous shrinkage was measured in this study to find its effect

63

on CTE. It was reported that autogeneous shrinkage of pastes containing slag increase

as the temperature increases (Matsuka et al., 2010). Thus, the measurements were

taken at 30±2°C since it was the maximum temperature used during the test period.

Even at 30±2°C, the strain due to autogeneous shrinkage was found to be less than

±1% when compared to the strain due to temperature change, during the test period.

Thus, the autogeneous shrinkage was considered to have insignificant effect therefore,

it was not included in computing CTE in this study. If the effect of autogeneous

shrinkage is significant, it is recommended to obtain measurements at each

temperature step in the test thereby include the values to compute CTE.

It was previously reported that for a specific concrete, the level of thermal

expansion or contraction at normal temperature conditions was similar for each unit

temperature change (Choktaweekarn, 2008; Amonamarittakul, 2011). Therefore,

length change measurements were done for every 5°C change of temperature. At the

same time, the autogeneous shrinkage was measured. Then, the CTE is calculated

based on Eq. (7.1).

TCTE AS

(7.1)

CTE is the coefficient of thermal expansion of paste (micron/oC), ε is strain

due to temperature change, εAS is the strain due to autogeneous shrinkage during the

period of ∆T which is the temperature change (oC). The plus sign is used when

heating up and minus is used when cooling down.


The experimental results of CTE of paste specimens with different water to

binder ratios and slag replacement levels are illustrated in Fig. 7.4 and Fig. 7.5. From

this experiment, it is observed that the CTE of paste specimens is a time-dependent

property which increases with time. This is due to the fact that continuity of the

structure increases when paste transfer from fresh state to hardened state. Many

64

studies have reported similar tendency for CTE experimental results done on paste,

mortar, and concrete (Berwanger & Sarker, 1976; Choktaweekarn, 2008;

Amonamarittakul, 2011).


Fig. 7.4 and Fig. 7.5 indicate that CTE results of paste specimens with w/b

0.25 are slightly higher than that of the specimens with w/b 0.40. The slight decrease

of CTE as w/b increases may be due to lower amount of cement content. Similar

results were observed previously by Berwanger and Sarker (1976) as well. However,

it was also reported that water behaves somewhat different from solid materials.

When the temperature changes, volume of water in capillary pores of paste increases.

However, water in the pores is capable of moving from pores to pores resulting in

different effect, from that of the solid, on CTE of paste (Choktaweekarn, 2008).


Effect of slag replacement level can be seen by comparing Fig. 7.4 with Fig.

7.5. The CTE of pastes decrease as the slag replacement level increases which is due

to lower CTE of slag itself. The CTE values of the constituents of concrete are shown

in Table 7.1. However, the tendency may change in long term as the hydration of slag

continues with time.

Table 7.1 CTE values of the constituents of concrete

Constituents in concrete CTE (micron/ °C)

Cement 14.4 (Choktaweekarn, 2008)

Slag 8.8 (Okura & Imaoka, 1979)

Quartz sand 10.4 (Klieger & Lamond, 1994)


Hydrated products Eq. (7.3)*

* The relationship was obtained from regression analysis in this study

65

Fig. 7.4 Test results of CTE of pastes with slag replacement levels= 0%, 45%, 60%,

and 75% at w/b=0.25

Fig. 7.5 Test results of CTE of pastes with slag replacement levels= 0%, 45%, 60%,

and 75% at w/b=0.40

66

7.4 Model for Predicting CTE

Model for simulating CTE of paste with slag is developed by adopting a time-

dependent model which computes CTE of paste containing fly ash (Choktaweekarn,

2008). In this model, the CTE of paste at a given time can be calculated based on

volume fractions of non-reacted binder content and hydrated product amounts.

Hydrated products increase with time resulting in reducing the non-reacted cement

and slag. The water content affect the CTE in the very early age, however, afterwards

the results are governed by the continuity of structure (Amonamarittakul, 2011). The

proposed model for computing CTE of slag cement paste is shown in Eq. (7.2).

hphp,pslaguslag,pcuc,pp CTE)t(ncCTE)t(nbCTE)t(na)t(CTE (7.2)

CTEp(t) is the CTE of paste at considered age (microns/oC), CTEc, CTEslag and

CTEhp are the CTE of cement, slag and the hydrated products which are given in

Table 7.1. np,uc(t), np,uslag(t) and np,hp(t) are the volumetric ratios of unhydrated cement,

non-reacted slag and hydrated products at the considered ages, which are computed

using similar principles as mentioned in thermal conductivity model in Chapter 6. The

constants a, b and c are derived by regression analysis from test results of pastes with

and without slag. The values are 0.284, 2 and 1.4 respectively.

The relationship for CTEhp is determined from regression analysis of the test

results of CTE. The hydrated product properties can change with different water to

binder ratios and slag replacement levels. As the w/b increases, the CTE of hydrated

products decreases due to reduction of cement amount. As slag replacement increases,

CTE of hydrated product decreases due to lower CTE of slag when compared to that

of cement. A relationship was obtained to compute CTE of hydrated products as a

function of w/b and slag replacement as shown in Eq. (7.3).

)0533.1b/w133.0(006.20s86.5s93.5CTE 2

hp (7.3)

67

where CTEhp denotes the CTE of hydrated product (microns/oC), w/b is the

water to binder ratio, and s is the slag replacement ratio in paste.

7.5 Verification of CTE Model

The proposed model is used to simulate CTE at various ages of the tested

specimens of paste with different water to binder ratios and slag replacement levels.

The comparison between the model simulations and test results are shown in Fig. 7.6

and Fig. 7.7. Fig. 7.6 indicates the comparison for paste specimens with different slag

replacement levels at a water to binder ratio of 0.25. Fig. 7.7 indicates the comparison

for paste specimens with different slag replacement levels at a water to binder ratio of

0.40.

It is seen that the model simulations show sufficient accuracy in predicting the

test results of CTE in paste specimens at different water to binder ratios and slag

replacement levels. The model predictions indicate that CTE decreases as water to

binder ratio and slag replacement level decreases.

Fig. 7.6 Comparison between test results and model simulations CTE of pastes with

slag replacement levels= 0%, 45%, 60%, and 75% at w/b=0.25

68

Fig. 7.7 Comparison between test results and model simulations CTE of pastes with

slag replacement levels= 0%, 45%, 60%, and 75% at w/b=0.40

69

Chapter 8

Model for Simulating Adiabatic Temperature of Mass Concrete

8.1 General

A computerized program was developed for simulating temperature in mass

concrete from a previous study (Saengsoy, 2003). It is capable of simulating the

temperature in fly ash concrete. This program was modified in order to obtain

adiabatic temperature rise of slag concrete. The development of the models to predict

hydration degrees of cement and slag, free water, specific heat were described

previously in Chapters 3, 4, and 5. If the total heat generation in concrete is known,

the temperature can be computed based on a basic equation which is shown in Eq.

(8.1).

mc

QT (8.1)

where Q is the cumulative heat of hydration (kcal), m is mass of the concrete

(kg), c is specific heat of the concrete (kcal/kg/°C), and ΔT is temperature rise at the

considered age (°C). The procedure for computing total heat generation in concrete is

described in this chapter which is followed by simulating adiabatic temperature rise of

concrete containing slag with different mix proportions.

8.2 Total Heat Generation of Concrete

All reactive compounds of cement involve in reactions which cause the

temperature rise in concrete. It is well known that the four major mineral compounds

of cement are C3S, C2S, C3A, and C4AF. The proportions of these compounds are

different in different types of cement. Thus their heat generation also varies

depending on the amount and reaction rate. On the other hand, reaction of slag also

generates heat which results in temperature rise in concrete.

70

Since the hydration degrees of cement compounds and slag are modeled

assuming that the reactions are independent, the total heat generation is also modeled

considering the similar assumption. The total heat generation is computed by the

summation of heat liberated from each cement compound and slag as shown in Eq.

(8.2).

)t(Q)t(Q)t(Q)t(Q)t(Q)t(Q)t(Q)t(Q slagAFETCAETCAFCACSCSC 434323

(8.2)

eff,imax,ii

i w*Q*100

)t()t(Q

(8.3)

ieff,i ww (8.4)

58.0b/w5.0s4037.0s579.0 2 (8.5)

Total heat generation is computed from Eq. (8.2) where Q(t) is the total heat

generation of concrete at the considered age t (kcal/kg of concrete), QC3S(t), QC2S(t),

QC3A(t), QC4AF(t), and Qslag(t) are cumulative heat generation of C3S, C2S, C3A, C4AF

and slag at the considered age t, respectively (kcal/kg of concrete). QC3AET(t) and

QC4AFET(t) are heat generation from ettringite and monosulphate production by C3A

and C4AF reactions with gypsum at the considered age t, respectively (kcal/kg of

concrete).

Heat generation from the compounds of cement and slag is computed based on

Eq. (8.3). Qi(t) is the cumulative heat generation from the ith

compound at the

considered age t (kcal/kg of concrete), αi(t) is the hydration degree of each compound

(%), Qi,max is the maximum cumulative heat generation at completion of the chemical

reaction of the ith

compound (kcal/kg of concrete), and wi,eff is the effective mass

fraction of each compound in the concrete. The compounds considered (i) in this

model are C3S, C2S, C3A, C4AF, and slag.

Slag involves in self-hydration and pozzolanic reactions (Kolani et al., 2012).

In this study, heat generation by slag is assumed as a single parameter because of

71

limited amount of data on the heat generated by the two separate reactions of slag.

Thus, it is recommended in the future to separate the heat generation from slag

hydration and pozzolanic reaction in order to interpret the two reactions

independently.

The effective mass fraction wi,eff , can be computed using Eq. (8.4) where wi is

the mass fraction of each compound in the concrete and Φ is the factor which includes

the effect of slag towards the adherence of cement particles. The adherence of cement

particles is defined as the amount of cement that can actually involve in hydration and

it depends on the type and source of cement (Choktaweekarn, 2008). This causes

some amount of cement to not properly mixed and actively involved in hydration

process. The use of slag may lead to better dispersion of the particles of cement

especially for higher fineness, enabling higher effective weight of cement to involve

in hydration. However, due to irregular particle shape of slag, the particles may

interlock together causing reduction of the dispersion effect in excessive replacements

of slag. It is evident that adherence of particles reduces as the amount of water in the

mix increases as well. Thus, a dispersion factor was introduced for interpreting the

effect of slag, as shown in Eq. (8.5). In this equation, s represents the slag

replacement ratio and w/b is the water to binder ratio of the concrete. This factor is

modeled based on the available test results of adiabatic temperature rise, mostly

Japanese cement sources. Thus, further modification should be done for different

sources of cements.

In the case of fly ash concrete, a separate dispersion factor was proposed

previously by (Choktaweekarn, 2008). A comparison of the dispersion effects of fly

ash and slag which affect the adherence of cement particles is shown in Fig. 8.1.

When comparing, it is seen that the dispersion factor for concrete with fly ash

replacements higher than 30% shows higher dispersion effect when compared to slag

concrete. This is mainly due to the spherical shape of fly ash which enables to slip

each other easily and reduce the inter particle friction forces. Whereas in the case of

slag, this factor decreases when the slag replacement ratio is very high. Slag may

increase dispersion in lower replacement ratios due to finer particles, however,

72

Fig. 8.1 Comparison of dispersion factors of fly ash and slag which affect the

adherence of cement particles in concrete

irregular shape of the slag particles may result in reduction of the dispersion effect in

excessive replacements.

The maximum cumulative heat generation by slag is estimated at 150kcal/kg

based on tendencies of heat rate in adiabatic temperature rise data obtained from Tada

et al. (2014). It is slightly higher than the value proposed by Kishi & Maekawa

(1996) as 110kcal/kg. This difference may be mainly due to dissimilar physical and

chemical properties of the slags.

The maximum heat generation of cement compounds is previously

summarized by Saengsoy (2003). However, the heat generation of each compound is

different which mainly depends on type and source of the cement. In this study, the

values for maximum heat generation of cement compounds were adopted from a

Japanese research (Kishi & Maekawa, 1996) for Japanese cement sources whereas for

other cases similar values were utilized as mentioned by Saengsoy (2003). The

73

theoretical values for heat of hydration of main cement compounds are summarized in

Table 8.1.

Table 8.1 Maximum heat generation values for major cement compounds

Major phase Maximum heat of hydration (kcal/kg)

Saengsoy (2003) This study

C3S 105 115

C2S 50 50

C3A 190 195

C4AF 85 85

8.3 Verifications of Proposed Adiabatic Temperature Model

Since the hydration degrees of cement and slag, free water content, specific

heat and total heat generation of concrete are separately modeled, the models can be

linked to compute the adiabatic temperature rise by using Eq. (8.1) as mentioned

earlier. In order to validate the proposed model, verifications are carried out on

concrete incorporating slag. Experimental data were obtained from various sources to

check the accuracy of the described model (Wang & Linger, 2010; Tada et al., 2014;

Taiheiyo Cement Corporation, 2014; TCC-Singapore, 2014).

Verifications for the experimental data obtained from Wang & Linger (2010)

are shown in Fig. 8.2 and Fig. 8.3. In this previous study, investigations were done at

two different initial temperatures, 20°C and 30°C. Comparison between the adiabatic

temperature of concrete incorporating 40% slag with a water to binder ratio of 0.392

and another concrete mixture with 40% slag and water to binder ratio of 0.493 are

shown in Fig. 8.2 and Fig. 8.3, respectively. The mix proportions and physical and

chemical compositions of binder are shown in Appendix A (Tables A1, A3, and A4,

respectively).

74

Fig. 8.2 Comparison of tested and predicted adiabatic temperature rise for concrete

with 40% slag, w/b=0.392 and initial temperatures of 20°C and 30°C


with 40% slag, w/b=0.493 and initial temperatures of 20°C and 30°C

75

The verifications shown in Fig. 8.2 and Fig. 8.3 indicate that the proposed

model is able to simulate adiabatic temperature rise of the tested slag concrete

specimens with a reasonable accuracy.

Tada et al. (2014) experimentally investigated adiabatic temperature rise of

concrete with and without slag replacements. Concrete containing only cement at w/b

ratios of 0.40 and 0.65 were tested. Concrete mixes containing 50%, 65%, and 75%

slag replacement levels at w/b ratios of 0.4 and 0.6, were tested to find the effect of

slag replacement. The mix proportions as well as physical and chemical compositions

of the binders are shown in Appendix A (Tables A2, A3, and A4, respectively). The

verifications for these experimental results are shown in Fig. 8.4 to Fig. 8.7.

Comparison of the adiabatic temperature of cement only concrete with w/b ratios of

0.40 and 0.65 are shown in Fig. 8.4. Comparison of the adiabatic temperature of

concrete incorporating 50% slag, 65% slag, and 75% slag with a w/b of 0.4 and 0.6

are shown in Fig. 8.5, Fig. 8.6, and Fig. 8.7, respectively.

The model simulations predict the experimental results with sufficient

accuracy. However, the model simulations predict lower results at early age and

continuously increase with time. This is because the results mainly follow hydration

degree of slag which is modeled to increase continuously with time. Moreover, these

variations of the predictions are considered to be due to lack of consideration of effect

of some compounds in cement and slag which are not taken into account in this study.

These compounds may alter the early age hydration rates of both cement and slag. It

was reported that the impurities included in recent cement production and the SO3

content of slag can alter the hydration rates of cement and slag (Shinwa et al., 2009).

Therefore, interaction of slag-cement behavior should be further investigated for an

enhanced accuracy by also taking into consideration of some chemical compounds of

the binders.

76

Fig. 8.4 Comparison of tested and predicted adiabatic temperature rise for cement

concrete at w/b=0.4 and 0.65


with 50% slag and at w/b=0.4 and 0.6

77





78

Taiheiyo Cement Corporation has experimentally investigated temperature of

slag concrete by using different sources of the slag and under different environmental

conditions. One such experimental result was obtained from a report which evaluated

a slag manufactured in Korea (R & D department I, 2014). A concrete mix with 65%

slag replacement level at a w/b of 0.50 was tested in this experiment. The mix

proportions and physical and chemical compositions of binder are shown in Appendix

A (Tables A1, A3, and A4, respectively). The verification is done by using the

proposed model and it is shown in Fig. 8.8.

A practical evaluation of temperature rise of a foundation construction was

done by Taiheiyo Cement Corporation (R & D department II, 2014), in Singapore. A

concrete mix with 74% slag blended cement at w/b of 0.46 was utilized and the

temperature measurements were taken inside a mass concrete foundation. The

temperature profile seemed to show characteristics of an adiabatic temperature

profile, at the center of this mass concrete structure. The proposed model is applied to

verify the measured temperature profile which is shown in Fig. 8.9. The mix

proportions and physical and chemical compositions of the binder are shown in

Appendix A (Tables A1, A2, and A3, respectively).

79


with 65% slag and at w/b=0.50


with 74% slag and at w/b=0.46

80

The model simulations predicted the experimental results as shown in Fig. 8.8

and Fig. 8.9 with sufficient accuracy. The model simulation seemed higher than that

of the experimental results at later age as shown in Fig. 8.9. This is due to the fact that

the heat loss to surrounding was ignored in the simulation. In the real mass concrete

construction, there could be a slight heat loss to the environment.

The proposed model can be applied in the real mass concrete construction to

investigate the temperature rise of slag concrete. However, few modifications are

required for the proposed model to enhance the accuracy of the predictions. The

adiabatic temperature simulation model for slag concrete can be further modified in

depth to investigate the effect of chemical properties of cement and slag. Moreover,

the accuracy of hydration degrees can be further enhanced as concerning different

types and sources of cement and slag. The effects of water reducing and retarding

admixtures are not included in this model which can be further investigated.

81

Chapter 9

Semi-Adiabatic Temperature Rise of Mass Concrete

9.1 General

Adiabatic temperature rise has been experimentally investigated by many

researchers in order to investigate the temperature profiles in an environment which

do not allow the concrete specimen to have any interaction with surrounding

environment. However, despite of having large dimensions, in practical situations of

mass concrete structures are somewhat affected by the surrounding. Therefore, it is

required to consider the dimensions, environment behavior and boundary conditions

in order to simulate the actual temperature profile inside the mass concrete.

In this study, an experimental investigation on semi-adiabatic temperature rise

of mass concrete samples with fly ash and slag was carried out. Then, the temperature

rise profiles at the center of the specimens were analyzed using heat rates. The heat

rate is vital since this may directly affect thermal cracking potential of mass concrete

structures.



The binders used throughout this study were Ordinary Portland Cement type I,

a blended cement with 50% low calcium fly ash and slag provided by Taiheiyo

Cement Corporation, Japan. The physical and chemical compositions of the cement

and slag are shown previously in Chapter 4, in Table 4.1 and Table 4.2, respectively.

The properties of fly ash in the blended cement are shown in Appendix B (Table B1

and Table B2). Fine and coarse aggregates with specific gravities of 2.58 and 2.72

were used in saturated surface-dry condition. Normal tap water was used as mixing

water. A napthalene-based superplastizer was used for mixes with w/b 0.4.

This experiment was conducted to obtain the semi-adiabatic temperature rise

profiles of mass concrete samples containing different replacements of fly ash and

82

slag at different water to binder ratios. Water to binder ratios (w/b) of 0.40, 0.50 and

0.60 were used. Slag replacements of 50%, 65%, and 75% and fly ash replacements of

30% and 50% were utilized. The replacement levels of fly ash and slag were decided

considering the practical limits of substitutions of mineral admixtures in mass

concrete as clarified by many researchers (Japan Concrete Institute, 2008; Slag

Cement Association, 2002; Thomas, 2007). The mix design of the binders used in this

study are shown in Table 9.1. All mix proportions of concrete used for testing semi-

adiabatic temperature rise of mass concrete samples are indicated in Appendix B

(Table B3).

Table 9.1 Mix design of the binders used in casting concrete

Mix w/b c/b r/b s/b

W40 S/R00 0.4 1.00 0 0

W40 R30 0.4 0.70 0.30 0

W40 R50 0.4 0.50 0.50 0

W40 S50 0.4 0.50 0 0.50

W40 S65 0.4 0.35 0 0.65

W40 S75 0.4 0.25 0 0.75

W50 S/R00 0.5 1.00 0 0

W50 R30 0.5 0.70 0.30 0

W50 R50 0.5 0.50 0.50 0

W50 S50 0.5 0.50 0 0.50

W50 S65 0.5 0.35 0 0.65

W50 S75 0.5 0.25 0 0.75

W60 S/R00 0.6 1.00 0 0

W60 R30 0.6 0.70 0.30 0

W60 R50 0.6 0.50 0.50 0

W60 S50 0.6 0.50 0 0.50

W60 S65 0.6 0.35 0 0.65

W60 S75 0.6 0.25 0 0.75

Remarks: w: water, c:cement, r:fly ash, s:slag, and b:binders (c+r+s).

83


Simplified semi-adiabatic temperature rise of mass concrete samples

incorporated with fly ash and slag were tested. The objective was to measure

temperature rise in the center of mass concrete samples with dimensions of

40×40×40cm. The thickness of the foam layer and plywood were 50mm and 15mm,

respectively. Thermocouples were installed at the center of the specimens before

casting. Type K nickel/chromium thermocouples, calibrated against a thermometer

were used. The readings were obtained using a data logger at 10-minute intervals. All

specimens were put in 30±2°C and 50-70% RH conditions. The temperature changes

inside the concrete samples were measured immediately after casting and continued

up to 7 days. The setup of formwork for measuring the semi-adiabatic temperature

rise is shown in Fig. 9.1.

Fig. 9.1 Setup of formwork for measuring semi-adiabatic temperature rise of a mass

concrete sample

84


The semi-adiabatic temperature rise profiles of the tested specimens having

different mix designs upto 8 days are shown in Fig. 9.2 to Fig. 9.7. It is seen that the

temperature in all the tested specimens rise up to a certain value which is followed by

a continous drop due to loss to the surrounding environment.

9.3.1 Effect of w/b

For concrete containing fly ash, Fig. 9.2, Fig. 9.3, and Fig. 9.4 show the effect

of w/b ratios whereas those for slag concrete samples are shown in Fig. 9.5, Fig.9.6,

and Fig. 9.7. It can be seen that for both mineral admixtures, the temperature of

concrete having similar replacement levels clearly reduces as the w/b increases from

0.4 to 0.6. This is due to a reduction of cement amount as the w/b ratios increase.

Then the weight of the cement compounds is reduced resulting in lowering the total

heat generation. Hence, temperature inside concrete specimen decreases as w/b

increases.

9.3.2 Effect of fly ash

The effect of fly ash on temperature of concrete with w/b of 0.4, 0.5, and 0.6 is

shown in Fig. 9.2, Fig. 9.3, and Fig. 9.4, respectively. It is observed that for mixes

having same w/b, peak temperatures are reduced as the fly ash replacement increases.

This is due to low reactivity and heat generation of fly ash concrete when compared to

cement-only concrete. Moreover, the time to attain peak temperature is slightly

delayed as the replacement increases. This is mainly because of the lower heat

generation and delayed hydration rate of cement at early age of fly ash concrete.

85

Fig 9.2 Test results of semi-adiabatic temperature of concrete with fly ash 0%, 30%,

and 50% at w/b=0.4


and 50% at w/b=0.5

86


and 50% at w/b=0.6

9.3.3 Effect of slag

The effect of slag on temperature of concrete with w/b of 0.4, 0.5, and 0.6 is

shown in Fig. 9.5, Fig. 9.6, and Fig. 9.7, respectively. It is observed that for mixes

having same w/b, peak temperatures are reduced as the slag replacement increases.

This is due to low reactivity and heat generation of slag concrete when compared to

cement-only concrete. Moreover, the time to attain peak temperature is slightly

delayed in slag concrete when compared to cement-only concrete. However, slight

acceleration of the time to attain peak temperature is observed as the slag replacement

increases. This may be because slag has the ability to accelerate the hydration of

certain compounds of cement (Ogawa et al., 1980; Hoshino et al., 2006). It may lead

to a slight acceleration of the time to attain peak temperature as the slag replacement

increases. However, further investigation is recommended to find the other probable

causes for this acceleration.

87

Fig 9.5 Test results of semi-adiabatic temperature of concrete with slag 0%, 50%, and

75% at w/b=0.4


75% at w/b=0.5

88


75% at w/b=0.6

Comparison of the fly ash and slag concrete can be done by analysing the

temperature profile characteristics. The initial temperature, peak temperature, time to

attain peak temperature, and heating/cooling slopes are calculated as shown in

Appendix C (Table C1). Lower peak temperature values are observed in fly ash

concrete when compared to that of slag concrete. This can be due to lower heat

generation from fly ash. The time to attain peak temperature is delayed in both cases

when compared to that of cement-only concrete. However, slag concrete showed

slightly quicker time to attain peak temperature as the replacement increased,

whereas, the time delayed in the case of fly ash concrete. This may be due to

acceleration of cement hydration by slag whereas retardation caused by fly ash.

However, further investigation is recommended to find the probable causes for this

acceleration.

The heating up and cooling down slopes are computed based on tangents of

semi-adiabatic temperature profiles (Nili & Salehi, 2010). Heating up slope is

calculated from the difference between maximum temperature and initial temperature.

89

Cooling down slope is calculated from the difference between the peak temperature

and the temperature observed at 90 hours after casting. The obtained values for the

concrete mixes with w/b of 0.5 are summarized in Fig. 9.8. It is clearly seen that fly

ash concrete significantly reduces heating up slopes than that of slag concrete. It

shows that fly ash concrete performs better in lowering heating rate in concrete. The

use of fly ash can be essentially beneficial for mass concrete structures to reduce the

temperature during the period in which the hydration heat is accumulated. The

cooling down slopes are not much different.

Fig 9.8 Heating up and cooling down slopes computed from temperature profiles of

fly ash and slag concrete with w/b=0.50

90

Chapter 10

Initial slump

10.1 General

Workability is known as the quality of concrete obtained such that it could be

placed, compacted, pumped, finished, etc. easily with enough resistance to

segregation at the same time (Kitticharoenkiat, 2000). It is apparent that the main

factor for workability is the amount of free water content in the concrete mix.

Addition of water results in increasing the inter-particle lubrication. However, the

amount of water in fresh concrete can be further divided into two sections. One of

them is the water retained by solid particles which moves together with solids. The

other is defined as free water which is independent from the solid particles. The

hydration process is insignificant when dealing with initial slump of concrete.

Slump tests are carried out in the laboratory and sites to estimate the degree of

wetness of concrete or in other words consistency of concrete. Consistency is a

practical consideration in securing a workable concrete. Slump prediction model was

previously proposed by Kitticharoenkiat (2000) on the basis that slump value linearly

varies with free water content. This was further modified by Khunthongkeaw (2001)

and Wangchuk (2003). The initial slump can be computed by Eq. (10.1).

)WW(SL ofr (10.1)

where SL is the slump value of fresh concrete (cm), α is the slope of slump-

free water content curve of fresh concrete (cm/kg/m3), Wfr is the volume of free

water in fresh concrete (kg/m3), and Wo is the amount of water just enough to

overcome inter particle surface forces.

The slope of slump-free water content has been verified to have relationship

with the ratio of paste volume to void content of aggregate phase which is γ

91

(Wangchuk, 2003). The slope of slump-free water content curves (α) increase with the

increase of γ. The relationship of the above two parameters is shown in Eq. (10.2).

944.14916.4374.463.21573.3 234 (10.2)

The free water content (Wfr) is described as the amount of water that is free,

by any means, from being restricted by all solid particles in the fresh concrete. This

could be calculated using Eq. (10.3).

aararpufr WW)t(WW)t(W (10.3)

where Wu is the unit water content of mix, Wrp(t) is the water restricted by

powder material at the considered age, Wra is the water restricted on the surface of

aggregates, and Waa is the additional free water due to filling effect of ultra-fine

particles. All parameters are provided in kg/m3 of concrete.

The water restricted by powder material includes water absorbed in powder

particles and water retained on surface of powder particles. The total amount of water

restricted by all powder materials (Wrp) is governed by water retain-ability of powder

materials. It can be obtained by Eq. (10.4). Water retain-ability of powder (βi)

depends on many factors such as porosity, surface condition, shape, size distribution

and loss of ignition of each powder. The β for cement and fly ash can be computed by

Eq. (10.5) and Eq. (10.6), respectively. It was stated that water retain-ability of

powder increases with the increase of environmental temperature so that this effect is

considered as well (Wangchuk, 2003).

n

1i

pipirp w*W (10.4)

49.0

c

55.0

c

74.0

pc /S**004.0 (10.5)

34.0

f

74.0

f

16.0

fpf /)LOI98.2(S*028.0 (10.6)

92

where βpi is the water retain-ability of the ith

powder material, and wpi is the

absolutely dried weight of the ith

powder material (kg/m3 of concrete). βpc and βpf are

the water retain-ability coefficients of cement and fly ash (g/g of dried weight), c and

f are the specific gravities of cement and fly ash, Sc and Sf are specific surface areas

of cement and fly ash (cm2/g), is the angularity factor of cement, and LOIf is the

loss on ignition of fly ash (%). The effects of time and temperature are added as

functions and were described previously in details by Wangchuk (2003).

The amount of water restricted at the surface of aggregates (Wra) can be

computed from Eq. (10.7). In this case, the water retain-ability of fine and coarse

aggregates are considered to be affected by specific surface areas, shape of the

particles, dimensions, and the specific gravities of the fine and coarse aggregates.

ggssra w*w*W (10.7)

where βs and βg are the surface water retain-ability coefficients of fine and

coarse aggregates excluding absorption (g/g of SSD aggregate), ws and wg are the

saturated surface-dry (SSD) weights of the fine and coarse aggregates (kg/m3 of mix),

respectively.

It was reported that using fine powders as partial replacements of cement

causes a filling effect on the amount of free water in concrete (Khunthongkeaw,

2001). Finer particles can fill in the voids amongst cement particles resulting in

moving out water entrapped in these voids. Hence, the amount of free water increases,

so reducing the water requirement of concrete to a certain extent. The additional

water due to filling effect of fly ash (Waa) was previously modeled in Eq. (10.8).

wfillaa VW * (10.8)

93

where Vfill is the volume of the fillable particles in the voids among cement

(Vc, m3 of concrete), and w is the specific gravity of water. A filling coefficient is

computed assuming that higher cement content results in larger amount of voids for

being filled by finer powder. This coefficient depends on the shape of particles,

specific surface areas of cement and fly ash, and fly ash replacement level. More

details are shown by Wangchuk (2003).

The minimum free water content required to initiate slump (Wo) is computed

by Eq. (10.9). The inter–particle surface forces were reported to vary significantly

with a certain amount of surface area of solid particles which was previously defined

as the effective surface area (Wangchuk, 2003). Moreover, the spherical particles can

reduce the inter-particle friction among larger particles; hence, a lubrication

coefficient was introduced as well. Therefore, the minimum free water content

required to initiate slump is modeled as a function of both effective surface area of

solid particles and lubrication coefficient.

L

SW

eff

o

76.05 *10*8

(10.9)

where Seff is the effective surface area of solid particles (cm2/m

3 of concrete)

and L is the lubrication coefficient to account for the lubrication effect of spherical

shape powder particles. Detailed explanation is provided by Wangchuk (2003).

After computing the water restricted by powder materials, water restricted by

the aggregates, filling effect by fine powder, free water content can be computed.

Then initial slump can be computed using Eq. (10.1) with substitutions from Eq.

(10.2) to Eq. (10.9).

In this study, initial slump was measured for concrete containing different

replacements of fly ash and slag with different water to binder ratios. The water

retain-ability coeffificents of powder materials were obtained through experiments.

Then the results were comparatively analyzed in order to identify the behavior of

94

initial slump of concrete mixes with fly ash and slag. The tested values of initial

slump of concrete with and without fly ash were verified using the model proposed by

Wangchuk (2003).

10.2 Experimental Procedure







were used in saturated surface-dry conditions. Normal tap water was used as mixing

water in all mixes. A dosage of 1.5% of a napthalene-based superplasticizer was used

for all concrete mixes with water to binder ratio of 0.4, in order to enhance

workability.

Water retain-ability of powder materials were tested using mini-slump test. A

simple method is adopted from Kitticharoenkiat’s study (Kitticharoenkiat, 2000). It

was measured by finding the lowest water to binder ratio (w/b) that initiates slump of

the paste using a mini-slump cone test. The experiment is started using a low w/b ratio

in which the slump value is zero. Then step by step increasing of the w/b was done

until the slump was initiated. A metal cone shaped mold (40±3 mm inside top

diameter, 90±3 mm inside bottom diameter, and 75±3 mm height) was used in which

the paste was cast in three layers. Each layer was compacted using a metal tamper

weighing 340±15 g. This method was used to find water retainability of slag, cement,

and blended cement with 50% low calcium fly ash.

Initial slump of concrete was measured in concrete containing different

replacements of fly ash and slag with different water to binder ratios. The mixes are

similar to that of the semi-adiabatic temperature measurement previously described in

Chapter 9. The initial slump was measured in accordance with ASTM C143.

95


The water retain-ability of cement, slag, and blended cement with 50% low

calcium fly ash were found to be 0.235, 0.25, and 0.225, respectively. The high water

retain-ability of slag may be due to the high fineness and irregular shape of its

particles which can increase inter-particle friction.

The initial slump of different mix designs are shown in Fig. 10.1 and Fig.

10.2. All the mixes at w/b ratio of 0.40 show higher initial slump than that of the

mixes with w/b of 0.5, due to the addition of 1.5% naphthalene based superplasticizer.

It is well known that the long molecules of superplasticizers have the ability to wrap

themselves around the powder materials and provide highly negative charge. Hence,

the powder materials attempt to repel each other which results in dispersion and

lowering the inter-particle friction forces.

The effect of w/b can be seen when comparing similar mixes with w/b of 0.5

and 0.6 in both Fig. 10.1 and Fig. 10.2. It is seen that as w/b ratio increases, the initial

slump increases due to increased lubrication by additional free water amount.

The effect of fly ash is shown in Fig. 10.1. It is observed that mixes with 50%

fly ash resulted in highest initial slump when compared to all other mixes. This is due

to the spherical particle shape of fly ash which can assist in reducing the inter-particle

friction forces, reducing adherence of cement particles and the ability of reducing the

water requirement due to filling effect as well. Thus, initial slump increases as the fly

ash replacement increases.

The effect of slag is shown in Fig. 10.2. It is observed for all tested w/b ratios

that the initial slump slightly reduces as the slag replacement increases. This result

may be due to high water retain-ability of slag particles as mentioned previously.

High temperature may also affect increasing water retain-ability of slag, therefore,

reducing the amount of free water for initiating the slump. Moreover, the irregular

shaped particles of slag can interlock each other firmly resulting in increasing inter

96

particle friction forces, and reduce dispersion of cement particles as the replacement

of slag is higher.

Fig 10.1 Tested results of initial slump of concrete with fly ash 0%, 30%, and 50% at

w/b=0.4, 0.5, and 0.6

Fig 10.2 Tested results of initial slump of concrete with slag 0%, 50%, 65%, and 75%

at w/b=0.4, 0.5, and 0.6

97

10.4 Verification of the Initial Slump Model for Concrete

The model developed by Wangchuk (2003) is used to predict initial slump of

concrete with and without fly ash. The parameters used for the cement, fly ash, and

slag to verify the initial slump model are summarized in Appendix D (Table D1). A

water reducing efficiency of 0.35 was used as an input in the model for the

napthalene-based superplasticizer used in this experiment (Wangchuk, 2003). The

comparison between tested and predicted values of initial slump for concrete with and

without fly ash is illustrated in Fig. 10.3. It is observed that the existing model is

capable of predicting initial slump with an accuracy in the range of ±3cm.

In the case of concrete containing slag, the comparison between tested and

predicted values of initial slump is illustrated in Fig. 10.4. The existing model is

capable of predicting initial slump with an accuracy in the range of ±2cm for fly ash

concrete cases and ±3cm for slag concrete cases. Slag concrete shows lower accuracy

of the predictions due to inaccuracies in computing the amount of additional water

due to filling effect from finer particles of slag. Existing model predicts the filling

effect from finer particles considering the limits of fly ash replacement levels. Thus, it

is required to modify the model considering the effect of slag replacement level.

Moreover, the effects on initial slump from loss on ignition and temperature are

different in the case of slag when compared to fly ash. Therefore, it is recommended

to further investigate the initial slump and slump models for the case of slag concrete.

98

Fig. 10.3 Comparison of tested and predicted initial slump values for concrete

containing fly ash 0%, 30%, and 50% for all tested w/b

Fig. 10.4 Comparison of tested and predicted initial slump values for concrete

containing slag 50%, 65%, and 75% for all tested w/b

99

Chapter 11

Compressive Strength

11.1 General

Significant amount of literature states the fact that partial replacement of

cement with mineral admixtures such as fly ash or slag, results in lower/delay

compressive strength development at early stages (Demirboga, 2003; Demirboga et

al., 2007; Uysal & Akyunsu, 2012). However, it is seen that higher later strength is

archived due to production of C-S-H at the expense of Ca(OH)2 by the pozzolanic

reaction. The effect of fly ash on compressive strength is different from that of the

effect of slag, due to different physical and chemical properties.

Numerous attempts have been undergone to model the compressive strength

of concrete containing fly ash (Kaewkhluab, 2002; Hung, 2005). Kaewkhluab (2002)

proposed a model to predict compressive strength at 28 days including several factors

which is shown in Eq. (11.1).

airwrLOIfeffc CaOdaysf ****]*)log([)28( 21

' (11.1)

where f’c (28 days) represents 28-day compressive strength of concrete cured

at normal temperature of 30°C. CaOeff is the amount of effective CaO in concrete

which depends on amount of CaO content of binders and fineness of fly ash. α1, α2 are

factors to include the effect of w/b on the compressive strength. λf, xγ, xLOI, xwr, xair

are the factors proposed for the effects of particle packing of fly ash, paste content,

loss on ignition of fly ash, water reducing admixture and air content, respectively. The

model was developed to compute compressive strength until the age of 1 year with the

considerations of hydration degrees. Hung (2005), further modified the previously

proposed model for different types of concretes and various curing temperatures.

In this study, compressive strength tests were conducted for concrete

containing different replacements of fly ash and slag at different water to binder

100

ratios. Then the results were comparatively analyzed in order to identify the behavior

of compressive strength with fly ash and slag. The compressive strength results at 28

days were verified using the model proposed by Kaewkhluab (2002). Simulating

compressive strength of concrete containing slag was done by using similar model,

although it is recommended to modify the proposed model considering the different

behavior of slag when compared to that of fly ash.

11.2 Experimental Procedure







were used in saturated surface-dry condition. Normal tap water was used as mixing

water in all mixes. A dosage of 1.5% of a napthalene-based superplasticizer was used

for all concrete mixes with water to binder ratio of 0.4, in order to enhance

workability.

This experiment was conducted to investigate the compressive strength

development of concrete samples containing different replacements of fly ash and

slag with different water to binder ratios. Water to binder ratios (w/b) of 0.40, 0.50

and 0.60 were used. Slag replacements of 50%, 65%, and 75% and fly ash

replacements of 30% and 50% were utilized. The mix designs of the concrete used in

this study are similar to that of the semi-adiabatic test which were previously shown

in Table 9.1.

Cube specimens with dimensions of 10×10×10cm were cast to obtain

compressive strength results. Water curing was done until the specimens were tested.

Three specimens were tested per each mix and average compressive strength was

computed. The measurements were taken at 3, 7, 28, and 91 days after casting.

101


The compressive strength development of the tested specimens having

different mix designs at 3, 7, 28, and 91 days are shown from Fig. 11.1 to Fig. 11.6.

The effect of age is evident where the compressive strength of all mixes increases

with age. The effect of w/b ratios and fly ash/slag replacement levels are described

below.

11.3.1 Effect of w/b

For concrete containing fly ash, Fig. 11.1, Fig. 11.2, and Fig. 11.3 show the

effect of w/b ratios whereas that for concrete containing slag are shown in Fig. 11.4,

Fig.11.5, and Fig. 11.6. It can be seen that for both mineral admixtures, the

compressive strength of concrete having similar replacement levels clearly reduces as

the w/b increases from 0.4 to 0.6. This is due to the fact that w/b directly affects

microstructure formation by affecting the mean inter-particle distance and volume of

capillary porosity (Hung, 2005). Thus higher w/b results in increasing the total

porosity in the paste mixes. It leads to reduction of the denseness of structure,

resulting in decreasing compressive strength of concrete.

11.3.2 Effect of fly ash

The effect of fly ash for different w/b is shown in Fig. 11.1, Fig. 11.2, and

Fig.11.3. It is observed that for mixes having similar w/b, compressive strength

reduces significantly as the replacement ratio of fly ash increases. Moreover, the

compressive strength is lower than that of cement-only concrete. This tendency is

evident up to 28 days. The compressive strength development ratio values are

computed and summarized in Fig.11.7 for all concrete mixes with w/b of 0.5. It is

defined as the ratio of compressive strength of concrete at a given age to its 28-day

compressive strength. From the compressive strength development ratios of concrete

containing fly ash, it is observed that the ratio at 91 days is higher as the replacement

increases from 0% to 50%. This reveals that the compressive strength gain rate of

concrete with fly ash is higher than that of cement-only concrete at later ages. This is

102

due to the pozzolanic reaction which is weak at early stage after casting of concrete.

After certain period of hydration process, pozzolanic reaction is activated and then the

reactivity of fly ash improves with age. It was previously reported that the period

which pozzolanic reaction starts rapidly, delays as the calcium amount is lower

(Neville, 1995). In this study a low calcium fly ash is utilized thus the strength

development ratio is higher than normal concrete at later age.

Fig 11.1 Test results of compressive strength of concrete with fly ash 0%, 30%, and

50% at w/b=0.40

103


50% at w/b=0.50


50% at w/b=0.60

104

11.3.3 Effect of slag

The effect of slag on compressive strength of concrete with different w/b is

shown in Fig. 11.4, Fig. 11.5, and Fig.11.6. It is observed that for mixes having

similar w/b, compressive strength reduces significantly as the replacement ratio of

slag increases. Moreover, the compressive strength is lower in slag concrete than that

of cement-only concrete at early age. However, in the case of w/b ratio of 0.4, the

early age results are slightly higher at slag replacements of 50% and 65%. This may

be because the addition of slag up to certain limit at low w/b ratios can result in

denser interfacial transition zone (ITZ), optimized structure and distribution of pores

(Shui et al., 2013). It was also reported that slag can reduce the amount and mean size

of Ca(OH)2 crystals, resulting in dense structure with high strength (Gao et al., 2005).

However, it is recommended to investigate the strength development behavior of slag

concrete at early age in order to understand the mechanism clearly.

It is seen that strength development ratios of slag concrete is higher than that

of fly ash concrete up to 28 days. This is due to the fact that slag is capable of reacting

itself with water in a slow rate and actively continues its pozzolanic reaction with

Ca(OH)2 which is produced from cement hydration as well (Wang & Linger, 2010).

Therefore, initial strength gaining rate is higher than that of the fly ash concrete.

However, as the replacement of slag increases, the reactivity becomes lower.

The results of a previous study by Oner & Akyuz (2007) indicated that highest

efficiency and strength of concrete was obtained when the slag content is not too high,

especially up to 60% of the total amount of binding materials. After a certain limit of

slag replacement, it is not efficient as a binder; instead, it will become as filler in

concrete. It was reported that slag could not enter and contribute to the chemical

reactions if the replacement is too high. However, this depends on other parameters

such as physical and chemical properties of slag and mix proportions as well.

105

It is reasonable to compare the effect of fly ash and slag by using compressive

strength development ratios which interpret the ratio of compressive strength of

concrete at a given time to its 28 day compressive strength. The comparison of

strength development ratios of the tested mixes are shown in Fig. 11.7. The ratios of

both fly ash and slag are lower than that of cement-only concrete up to 28 days. It is

seen that slag concrete can gain higher strength rate at early stage than that of the fly

ash concrete. However, in terms of later age strength rate, fly ash concrete is higher

than that of slag concrete since the pozzolanic reaction of fly ash continues further at

later age as well.

Fig 11.4 Test results of compressive strength of concrete with slag 0%, 50%, 65%,

and 75% at w/b=0.40

106


and 75% at w/b=0.50


and 75% at w/b=0.60

107

Fig 11.7 Comparison of strength development ratios of concrete with fly ash 0%,

30%, 50%, and concrete with slag 50%, 65%, 75% at w/b=0.50

11.4 Verification of the 28-day Compressive Strength Model

Kaewkhluab (2002) proposed a model to predict compressive strength of fly

ash concrete at 28 days including several factors. It is shown previously in Eq. (11.1).

The 28-day compressive strength values obtained by experimental procedure are

verified using this equation. The parameters used for the cement, fly ash, and slag to

verify the 28-day compressive strength model are summarized in Appendix D (Table

D2). A water reducing efficiency of 0.35 was used as an input in the model for the

napthalene-based superplasticizer used in this experiment (Wangchuk, 2003). The

comparison between tested and predicted values of compressive strength for fly ash

concrete and slag concrete are illustrated in Fig. 11.8 and Fig. 11.9, respectively.

It is observed that the existing model is capable of predicting 28-day

compressive strength with sufficient accuracy. The predictions are verified with a

possible error of ±20%. However, mostly the predictions are lower than that of the

test results. This may be due to the improvement of cement properties recently.

Recent cements have slightly higher strength than the cements produced earlier

108

probably due to addition of some compounds into the cement. In the case of concrete

containing slag, the most of the model predictions are lower than that of the test

results as well. The predictions are still satisfactory. However, it is recommended to

further modify the existing model for both cases of fly ash and slag for higher

accuracies.

Fig. 11.8 Comparison of tested and predicted 28-day compressive strength values for

concrete containing fly ash 0%, 30%, and 50% for all tested w/b

109

Fig. 11.9 Comparison of tested and predicted 28-day compressive strength values for

concrete containing slag 50%, 65%, and 75% for all tested w/b

110

Chapter 12

Conclusions and Recommendations for Future Studies

Predicting temperature rise is essentially useful for investigating thermal

cracking potential especially at early stage of mass concrete containing slag.

Therefore this study mainly focused on investigating thermal properties and

temperature rise through experiments which were followed by modeling these

properties based on hydration degrees of cement and slag. The important conclusions

obtained from this study can be summarized as follows.

12.1 Conclusions

1. The hydration degree which represents reaction rate was modelled as a

function of age, water to cement ratio and concrete temperature. Since, the model was

able to predict hydration degree of each reactive compound, the average hydration

degree was computed based on the assumption that each cement compound reacts

independently.

2. The hydration degree of slag in this study was defined as the weight fraction

of already reacted slag per total slag in the concrete mix. The model equation was

developed using the method of back analysis from adiabatic temperature results of

slag concrete. The key parameters of the model for predicting hydration degree of slag

were age, water to binder ratio, concrete temperature, slag replacement ratio and

fineness of the slag. The dispersion effect was considered to affect hydration degree

of C3S and C3A in cement at very early age. Thus, equations for dispersion effect

were proposed for acceleration of C3S and C3A hydrations.

3. The free water content of paste specimens clearly reduced with age due to

the consumption of water by cement and slag hydration. Free water increased as the

slag substitution level was higher. This was due to lower reactivity of slag at early age

when compared to cement. Simulating free water content in pastes with slag was

111

achieved by adopting a time-dependent model which computes free water based on

chemically bound water and gel water amounts. The model simulations showed

sufficient accuracy in predicting the test results of paste specimens at different water

to binder ratios and slag replacement levels.

4. The specific heat of pastes containing slag was experimentally measured

using transient method. It was concluded that the specific heat is a time-dependent

property which is significantly affected by the availability of free water at the

considered age. Time-dependent model was proposed based on weight fraction and

specific heat of each constituent. The model simulations showed sufficient accuracy

in predicting the tested results.

5. Transient method provided the outputs for thermal conductivity of pastes

containing slag. An increase of slag replacement level resulted in a slight decrease of

the values due to the lower thermal conductivity of slag when compared to cement.

The thermal conductivity model was proposed based on volume fraction and thermal

conductivity of each constituent. The model indicated a certain level of accuracy of

predictions.

6. CTE of pastes was experimentally obtained by measuring change in length

for every 5°C change of temperature. The CTE of pastes decreased as the slag

replacement level increased which was partly due to lower CTE of slag itself. The

model of CTE of paste at a given time was proposed based on volume fractions of

non-reacted binder content and hydrated product amounts. The model simulations

showed sufficient accuracy in predicting the test results.

7. Since the hydration degrees of cement and slag, free water content, specific

heat and total heat generation of concrete were separately modeled, the models can be

linked to compute the adiabatic temperature rise. Verifications of the proposed

adiabatic temperature rise model were done by using different sources of previous test

results. The model simulations were done with a reasonable accuracy of the predicted

results.

112

8. The semi-adiabatic temperature profiles indicated that fly ash concrete

performs better than that of slag concrete in terms of reducing heating up slopes. This

is due to low heat generation, low thermal conductivity, and high specific heat of fly

ash concrete when compared to the slag concrete, especially at early stage.

9. Initial slump increased as the fly ash replacement increased because the

spherical particle shape of fly ash can assist in reducing the inter-particle friction

forces and reduce the water requirement. Slag replacements caused slightly reductions

of the initial slump as the replacement increased. This may be due to high water

retain-ability of slag particles. A model for initial slump, computed based on free

water amount, was verified for both cases of fly ash and slag concrete and it was able

to accurately predict the test results.

10. The compressive strength results indicated that slag concrete possess

higher compressive strength than that of the fly ash concrete. However, strength

development ratios proved that fly ash concrete performed better at later ages. The 28-

day compressive strength model was verified for both fly ash and slag concrete and it

was able to accurately predict the tested values.

113

12. 2 Recommendations for Future Study

1. Effect of chemical composition of slag should be included in modeling

hydration degree of slag. Thereby, different types and sources of slag can be verified

for their hydration degrees.

2. Slag possesses both self-hydration and pozzolanic reactions. The model

equation for hydration degree of slag should be further divided into two sections to

separate these reactions.

3. The properties and behavior of hydrated products can be thoroughly

investigated with different mix designs of paste, mortar, and concrete. Hence, the

accuracy of the models can be further enhanced by understanding the microstructural

behavior of the hydrated products.

4. The maximum heat generation of cement compounds and slag can be varied

with different sources of the cement and slag and may depend on the properties of the

binders. Hence it is required to investigate the maximum heat generation of different

types of binders.

5. The effect of super plasticizers and retarding agents was not included in this

study. Hence, it can be further investigated.

6. The proposed adiabatic temperature simulation model is capable of

predicting the hydration degree within the limits of slag replacement from 0.40 to

0.75, water to binder ratio from 0.35 to 0.65, and slag fineness from 3500 cm2/g to

6000 cm2/g. The prediction range of these parameters can be further enhanced.

7. Finite element analysis shall be applied to simulate the tested semi-adiabatic

temperature profiles in order to investigate heat transfer considering the surrounding

environment.

8. The proposed model can be applied in actual construction to investigate the

temperature of slag concrete. Then, thermal cracking potential can be evaluated by

comparing the restrained strain to tensile strain capacity at any point in mass concrete

structures.

114

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Appendices

123

Appendix A

Mix proportions, physical and chemical properties of the binders

used in adiabatic temperature rise investigations by various previous

studies which used to verify the proposed model

Table A1 Mix proportions of concrete used in adiabatic temperature rise tests by

previous studies

Source

Mix proportions (kg/m3)

Cement Slag Water Sand Gravel w/b Slag ratio

Wang & Linger,

2010: Mix 1

1

240

1

160

1

157

6

645

1

1129

0

0.392

0.40

Wang & Linger,

2010: Mix 2

1

180

1

120

1

148

6

757

1

1129

0

0.493

0.40

Taiheiyo Cement

Corp. (testing of

Korean slag, 2014)

1

115

1

215

1

165

6

772

1

1079

0

0.50

0.65

Taiheiyo Cement

Corp. (testing in

Singpore, 2014)

1

91

1

259

1

160

6

845

1

980

0

0.46

0.74

124

Table A2 Mix proportions of concrete used in adiabatic temperature rise tests by Tada

et al. (2014)

Source

Mix proportions (kg/m3)

Cement Slag Water Sand Gravel w/b Slag ratio

s=0.00, w/b=0.40 400 0 160 798 1000 0.40 0.00

s=0.00, w/b=0.65 254 0 165 834 1094 0.65 0.00

s=0.50, w/b=0.40 206 207 165 736 1047 0.40 0.50

s=0.50, w/b=0.60 137 138 165 897 1001 0.60 0.50

s=0.65, w/b=0.40 145 268 165 734 1044 0.40 0.65

s=0.65, w/b=0.60 96 179 165 858 1038 0.60 0.65

s=0.75, w/b=0.40 103 310 165 732 1041 0.40 0.75

s=0.75, w/b=0.60 69 206 165 895 995 0.60 0.75

Table A3 Physical properties of binders used in adiabatic temperature rise tests by

previous studies

Source Binder type Blaine fineness

(cm2/g)

Specific gravity

(g/cm3) Ig. Loss (%)

Wang & Linger,

2010

Cement I 3380 3.15 2

Slag 4250 2.90 0.8

Taiheiyo Cement

Corp. (testing of

Korean slag,

2014)

Cement I 3320 3.15 1.7

Slag 4070 2.98 1.09

Taiheiyo Cement

Corp. (testing in

Singpore, 2014)

74% slag

blended

cement

3560 3.10 0.67

Tada et al. (2014) Cement I 3320 3.16 1.7

Slag 4220 2.89 0.96

125

Table A4 Chemical properties of binders used in adiabatic temperature rise tests by

previous studies

Source Binder

type

Chemical composition (% by weight)


Wang & Linger,

2010

Cement I -

Slag 33.4 15 0.5 43.1 6.6 0

Taiheiyo

Cement Corp.

(testing of

Korean slag,

2014)

Cement I

20.65 5.66 2.93 63.10 2.50 2.07

Slag 3

32.72

1

13.14

0

0.40

4

44.24

5

5.09

2

2.10

Taiheiyo

Cement Corp.

(testing in

Singpore, 2014)

74% slag

blended

cement

2

27.05

1

11.24

1

1.51

5

50

5

5.55

2

2.95

Tada et al.

(2014)

Cement I 20.65 5.66 2.93 63.10 2.50 2.07

Slag 34.54 14.06 0.29 44.38 5.59 0

126

Appendix B

Physical and chemical properties of fly ash and mix proportions of

concrete used in semi-adiabatic temperature rise, initial slump, and

compressive strength experimental investigations

Table B1 Physical properties of fly ash

Physical properties Fly ash

Specific gravity 2.27

Fineness (cm2/g) 3490

Loss on ignition (%) 1.9

Table B2 Chemical properties of fly ash

Chemical

properties

%


Fly ash 66.8 22.6 3.1 1.1 0.4 0.1

127

Table B3 Mix proportions of the concrete cast for measuring semi-adiabatic

temperature, initial slump, and compressive strength

Mix Cement

(kg/m3)

Fly ash

(kg/m3)

Slag

(kg/m3)

Water

(kg/m3)

Sand

(kg/m3)

Gravel

(kg/m3)

W40 S/R00 421.6 - - 168.6 798.7 1029.2

W40 R30 280.6 120.2 - 160.3 798.7 1029.2

W40 R50 194.0 194.0 - 155.2 798.7 1029.2

W40 S50 206.5 - 206.5 165.2 798.7 1029.2

W40 S65 143.7 - 266.9 164.2 798.7 1029.2

W40 S75 102.2 - 306.7 163.6 798.7 1029.2

W50 S/R00 369.9 - - 184.9 798.7 1029.2

W50 R30 247.7 106.2 - 176.9 798.7 1029.2

W50 R50 171.9 171.9 - 171.9 798.7 1029.2

W50 S50 181.7 - 181.7 181.7 798.7 1029.2

W50 S65 126.5 - 234.9 180.7 798.7 1029.2

W50 S75 90.1 - 270.14 180.1 798.7 1029.2

W60 S/R00 329.6 - - 197.7 798.7 1029.2

W60 R30 221.7 95.0 - 190.0 798.7 1029.2

W60 R50 154.4 154.4 - 185.2 798.7 1029.2

W60 S50 162.2 - 162.2 194.6 798.7 1029.2

W60 S65 112.9 - 209.8 193.7 798.7 1029.2

W60 S75 80.5 - 241.4 193.0 798.7 1029.2

128

Appendix C

Analysis of the tested semi-adiabatic temperature profile

characteristics

Table C1 Initial temperature, peak temperature, time to attain peak temperature, and

heating and cooling slopes of all tested concrete mixes for semi-adiabatic temperature

Mix

Initial

temp.

(°C)

Peak

temp.

(°C)

Time to

attain peak

temp. (°C)

Temp at 90

hours (°C)

Heating

slope

(°C/day)

Cooling

slope

(°C/day)

W40 S/R00 31.5 71.0 16.7 40.1 2.4 0.4

W40 R30 32.1 61.6 19.2 40.3 1.5 0.3

W40 R50 31.9 51.5 23.8 39.5 0.8 0.2

W40 S50 31.1 65.5 17.5 39.7 2.0 0.4

W40 S75 31.6 56.1 15.2 36.9 1.6 0.3

W50 S/R00 30.7 66.3 16.8 40.2 2.1 0.4

W50 R30 30.3 56.0 19.5 40.3 1.3 0.2

W50 R50 29.6 47.5 24.8 38.3 0.7 0.1

W50 S50 30.4 60.2 19.0 41.2 1.6 0.3

W50 S75 31.1 54.5 15.8 37.1 1.5 0.2

W60 S/R00 30.5 61.0 17.2 40.6 1.8 0.3

W60 R30 30.7 51.5 22.5 40.0 0.9 0.2

W60 R50 31.3 46.9 26.3 38.0 0.6 0.1

W60 S50 31.0 57.2 21.3 41.5 1.2 0.2

W60 S75 31.3 51.7 20.2 37.6 1.0 0.2

129

Appendix D

Parameters used for the cement, fly ash, and slag to verify the initial

slump and 28-day compressive strength models

Table D1 Parameters used for the cement, fly ash, and slag to verify the initial slump

model

Binder Cement Fly ash Slag

Specific gravity 3.16 2.27 2.89

Fineness (cm2/g) 3570 3490 4330

Loss on ignition (%) 2.06 1.9 0.96

Water retain-ability 0.225 0.19 0.25

Angularity factor (ψ) 1.4 1.2 1.55

Table D2 Parameters used for the cement, fly ash, and slag to verify the 28-day

compressive strength model

Binder Cement Fly ash Slag

Specific gravity 3.16 2.27 2.89

Fineness (cm2/g) 3570 3490 4330

Loss on ignition (%) 2.06 1.9 0.96

CaO (%) 63.4 1.1 43.26

Angularity factor (ψ) 1.4 1.2 1.55

models for predicting thermal properties and temperature...

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