Chloride-Induced Steel Corrosion in Concrete Under Service Loads 9811541078, 9789811541070

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Hailong Ye Chuanqing Fu Ye Tian Nanguo Jin

Chloride-Induced Steel Corrosion in Concrete Under Service Loads

Chloride-Induced Steel Corrosion in Concrete Under Service Loads

Hailong Ye Chuanqing Fu Ye Tian Nanguo Jin •

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•

Chloride-Induced Steel Corrosion in Concrete Under Service Loads

123

Hailong Ye The University of Hong Kong Hong Kong, China

Chuanqing Fu Zhejiang University of Technology Hangzhou, China

Ye Tian Zhejiang University Hangzhou, China

Nanguo Jin Zhejiang University Hangzhou, China

ISBN 978-981-15-4107-0 ISBN 978-981-15-4108-7 https://doi.org/10.1007/978-981-15-4108-7

(eBook)

© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2020 This work is subject to copyright. All rights are solely and exclusively licensed by the Publisher, whether the whole or part of the material is concerned, specifically the rights of translation, reprinting, reuse of illustrations, recitation, broadcasting, reproduction on microfilms or in any other physical way, and transmission or information storage and retrieval, electronic adaptation, computer software, or by similar or dissimilar methodology now known or hereafter developed. The use of general descriptive names, registered names, trademarks, service marks, etc. in this publication does not imply, even in the absence of a specific statement, that such names are exempt from the relevant protective laws and regulations and therefore free for general use. The publisher, the authors and the editors are safe to assume that the advice and information in this book are believed to be true and accurate at the date of publication. Neither the publisher nor the authors or the editors give a warranty, expressed or implied, with respect to the material contained herein or for any errors or omissions that may have been made. The publisher remains neutral with regard to jurisdictional claims in published maps and institutional affiliations. This Springer imprint is published by the registered company Springer Nature Singapore Pte Ltd. The registered company address is: 152 Beach Road, #21-01/04 Gateway East, Singapore 189721, Singapore

Preface

Concrete is the most economical and suitable material for the construction of the vast variety of marine structures in use nowadays, such as super-span bridges, undersea tunnels, and offshore drilling platforms. However, the hostile and highly complex marine environment greatly accelerates the deterioration of reinforced concrete structures, and thus its durability performance in the marine environment has been an active area of research in recent decades. Among the many detrimental factors, the chloride-induced corrosion of steel is the most fundamental cause of performance degradation of marine structures in China and worldwide. Such corrosion damage of reinforcement in concrete infrastructure has led to a number of structural failures in recent decades. Also, this progressive degradation of antique residential buildings poses a threat to the safety of the general public and a great challenge for structural maintenance. The total cost of corrosion in China is approximately 3% of the annual national gross domestic product, about 10% of which is related to civil infrastructure. This book summarizes and outlines the main findings obtained in the decade studies conducted at the School of Civil Engineering, Zhejiang University, China, with a specific focus on the impact of service loads on chloride-induced steel corrosion in concrete structures. The coupling effects of mechanical loads and reinforcement corrosion on the degradation of concrete structures are highlighted. This book includes the impact of service loads of various forms (e.g., flexural loads, fatigue loads) on the transport properties of concrete, the influence of environmental conditions (e.g., salt-fog, carbonation) on the chloride ingress in concrete, natural and accelerated corrosion propagation process in concrete beams under service loads, mechanical degradation of concrete beams corroded under service loads, among others. The outcomes of this book highlight the importance and significance of the coupling and combinational effects of mechanical loads and environmental factors on the deterioration process of reinforced concrete structures. The book contains six chapters. In Chap. 1, brief introduction to the characteristics of the marine environments, the physiochemical properties of modern concrete materials, the transport mechanism of chloride and other deleterious agents into concrete materials, and the electrochemical process of steel corrosion in v

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Preface

concrete are given. In Chap. 2, the effects of stress on the moisture transfer and chloride diffusion coefficient of concrete are discussed with the introduction of the loading effect factor. In Chap. 3, the effects of geometrical characteristics of cracks on the moisture influential depth and chloride diffusion coefficient of pre-cracked concrete are discussed. In Chap. 4, the influence of various environmental conditions on the chloride penetration in concrete is discussed, including cyclic seawater wet-dry, salt-fog wet-dry, and carbonation. In Chap. 5, the corrosion propagation characteristics in pre-cracked concrete beams corroded under sustained loading are discussed. In Chap. 6, the mechanical degradation of pre-cracked concrete beams corroded under sustained service loads is discussed. This book is intended for researchers and structural and concrete materials engineers who are interested in having a systematic understanding of the impact of service loads on the various stages of service-life performance of reinforced concrete structures. Hong Kong, China Hangzhou, China Hangzhou, China Hangzhou, China

Hailong Ye Chuanqing Fu Ye Tian Nanguo Jin

Acknowledgments

The financial supports from the National Basic Research Program of China (Grants No. 2015CB655103) and National Natural Science Foundation of China (Grants Nos. 51808475, 51678529, and 51308503) are gratefully acknowledged. The authors would like to thank the colleagues, postdoctoral researchers, and postgraduate students who have worked on those projects over years, including Dr. Qiang Li, Dr. Bei Li, Dr. Zhonggou Chen, Dr. Zhi Wang, Dr. Jun Chen, Mr. Yanxin Huang, Mr. Tao Huang, Mr. Yibin Xu, Mr. Wei Dai, Mr. Mingyue Du, and Miss. Wei Chen. Their hard-working, dedication, and intelligence have contributed considerably to the work presented in this book. Any opinions, findings, and conclusions or recommendations expressed in this book are those of the authors and do not necessarily reflect the views of the sponsors, any institutions, or organizations.

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Contents

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1 1 1 4 5 6 6 7 8

2 Chloride Ingress in Stressed Concrete . . . . . . . . . . . . . . . . . . . . 2.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.2 The Loading Effect Factor . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.2.1 Damage Influence Factor . . . . . . . . . . . . . . . . . . . . . . 2.2.2 Definition of the Damage Variable . . . . . . . . . . . . . . . 2.2.3 Derivation of Loading Effect Factor . . . . . . . . . . . . . . 2.3 Effect of Flexural Loading on Chloride Ingress . . . . . . . . . . . 2.3.1 Experimental Program . . . . . . . . . . . . . . . . . . . . . . . . 2.3.2 Influence of Flexural Load on Saturated Concrete . . . . 2.3.3 Influence of Flexural Load on Non-saturated Concrete . 2.4 Fatigue Loads on Moisture Transport . . . . . . . . . . . . . . . . . . . 2.4.1 Surface Factor . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.4.2 Damage Variable of Fatigue-Damaged Concrete . . . . . 2.4.3 Experimental Programs . . . . . . . . . . . . . . . . . . . . . . . 2.4.4 Moisture Transport in Surface Concrete . . . . . . . . . . . 2.4.5 Influence Factors on the Surface Factor . . . . . . . . . . . . 2.4.6 Surface Mass Transfer Area . . . . . . . . . . . . . . . . . . . .

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11 11 12 12 13 16 18 18 21 23 24 24 24 25 29 31 32

1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.1 Background . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.2 Marine Environment . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.3 Composition and Properties of Concrete . . . . . . . . . . . . 1.4 Supplementary Cementitious Materials . . . . . . . . . . . . . . 1.5 Pore Structure and Transport Properties of Concrete . . . . 1.6 Electrochemical Reaction of Chloride-Induced Corrosion 1.7 Content of This Book . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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2.5 Fatigue Loads on Chloride Ingress . . . . . . . . . . . . . . . 2.5.1 Experimental Program . . . . . . . . . . . . . . . . . . . 2.5.2 Chloride Profiles in Fatigue-Damaged Concrete . 2.5.3 Fatigue Loading Effect Factor . . . . . . . . . . . . . . 2.6 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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3 Chloride Ingress in Cracked Concrete . . . . . . . . . . . . . . . . . . . 3.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2 Moisture Influential Depth . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.1 Experimental Program . . . . . . . . . . . . . . . . . . . . . . . 3.2.2 Cyclic RH Change . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.3 Cyclic Temperature Change . . . . . . . . . . . . . . . . . . . 3.2.4 Fog Environment . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.5 Rainfall Environment . . . . . . . . . . . . . . . . . . . . . . . . 3.3 Chloride Diffusion in Cracked Concrete in Saturated States . 3.3.1 Experimental Program . . . . . . . . . . . . . . . . . . . . . . . 3.3.2 Chloride Diffusion Coefficient . . . . . . . . . . . . . . . . . 3.4 Chloride Penetration into Cracked Concrete Under Wet-Dry Cycles . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.4.1 Experimental Program . . . . . . . . . . . . . . . . . . . . . . . 3.4.2 Chloride Profiles in Cracks . . . . . . . . . . . . . . . . . . . . 3.5 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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63 68 68 72 72 72 76 80 80 83 84

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4 Influence of Environmental Condition on Chloride Ingress into Loaded Concrete . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2 Cyclic Wet-Dry Salt-Fog Environment . . . . . . . . . . . . . . . . . . 4.2.1 Experimental Program . . . . . . . . . . . . . . . . . . . . . . . . 4.2.2 Internal RH Evolution in Concrete Under Salt-Fog Environments . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2.3 Influence of Exposure Duration on Chloride Profiles . . 4.2.4 Influence of Binder Type on Chloride Profiles . . . . . . . 4.2.5 Influence of Exposure Orientation on Chloride Profiles 4.3 Cyclic Carbonation and Wet-Dry Condition . . . . . . . . . . . . . . 4.3.1 Experimental Program . . . . . . . . . . . . . . . . . . . . . . . . 4.3.2 Influence of Carbonation on Chloride Profiles . . . . . . . 4.3.3 Influence of Sustained Loading . . . . . . . . . . . . . . . . . . 4.4 Interaction Between Carbonation and Chloride Attack . . . . . . 4.5 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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5 Corrosion Development in Concrete Beams Under Service Loads . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.2 Experimental Program . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.3 Crack Mapping . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.4 Optical Microscopy Examination . . . . . . . . . . . . . . . . . . . . 5.4.1 3D Laser Scanning . . . . . . . . . . . . . . . . . . . . . . . . 5.5 EMPA . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.6 The Effect of Cover Thickness on Corrosion Characteristics 5.6.1 Experimental Program . . . . . . . . . . . . . . . . . . . . . . 5.6.2 Rust Distribution . . . . . . . . . . . . . . . . . . . . . . . . . . 5.6.3 Parameter Analysis . . . . . . . . . . . . . . . . . . . . . . . . 5.6.4 Correlation Between q and Crack Width . . . . . . . . . 5.7 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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85 85 85 89 91 91 96 101 101 104 108 109 112 113

6 Mechanical Degradation of Concrete Beams Corroded Under Service Loads . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.2 Experimental Program . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3 Cracking Map . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.4 Distribution of Mass Loss of Corroded Rebar . . . . . . . . . . 6.5 Loads-Deflection Responses . . . . . . . . . . . . . . . . . . . . . . . 6.6 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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115 115 116 119 127 130 138 138

Abbreviations

3D ASTM EDX EPMA FA GGBS HPC HRB ITZ MIP OPC PC RC RH SCMs SF w/b w/c

3 Dimensions American Society for Testing and Materials Energy-dispersive X-ray spectroscopy Electron probe microanalysis Pulverized fly ash Ground granulated blast-furnace slag High-performance concrete Hot-rolled ribbed bar Interfacial transition zone Mercury intrusion porosimetry Ordinary Portland cement Portland cement Reinforced concrete Relative humidity Supplementary cementitious materials Silica fume Water-to-binder ratio Water-to-cement ratio

Cement Chemistry Notation C S A M S C H

CaO SiO2 Al2O3 MgO SO3 CO2 H 2O

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C–S–H AFt AFm

Abbreviations

Calcium–silicate–hydrate A group of tri-substituted calcium aluminate-ferrite hydrates A group of mono-substituted calcium aluminate-ferrite hydrates

List of Figures

Fig. 1.1

Fig. 1.2 Fig. 2.1 Fig. 2.2 Fig. 2.3

Fig. 2.4

Fig. 2.5

Fig. 2.6 Fig. 2.7 Fig. 2.8

Fig. 2.9 Fig. 2.10

Illustration of multiple physical and chemical deterioration actions on concrete structures (modified from Alexander 2016) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Illustration of the electrochemical corrosion reactions at the steel-concrete interface . . . . . . . . . . . . . . . . . . . . . . . . . Relationship between damage variable d and f ðdÞ . . . . . . . . . a Correlation between tortuosity and porosity; b Correlation between constrictivity and peak radius (unit in m) . . . . . . . . . MIP Pore size distribution of damaged concrete; a Pore size distribution of damaged concrete with different uniaxial tensile strain levels; b Pore size distribution of damaged concrete with different uniaxial compressive strain levels . . . . Layout of the reinforced concrete beam. The thickness of concrete covers was 15 mm at both the top and bottom of the cross section of the beam (unit: mm) . . . . . . . . . . . . . . Four-point bending systems of the RC beam to initiate the transverse cracks and sustain the load during the entire experiment . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Fitted line for the spring calibration in the loading system . . . Powder drilling in beams; a Maps of drilling positions in a beam; b Beams after drilling . . . . . . . . . . . . . . . . . . . . . . Chloride distribution in tensile and compressive zones of pure bending (midspan) section of the saturated; a OPC concrete; b GGBS concrete . . . . . . . . . . . . . . . . . . . . . Chloride diffusion coefficient in different concrete specimens . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . The relationship between chloride diffusion coefficient and strain level of stressed concrete; a Tensile strain; b Compressive strain. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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List of Figures

Fig. 2.11

Fig. 2.12 Fig. Fig. Fig. Fig.

2.13 2.14 2.15 2.16

Fig. 2.17

Fig. 2.18

Fig. 2.19 Fig. 2.20 Fig. 2.21 Fig. 2.22

Fig. 2.23

Fig. 2.24

Fig. 2.25

Fig. 2.26 Fig. 3.1 Fig. 3.2

a Chloride profiles in the compressive zone of pure bending section subjected to cyclic wet-dry; b Chloride profiles in the tensile zone of pure bending section subjected to cyclic wet-dry . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . a Configuration of specimens for uniaxial tension fatigue testing (unit: in mm) b Specimens after casting . . . . . . . . . . . Tensile fatigue loading system . . . . . . . . . . . . . . . . . . . . . . . . Tensile load time relation curve . . . . . . . . . . . . . . . . . . . . . . . Scheme of the setups for moisture desorption experiments . . . a Specimens arrangement in the tunnel; b Temperature and humidity sensor; c Placed sensor; d Airflow generating and the wind tunnel; e Wind speed dial . . . . . . . . . . . . . . . . . Cumulative water desorption for concrete of unit square meters; a Effect of fatigue loading and w/b on the moisture loss; b Effect of fatigue loading level and SCMs on the moisture loss . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Decrease of RH for surface concrete; a Effect of fatigue loading and w/b on the RH; b Effect of fatigue loading level and SCMs on the RH evolution . . . . . . . . . . . . . . . . . . . Influence of wind speed on the surface factor . . . . . . . . . . . . . Influence of temperature on the surface factor . . . . . . . . . . . . Correlation between magnification factor and damage variable in tensile fatigue-damaged concrete . . . . . . . . . . . . . . a The prepared cubic specimens for chloride penetration tests; b Illustration of the powder collection for determining the chloride content in concrete . . . . . . . . . . . . . . . . . . . . . . . Free chloride profiles in PC mixture concrete subjected to various levels of fatigue loads in immersion at different ages; a 30 days; b 45 days; c 60 days (Note the number between mixture ID and the exposure duration is the strain gauge number; for example, PC15-2–45 days represents the cubic specimen for chloride penetration tests is cut from the location of strain gauge #2 in PC15 at 45 days). . . . . . . . . . . . . . . . . . Free chloride profiles in PC mixture concrete subjected to various levels of fatigue loads in wet-dry condition at different ages; a 30 days; b 45 days; c 60 days . . . . . . . . . The fitted values for apparent chloride diffusion coefficient for concrete subjected to various levels of fatigue loads; a Immersion; b Wet-dry cycles . . . . . . . . . . . . . . . . . . . . . . . . Correlation between magnification factor and residual strain in fatigue-damaged concrete . . . . . . . . . . . . . . . . . . . . . . . . . . Configuration of prefabricated specimens . . . . . . . . . . . . . . . . Arrangement of sensors along the crack path of cracked specimens . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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List of Figures

Fig. 3.3 Fig. 3.4

Fig. 3.5

Fig. 3.6

Fig. 3.7

Fig. 3.8

Fig. 3.9

Fig. 3.10

Fig. 3.11

Fig. 3.12 Fig. 3.13

Fig. 3.14

Fig. 4.1 Fig. 4.2

Arrangement of specimens in the environmental chamber . . . . . Time-dependent changing of RH in cracks of concrete subjected to a cyclic change of external artificial RH. a specimen 1; b specimen 2; c specimens 3 . . . . . . . . . . . . . . . . Time-dependent RH change in cracks of concrete subjected to a cyclic change of external artificial temperature. a specimen 1; b specimen 2; c specimen 3. . . . . . . Time-dependent changing of RH in cracks of concrete subjected to simulated fog environment. a specimen 1; b specimen 2; c specimens 3 . . . . . . . . . . . . . . . . . . . . . . . . . . . Time-dependent changing of RH in cracks of concrete subjected to artificial rainfall and the subsequent drying. a specimen 1; b specimen 2; c specimens 3 . . . . . . . . . . . . . . . . EMPA mapping of chloride concentration of a OPC concrete subjected to immersion for 30 days b OPC concrete subjected to immersion for 60 days c GGBS concrete subjected to immersion for 30 days d GGBS concrete subjected to immersion for 60 days . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . a EMPA Line scanning of chloride concentration for GGBS concrete subjected to immersion for 30 days b Area Line scanning of chloride concentration for GGBS concrete subjected to immersion for 60 days . . . . . . . . . . . . . . . . . . . . . . a Point EMPA scanning results of chloride concentration in the crack surface for OPC concrete subjected to immersion for 60d b Point EMPA scanning results of chloride concentration in the crack surface for GGBS concrete subjected to immersion for 60d . . . . . . . . . . . . . . . . . . . . . . . . . . a Chloride distribution perpendicular to the crack surface of GGBS concrete at surface crack width 0.1 mm b Chloride distribution perpendicular to the crack surface of GGBS concrete at surface crack width 0.15 mm . . . . . . . . . . . . . . . . . . Definition of the cracked zone . . . . . . . . . . . . . . . . . . . . . . . . . . a Chloride ions profiles in the cracked zone for the specimen with a surface crack width of 0.1 mm. b Chloride ions profiles in the cracked zone for the specimen with a surface crack width of 0.2 mm . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . a Chloride ions profiles for the specimen with a surface crack width of 0.1 mm. b Chloride ions profiles for the specimen with a surface crack width of 0.2 mm . . . . . . . Salt-fog chambers during; a salt-fog spraying (wetting period); b drying period . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Illustration of specimen placement orientation in the salt-fog chamber; A an angle of 0° to the horizontal; B an angle of 60° to the horizontal; C an angle of 30° to the horizontal . . . . . . . . .

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Fig. 4.3 Fig. 4.4

Fig. 4.5 Fig. 4.6

Fig. 4.7

Fig. 4.8

Fig. 4.9 Fig. 4.10

Fig. 4.11 Fig. 4.12 Fig. 4.13

Fig. 4.14 Fig. 5.1

List of Figures

Arrangements of RH sensors in the mortar (unit: in mm) a Top view; b Side view; c During specimen preparation . . . Evolution of RH in the internal mortar exposed to the salt-fog wet-dry cycles a OPC_42; b OPC_47; c OPC_52; d SL_47; e FA_47; f SF_47 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Pore size distribution of mortar a OPC with various w/c ratios; b OPC with/without SCMs . . . . . . . . . . . . . . . . . . . . . . . . . . . Free chloride concentration profiles in mortar exposed to cyclic wet-dry salt-fog environments; a OPC_42; b OPC_47; c OPC_52; d SL_47; e FA_47; f SF_47 . . . . . . . . . . . . . . . . . Free chloride concentration profiles in mortar exposed to cyclic wet-dry salt-fog environments; a OPC with various w/c ratios after 30 salt-fog cycles; b OPC with/without SCMs after 15 cycles; c OPC with/without SCMs after 30 cycles; d OPC with/without SCMs after 45 cycles . . . . . . . . . . . . . . . . . . . . . Free chloride concentration profiles in OPC_47 mixture with three different exposure orientations after; a 30 salt-fog cycles; b 45 salt-fog cycles . . . . . . . . . . . . . . . . a Configuration of a sample with loading system. b Setup for chloride penetration test . . . . . . . . . . . . . . . . . . . . Free chloride distributions in stress-free concrete; a PC mixture exposed to condition I (control) and II; b PC mixture exposed to condition I and III; c FS mixture exposed to condition I and II; d FS mixture exposed to condition I and III (Note The nomenclature PC_I10_00 means PC mixture was exposed to 10 wet-dry-carbonation cycles at exposed condition I. The last two digitals 00 mean stress-free status. Other nomenclatures use the same rules) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Carbonation depths of stress-free concrete samples . . . . . . . . . Pore size distribution for PC concrete obtained at different carbonation regions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Free chloride distributions in stressed concrete; a PC mixture exposed to condition II; b PC mixture exposed to condition III; c FS mixture exposed to condition II; d FS mixture exposed to condition III (Note The last two digitals 03 and 06 in the nomenclature mean concrete samples were applied with 30% and 60% of ultimate loading capability, respectively) . . . . . . . Carbonation depths of stressed concrete samples . . . . . . . . . . Load-induced crack mapping of the RC beam (crack width and length in mm) . . . . . . . . . . . . . . . . . . . . . . .

..

64

..

65

..

67

..

69

..

71

..

73

..

74

.. ..

77 78

..

79

.. ..

81 82

..

86

List of Figures

Fig. 5.2

Fig. 5.3

Fig. 5.4

Fig. 5.5

Fig. 5.6 Fig. 5.7

Fig. 5.8

Fig. 5.9

Fig. 5.10

The RC beam for sample preparation in 3D laser scanning and EMPA analysis. Half of the beam was further sliced for optical and EMPA observation (left side), while another half was destructed to obtain the corroded rebar for 3D laser scanning (right side) . . . . . . . . . . . . . . . . . . . . . . . . . . . . Illustrations of the quantitative determination of the thickness of rust at the rebar-concrete interface and the corrosion-induced crack widths. a Optical images of polished corroded rebar embedded in concrete. Total 20 points along the circumference of rebar was measured for rust thickness; b measurement of the thickness of rust around corroded rebar; c calculation of total corrosion-induced crack width as a function of its distance from the center of rebar . . . . . . . An example of the prepared EMPA specimens for analyzing the compositional and phase distribution around the corroded rebar in concrete . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Illustration of the reconstructed corroded rebar models in the processing software interface. a Reconstructed rebar model with a scanning interval of 1.0 mm; b analysis of the loss fraction of cross-sectional area of corroded rebar in a 20 mm segment . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Crack maps of corroding RC beam after natural exposure for 4 years (unit: mm) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Observation of close of transverse cracks in the RC beam after 4-year natural corrosion. a Self-healing of transverse cracks by further hydration of cement; b transverse cracks filled with rusts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Correlation between the initial width of load-induced cracks and the width of corrosion-induced cracks in the tension surface of RC beam (The crack width measurement was taken at an interval of 4 mm for the two longitudinal rebars in the tension surface of RC beam) (unit: mm) . . . . . . . . . . . . a Evolution of total corrosion-induced crack widths as a function of distance from rebar; b correlation of corrosion-induced crack width at concrete surface and the average thickness of rusts . . . . . . . . . . . . . . . . . . . . . . Correlation between loss fraction of cross-sectional area of corroded rebar and the width of initial transverse cracks along the beam. Each data point represents the average loss fraction of cross-sectional area of corroded rebar within every 20 mm analyzing segment (unit: mm) . . . . . . . . . . . . . .

xix

..

86

..

87

..

88

..

88

..

89

..

90

..

90

..

93

..

94

xx

Fig. 5.11

Fig. 5.12

Fig. 5.13

Fig. 5.14

Fig. 5.15 Fig. 5.16 Fig. 5.17 Fig. 5.18 Fig. 5.19 Fig. 5.20

Fig. 5.21

Fig. 5.22 Fig. 5.23

Fig. 5.24

Fig. 5.25

List of Figures

a The region in the longitudinal rebar which is contacted with stirrups shows less corrosion in comparison to other regions; b the stirrups placed outside the longitudinal rebar suffers from severe corrosion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Correlation between the average loss fraction of the cross-sectional area of corroded rebar and the maximal loss fraction of the cross-sectional area of corroded rebar within a 20 mm analyzed segment of corroded rebar . . . . . . . Representative backscattered electron images of the rust layer in the corroded rebar and concrete interface obtained from various EMPA samples . . . . . . . . . . . . . . . . . . . . . . . . . Correlation of the weight percentage of oxygen and the weight percentage of iron in the rust layer obtained by EMPA point analysis. The composition of millscale is relatively constant, while that of rust varies considerably at different locations . . . Configuration of specimens and the layouts of the rebars (unit: mm) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Illustration of the sectioning planes for each specimen during the steel corrosion propagation stage (unit: mm) . . . . . Illustration of the measurement of rust thickness around the circumference of corroded rebar (unit: mm) . . . . . . . . . . . Illustration of the calculation of the total area of the rust along the circumference of corroded rebar . . . . . . . . . . . . . . . Illustration of the measuring locations for the corrosion-induced concrete cracking at surface (unit: mm) . . . Time-dependent spatial distribution of rust thickness as a function of coordinate along the circumference of corner rebar. a C10; b C15; c C20; d C25 (d means day) . . . . . . . . . Conceptual models of the rust development in corroded rebar. a Side-located rebar or corner-located rebar with C2/C1 being larger than 1.5; b corner-located rebar with C2/C1 being smaller than 1.5 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . For C10 position at 114 days, the rust thickness profiles need to be fitted by three Gaussian functions . . . . . . . . . . . . . a Subdivision of the rust layer for C10 position; b time-dependent evolution of the Arust for these three sub-regions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Analysis of the fitting parameters in the Gaussian models as a function of C2/C1 ratio, a a1; b a2; c h. The parameters of the corner-located rebar in C10 position (i.e., C2/C1 = 1) are not included because its rust distribution patterns are different from the other three cases (d means day) . . . . . . . . . Time-dependent evolution of q and crack width for rebar located at various positions. a C10; b C15; c C20; d C25 . . .

..

94

..

96

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99

. . 101 . . 101 . . 102 . . 103 . . 104 . . 105

. . 106

. . 107 . . 108

. . 109

. . 110 . . 111

List of Figures

Fig. 6.1

Fig. 6.2 Fig. 6.3

Fig. 6.4

Fig. 6.5

Fig. 6.6

Fig. 6.7 Fig. 6.8

Illustration of the setup for accelerated corrosion techniques. a Artificial climate exposure method; b Impressed current method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Tests of load bearing capacity of the corroded RC beams . . . Distribution of mass loss ratio of segmented rebar along the longitudinal direction of corroded rebar in RC beams accelerated by artificial climate exposure method. a Specimens without sustained loads; b Specimens with 30% sustained loads; c Specimens with 60% sustained loads . . . . . . . . . . . . . Distribution of mass loss ratio of segmented rebar along the longitudinal direction of corroded rebar in RC beams accelerated by impressed current method . . . . . . . . . . . . . . . . Correlation between the average mass loss ratio of corroded rebar and the measured maximum crack width (the regression was performed by assuming that the line trend passes through the original point, i.e., the intercept is set to be zero) . . . . . . . Load and deflection responses of corroded RC beams accelerated by artificial climate exposure method. a Specimens without sustained loads; b Specimens with 30% sustained loads; c Specimens with 60% sustained loads . . . . . . . . . . . . . Load and deflection responses of corroded RC beams accelerated by impressed current method . . . . . . . . . . . . . . . . Correlation between the ratio of ultimate load capacity of corroded beams w.r.t. the reference beam and the mass loss of corroded rebar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

xxi

. . 118 . . 119

. . 128

. . 129

. . 129

. . 131 . . 132

. . 132

List of Tables

Table 2.1 Table Table Table Table Table Table Table Table Table Table Table

2.2 2.3 2.4 3.1 3.2 3.3 3.4 4.1 4.2 4.3 5.1

Table 5.2

Table 5.3

Table 6.1

Table 6.2

Mix proportion (kg/m3) and mechanical properties of concrete . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Fatigue loading parameters and specimen number . . . . . . . . . Environmental conditions . . . . . . . . . . . . . . . . . . . . . . . . . . . . Fatigue loading protocol . . . . . . . . . . . . . . . . . . . . . . . . . . . . Crack characteristics of prefabricated concrete specimens . . . Crack widths of specimens . . . . . . . . . . . . . . . . . . . . . . . . . . Characteristics of crack for specimen 1 . . . . . . . . . . . . . . . . . Characteristics of crack for specimen 2 . . . . . . . . . . . . . . . . . Mixing proportion of mortar (kg/m3) . . . . . . . . . . . . . . . . . . . Mixing proportion of concrete (kg/m3) . . . . . . . . . . . . . . . . . Exposure conditions for one wet-dry-carbonation cycle . . . . . Representative optical microscopy images of corroded rebar-concrete cross section sliced from four different loading zones in the RC beam . . . . . . . . . . . . . . . . . . . . . . . . EMPA mapping of the corroded rebar-concrete sliced specimens obtained in pure flexural tension and shear-tension zones in the RC beam. The upper sides of the figures are the exposure surface to the external environment. Specimens SP1–6 are from pure flexural tension zone, while SP7, 8 are from shear-tension zone . . . . . . . . . . . . . . . EMPA point microanalysis data (weight percentage of the iron and oxygen element) corresponding to the EDS ID shown in Fig. 5.13 . . . . . . . . . . . . . . . . . . . . . Designed RC beams with various levels of sustained flexural loads and two different methods for corrosion acceleration . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Cracking maps of corroded RC beams accelerated by artificial climate exposure method (For each beam, the back surface, the front surface, and the surface in flexural tension are shown, in the order from top to the bottom) . . . . . . . . . . . . . . . . . . .

. . . . . . . . . . .

25 26 28 33 45 48 56 57 62 73 75

..

92

..

97

. . . . . . . . . . .

. . 100

. . 116

. . 120 xxiii

xxiv

Table 6.3

Table 6.4 Table 6.5

List of Tables

Cracking maps of corroded RC beams accelerated by impressed current method (only the surface at flexural tension is shown for each beam because the front and back surfaces are crack-free) . . . . . . . . . . . . . . . . . . . . . . . . 126 Failure patterns of the corroded RC beams accelerated by artificial climate exposure method . . . . . . . . . . . . . . . . . . . . . 133 Failure patterns of the corroded RC beams accelerated by impressed current method . . . . . . . . . . . . . . . . . . . . . . . . . . . 136

Chapter 1

Introduction

Abstract This chapter introduces the characteristics of the marine environments, the physiochemical properties of modern concrete materials, the transport mechanism of chloride and other deleterious agents into concrete materials, and the electrochemical process of steel corrosion in concrete. This chapter lays the foundation for a better understanding of the degradation mechanisms of concrete structures subjected to combined chloride-induced steel corrosion and service loads.

1.1 Background Since the early 1950s, significant attention has been given to the issue of chlorideinduced steel corrosion in reinforced concrete (RC) structures (Al-Sulaimani et al. 1990; Lewis and Copenhagen 1959; Page and Treadaway 1982). Chloride-induced steel corrosion has been one of the most common deterioration mechanisms in the existing RC structures, particularly those exposed to the marine environments such as the coastlines of Southern China. The structural concrete exposed to the marine environment degrades due to a combinational effect of chloride-induced steel corrosion, mechanical loading, and others (Gao et al. 2013; Maage et al. 1996; Mehta 2002; Poupard et al. 2006; Tang et al. 2015; Xi et al. 2001; Ye and Jin 2019; Yoon et al. 2000). Any damage in concrete will soon be revealed and manifested in such harsh environments since the couplings of environmental factors and mechanical loads potentially accelerate the deterioration of in-service RC structures.

1.2 Marine Environment The marine environment is one of the most inhospitable service conditions for RC structures. The structural concrete in the marine environments is subjected to four different exposure conditions, classified as submerged zone, tidal zone, splash zone, and atmospheric zone, as shown in Fig. 1.1 (Mehta 2002). © The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2020 H. Ye et al., Chloride-Induced Steel Corrosion in Concrete Under Service Loads, https://doi.org/10.1007/978-981-15-4108-7_1

1

2

1 Introduction

Rain reducing surface salt concentration Diffusion in response to salt concentration

Airborne salt and occasional saltwater inundation

Evaporation giving a salt concentration

Atmospheric zone

Water table Capillary absorption into partially saturated concrete

Permeation by pressure head

Splash zone

Wick action Splash/spray

Tidal range

Diffusion of salt from seawater

Tidal zone

Submerged zone

Fig. 1.1 Illustration of multiple physical and chemical deterioration actions on concrete structures (modified from Alexander 2016)

For RC structures exposed to the submerged zone, the concrete degrades due to direct contact with seawater. The hydrostatic pressure of seawater on the submerged concrete pushes the corrosive ions and molecules into the connected pores of the concrete. The typical composition of seawater is mainly 3.5% by weight of salts, including principal ions of Na+ , Mg2+ , Cl− , and SO4 2− . In addition, oxygen (O2 ) gas, carbon dioxide (CO2 ), and hydrogen sulfide (H2 S) dissolved in the seawater can also participate in the physiochemical and electrochemical process of chloride-induced steel corrosion. The CO2 concentration in some geographic locations can be relatively high, which may result in a noticeable pH reduction in concrete pore solution and decline chloride binding capacity of concrete. The influence of carbonation on the chloride penetration in concrete will be discussed in detail in Chap. 4. In addition, the leaching of cementitious constitutes in concrete in the submerged zone leads to

1.2 Marine Environment

3

the decomposition of hydration products, enlargement of porosity, and weakening of strength. The Mg2+ and SO4 2− may also react with the constituents of concrete to form deleterious products that cause other durability problems such as sulfate attack. The tidal and splash zones are considered the most unfavorable due to the cyclic wet-dry action, which accelerates chloride ingress and steel corrosion (Boddy et al. 1999; Hong and Hooton 1999; Ye et al. 2012). In a typical tidal, the concrete portion in tidal zones experiences twice-a-day cycles of wet and dry. This wet-dry action may be accompanied by the heating-cooling action due to the temperature difference in air and seawater, as well as freeze-thaw action in the cold climates. In regards to the splash zone, the wave action mainly caused by the action of wind on seawater can result in more frequent and unpredictable cyclic wet-dry action, in addition to the erosion action, depending on the wave energy. In Southern China, typhoons and hurricanes are capable of generating exceedingly high and strong waves, causing severe impacts on marine structures. The cyclic wet-dry action renders the salts in seawater enter the surface layer of concrete more quickly that in a saturated condition. The moisture evaporation during the drying stage results in concentration increase of salts in the pore solution of concrete, and probable crystallization with salts deposited at the surface pores. The capillary suction during wetting stages drives the salts into concrete, resulting in massive accumulation of chloride ions at the skin layers. Also, the hysteresis behavior of moisture absorption and desorption contributes to the so-called maximum phenomenon in chloride profiles in surface concrete. In addition, the O2 and CO2 gases can enter the concrete cover layer of RC concrete during drying stages, participating in the steel corrosion reaction. For concrete exposed to atmospheric zone, the salt-fog environment also results in a seasonal wet-dry action on concrete. In summer, the coastal fog is formed which carries droplets of sweater caused by the wave action. The salt-bearing fog results in seawater condensation at the surface of concrete, in which the chloride concentration at the surface increases gradually over time up to a few years after construction. In the winter, the cold air from land passes over the warmer and more humid environment of seawater, causing low stratus clouds. The salt-fog environment affects a wide range of infrastructure since the wind can carry seawater over long distances inland. The temperature of seawater and air above the seawater affects the degradation process of RC structures under chloride-induced steel corrosion. The surface temperature of seawater varies with the geographic location and climate from the freezing temperature of −2 °C to above 30 °C. Temperature affects the kinetics and pathway of physicochemical and electrochemical reactions involved in the corrosion process in RC structures, such as chloride diffusion coefficients, moisture evaporation rate, electrical resistivity of concrete, among others. In Southern China, the daytime air temperature can reach 40 °C, and the high humidity is frequently caused by the high evaporation of moisture at high-temperature conditions.

4

1 Introduction

1.3 Composition and Properties of Concrete Concrete is a composite material made of aggregates (filler) and binder. Modern concrete also contains chemical admixtures (e.g., superplasticizer, air entrainer) and supplementary cementitious materials (SCMs), in addition to the cement, water, and aggregates. Ordinary portland cement (OPC), which is manufactured usually by calcining limestone and clays at 1450 °C (called clinker) and then ground with gypsum, is the most commonly used binder. Other alternative cementitious materials, including geopolymer and alkali-activated materials, ye’elimite cement, magnesium oxychloride cement, and calcium aluminate cement, are used in some specific circumstance or environments (Juenger et al. 2011; Scrivener et al. 2018). The main constituent components of clinker are tricalcium and dicalcium silicates (C3 S and C2 S), the aluminate and ferroaluminate of calcium (C3 A and C4 AF). The C3 S and C2 S are also called alite and belite in cement terminology. Gypsum (CSH2) is incorporated into clinkers to avoid the flash setting due to the hydration of aluminates (Taylor 1997). Upon mixing with water, the compounds of OPC dissolves and precipitates out, forming various types of hydration product. The main hydration products in hardened OPC paste are calcium–silicate–hydrate (C–S–H) and portlandite (CH), which are produced due to hydration of C3 S and C2 S. The hydration reaction can be written as (Mindess et al. 2003): 2C3 S + 6H = C3 S2 H3 + 3Ca(OH)2

(1.1)

2C2 S + 4H = C3 S2 H3 + Ca(OH)2

(1.2)

However, it should be noted that the stoichiometry of C–S–H is not fixed, but varies considerably spatially in the microstructure of hardened cement pastes (Richardson 2008). The nature of C–S–H has always been an intriguing subject among the cement research community (Jennings 2008; Mindess et al. 2003; Richardson 2008). Well acknowledged nanostructure models of C–S–H include the colloidal model proposed by Powers and Brownyard (Powers and Brownyard 1946), the layered structure model proposed by Feldman and Sereda (Feldman and Sereda 1968), and the colloidal model proposed by Jennings and his co-workers (Jennings 2000, 2008). The C–S–H accounts for about 50–60% of the volume of hydrated cement pastes and contributes dominatingly to the strength, shrinkage and creep, transport properties, and chemical durability of concrete (Allen et al. 2007; Jennings 2000, 2008; Youssef et al. 2011). The hydration of C3 A without the presence of gypsum results in the rapid formation of hydrogarnet phases, which causes ‘flash’ setting of concrete and allows no time for concrete placement and finishing. The hydration C3 A in the presence of gypsum results in concurrent formation of calcium aluminate tri-sulfate hydrate (AFttype phase, e.g., ettringite) and/or calcium aluminate monosulfate hydrate (AFm-type phase, e.g., monosulfate), depending on the gypsum-to-C3 A ratio. A higher gypsumto-C3 A ratio results in more ettringite formation over monosulfate. In a typical OPC

1.3 Composition and Properties of Concrete

5

with 5% gypsum, the formation of ettringite occurs first before the setting of cement (Mindess et al. 2003), C3 A + 3CSH2 + 26H → C6 AS3 H32

(1.3)

As the gypsum is depleted while C3 A is still available, the ettringite will become unstable and reacts with C3 A to form monosulfate, 2C3 A + C6 AS3 H32 + 4H → 3C4 ASH12

(1.4)

However, if sulfate ions become available again due to a number of reasons, the monosulfate will transform back to ettringite, C4 ASH12 + 2CSH2 + 16H → C6 AS3H32

(1.5)

Since ettringite is volumetrically expansive, the transformation of monosulfate to ettringite causes expansion and cracking (i.e., sulfate attack). The sulfate attack may occur in concrete exposed to the marine environment due to the presence of sulfate ions in seawater, although evidence also shows the presence of chloride results in instability of ettringite. Low C3 A content in cement is preferred in terms of sulfate resistance since the amount of AFm-type phase is reduced. However, high C3 A cement may be favored in some circumstances in terms of chloride and carbonation resistance of concrete. The AFm-type phases have a layered structure with the intercalation sites accommodatable by hydroxyl, sulfate, and carbonate groups (Matschei et al. 2007). Due to the layered structure and ion-exchange capacity, AFm-type phases play an important role in chloride binding and CO2 absorption in concrete, all of which can influence the chloride-induced steel corrosion in RC structures (Balonis et al. 2011). The hydration reaction of C4 AF is not introduced, as it is similar to those of C3 A, although at a comparatively low significance.

1.4 Supplementary Cementitious Materials Supplementary cementitious materials (SCMs), including ground granulated blastfurnace slag (GGBS), pulverized fly ash (FA), and silica fume (SF), are widely used in modern concrete to partially replace OPC (Lothenbach et al. 2011). Some of the SCMs are industrial byproducts or locally accessible pozzolans which reduces the CO2 emission and energy consumption of concrete products. The incorporation of SCMs into concrete affects the cement hydration, phase assemblage, and pore structure, thus influencing the durability properties related to chloride-induced steel corrosion (Khatri et al. 1995; Lothenbach et al. 2011; Thomas et al. 2012). For instance, silica-rich SCMs reduce the amount of portlandite and lower the Ca/Si ratio in C–S–H in concrete through pozzolanic reaction, while aluminum-rich SCMs affect the amount and types of AFm and AFt-type phases in hardened concrete. The alteration in the type and composition of hydrated phases in concrete due to SCMs affect

6

1 Introduction

its resistance against chloride binding, carbonation, and sulfate attack. In addition, the change in the pore structure of concrete affects the transport properties and electrical resistivity, which are closely related to corrosion initiation and propagation stages of steel embedded in concrete.

1.5 Pore Structure and Transport Properties of Concrete Concrete is a multi-scale porous material, which allows the penetration of external gas and liquid substances into it. The resistance of concrete against the penetration of aggressive ions and molecules is closely related to the phenomenon and process involved in the chloride-induced steel corrosion, including moisture, chloride ions, CO2 and O2 ingress, and electrical current transfer within concrete. The moisture content in concrete, which mainly depends on the environmental condition and pore structure, affects the penetration process of ions and gas. The ions (e.g., chloride and sulfate ions) have to be dissolved in the liquid phase; while gases (e.g., O2 gas) have to transport via gaseous space in pores into interior concrete. The classification of water types in concrete is still controversial; but roughly speaking, there are three types, i.e., capillary water, gel water, and chemically bound water. Only the capillary water contributes considerably to the transport properties of concrete. There are many mechanisms that ions and molecules get into the interior of concrete, including diffusion, permeation, migration, and capillary suction (Claisse 2005; Samaha and Hover 1992). The diffusion is related to the movement of substance due to concentration gradients, such as the movement of chloride in saturated concrete. The diffusion process is described using Fick’s laws. The permeation is related to the pressure gradients, which is typically described by Darcy’s law for liquid. The migration of ions is driven by the electric field. Capillary suction is due to the capillary action inside capillaries of concrete driven by the surface tension.

1.6 Electrochemical Reaction of Chloride-Induced Corrosion Due to the variation in chloride concentration at the interfaces between steel reinforcements and concrete, localized corrosion typically takes place, leading to nonuniform corrosion (i.e., pitting corrosion). Figure 1.2 illustrates the electrochemical corrosion mechanisms and the corresponding half-cell reactions for pitting corrosion, in which anodic iron oxidation (1.6) and cathodic oxygen reduction reaction (1.7) take places (Revie 2008): Anodic (oxidation) half-cell reaction: Fe - 2e− → Fe2+ Cathodic (reduction) half-cell reaction: O2 + 2H2 O + 4e− → 4OH−

(1.6) (1.7)

1.6 Electrochemical Reaction of Chloride-Induced Corrosion

7

K+, Na+, Ca2+, Fe2+

Concrete

Pore soluƟon

OH-, Cl-, SO42OHRust

H2O, O2

Rust

ClFe2+

Steel

Pit

Fe2+

neFe

Anode

Cathode

Fig. 1.2 Illustration of the electrochemical corrosion reactions at the steel-concrete interface

The anodic process liberates electrons in the metallic phase and forms dissolved iron ions. The electrons are transported within the metal to the cathodic regions and consumed there. In order to complete the circuit, the current flows inside the concrete from the anodic to the cathodic regions, in the form of ionic movement in the pore solution of concrete. It is clear that the states and properties of concrete can have a considerable influence on the corrosion process. In addition, the corrosion rate of steel is determined by the slowest of the above processes. As the electrical resistance of reinforcement in concrete is always negligible and would not be the kinetically controlling process. Thus, the corrosion rate can be controlled by the following three forms: (i) Anodic control: reinforcement remains passive; (ii) Cathodic control: the rate of oxygen reaching the surface of reinforcement is low; (iii) Ohmic control: the electrical resistance of the concrete is high (Bertolini et al. 2013; Broomfield 2003).

1.7 Content of This Book In most in-service RC structures, the sustained loads have a noticeable impact on several important aspects in regards to the chloride-induced steel corrosion, including transport properties, corrosion initiation and propagation process, and mechanical and failure behaviors of concrete members. The motivation of this work is to enable a better understanding of the influence of service loads on the chloride-induced steel corrosion of reinforced concrete. In Chap. 2, the effects of stress on the moisture and chloride diffusion coefficient of concrete are discussed with the introduction of the loading effect factor. In Chap. 3, the effects of cracks on the moisture influential depth and chloride diffusion coefficient of concrete are discussed. In Chap. 4, the influence of various environmental conditions on the chloride penetration in concrete is discussed, including

8

1 Introduction

cyclic seawater wet-dry, salt-fog wet-dry, and carbonation. In Chap. 5, the corrosion propagation characteristics in pre-cracked concrete beams corroded under sustained loading are discussed. In Chap. 6, the mechanical degradation of concrete beams corroded under sustained service loads is discussed.

References Al-Sulaimani, G. J., Kaleemullah, M., & Basunbul, I. A. (1990). Influence of corrosion and cracking on bond behavior and strength of reinforced concrete members. Structural Journal, 87(2), 220– 231. Alexander, M. (2016). Marine concrete structures: design, durability and performance. Woodhead Publishing. Allen, A. J., Thomas, J. J., & Jennings, H. M. (2007). Composition and density of nanoscale calcium–silicate–hydrate in cement. Nature Materials, 6(4), 311. Balonis, M., Medala, M., & Glasser, F. P. (2011). Influence of calcium nitrate and nitrite on the constitution of AFm and AFt cement hydrates. Advances in Cement Research, 23(3), 129–143. Bertolini, L., Elsener, B., Pedeferri, P., Redaelli, E., & Polder, R. B. (2013). Corrosion of steel in concrete: Prevention, diagnosis, repair. Wiley. Boddy, A., Bentz, E., Thomas, M. D. A., & Hooton, R. D. (1999). An overview and sensitivity study of a multimechanistic chloride transport model. Cement and Concrete Research, 29(6), 827–837. Broomfield, J. P. (2003). Corrosion of steel in concrete: Understanding, investigation and repair. CRC Press. Claisse, P. (2005). Transport properties of concrete. Concrete International, 27(1), 43–48. Feldman, R. F., & Sereda, P. J. (1968). A model for hydrated Portland cement paste as deduced from sorption-length change and mechanical properties. Matériaux et Construction, 1(6), 509–520. Gao, J., Yu, Z., Song, L., Wang, T., & Wei, S. (2013). Durability of concrete exposed to sulfate attack under flexural loading and drying–wetting cycles. Construction and Building Materials, 39, 33–38. Hong, K., & Hooton, R. D. (1999). Effects of cyclic chloride exposure on penetration of concrete cover. Cement and Concrete Research, Elsevier, 29(9), 1379–1386. Jennings, H. M. (2000). Model for the microstructure of calcium–silicate–hydrate in cement paste. Cement and Concrete Research, 30(1), 101–116. Jennings, H. M. (2008). Refinements to colloid model of CSH in cement: CM-II. Cement and Concrete Research, 38(3), 275–289. Juenger, M. C. G., Winnefeld, F., Provis, J. L., & Ideker, J. H. (2011). Advances in alternative cementitious binders. Cement and Concrete Research, 41(12), 1232–1243. Khatri, R. P., Sirivivatnanon, V., & Gross, W. (1995). Effect of different supplementary cementitious materials on mechanical properties of high performance concrete. Cement and Concrete Research, 25(1), 209–220. Lewis, D. A., & Copenhagen, W. J. (1959). Corrosion of reinforcing steel in concrete in marine atmospheres. Corrosion, 15(7), 60–66. Lothenbach, B., Scrivener, K., & Hooton, R. D. (2011). Supplementary cementitious materials. Cement and Concrete Research, 41(12), 1244–1256. Maage, M., Helland, S., Poulsen, E., Vennesland, O., & Carl, J. E. (1996). Service life prediction of existing concrete structures exposed to marine environment. Materials Journal, 93(6), 602–608. Matschei, T., Lothenbach, B., & Glasser, F. P. (2007). The AFm phase in Portland cement. Cement and Concrete Research, 37(2), 118–130. Mehta, P. K. (2002). Concrete in the marine environment. CRC Press. Mindess, S., Young, F. J., & Darwin, D. (2003). Concrete. Technical Documents.

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Page, C. L., & Treadaway, K. W. J. (1982). Aspects of the electrochemistry of steel in concrete. Nature, 297(5862), 109. Poupard, O., L’hostis, V., Catinaud, S., & Petre-Lazar, I. (2006). Corrosion damage diagnosis of a reinforced concrete beam after 40 years natural exposure in marine environment. Cement and Concrete Research, 36(3), 504–520. Powers, T. C., & Brownyard, T. L. (1946). Studies of the physical properties of hardened Portland cement paste. ACI Journal Proceedings. Revie, R. W. (2008). Corrosion and corrosion control: an introduction to corrosion science and engineering. Wiley. Richardson, I. G. (2008). The calcium–silicate–hydrates. Cement and Concrete Research, 38(2), 137–158. Samaha, H. R., & Hover, K. C. (1992). Influence of microcracking on the mass transport properties of concrete. Materials Journal, 89(4), 416–424. Scrivener, K. L., John, V. M., & Gartner, E. M. (2018). Eco-efficient cements: Potential economically viable solutions for a low-CO2 cement-based materials industry. Cement and Concrete Research, 114, 2–26. Tang, S. W., Yao, Y., Andrade, C., & Li, Z. J. (2015). Recent durability studies on concrete structure. Cement and Concrete Research, 78, 143–154. Taylor, H. F. W. (1997). Cement chemistry. Thomas Telford. Thomas, M. D. A., Hooton, R. D., Scott, A., & Zibara, H. (2012). The effect of supplementary cementitious materials on chloride binding in hardened cement paste. Cement and Concrete Research, 42(1), 1–7. Xi, Y., Willam, K., Ababneh, D. M. F., Nakhi, A., Kong, J. S., & Nogueira, C. L. (2001). Accelerated testing and modeling of concrete durability subjected to coupled environmental and mechanical loading (pp. 45–56). Long term durability of structural materials: Elsevier. Ye, H., & Jin, N. (2019). Degradation mechanisms of concrete subjected to combined environmental and mechanical actions: a review and perspective. Computers and Concrete, 23(2), 107–119. Ye, H., Jin, N., Jin, X., & Fu, C. (2012). Model of chloride penetration into cracked concrete subject to drying-wetting cycles. Construction and Building Materials, 36, 259–269. Yoon, S., Wang, K., Weiss, W. J., & Shah, S. P. (2000). Interaction between loading, corrosion, and serviceability of reinforced concrete. Materials Journal, 97(6), 637–644. Youssef, M., Pellenq, R. J.-M., & Yildiz, B. (2011). Glassy nature of water in an ultraconfining disordered material: The case of calcium−silicate−hydrate. Journal of the American Chemical Society, 133(8), 2499–2510.

Chapter 2

Chloride Ingress in Stressed Concrete

Abstract This chapter focuses on the influence of microscopic damage and cracks on the chloride ingress in concrete. In this chapter, a loading effect factor is defined and the influence of flexural and fatigue loads-induced damage on the moisture transfer and chloride ingress in concrete is discussed.

2.1 Introduction The concrete materials in most in-service RC structures are under stressed conditions caused by either mechanical loads, environmental loads (e.g., thermal stress), or both. The loading can cause microstructural alteration (Gerard et al. 1998) and trigger micro- and macro-cracking of concrete (Gowripalan et al. 2000), any of which unambiguously impacts the chloride penetration and steel corrosion process in concrete. Based on the form of external mechanical loads (e.g., uniaxial or multi-axial sustained loads, pure shear, flexural loads, or cyclic loads), the microstructural alteration of concrete varies; as such, their influence on the transport properties of concrete are different. Establishing the quantitative relationship between the damage and its impact on the moisture transfer and chloride transport in concrete is of significant importance. By summarizing the results of these existing studies (Hoseini et al. 2009), there exists critical compressive stress of approximately 70–80% of the ultimate load, below which the mass penetration decreases with the load increases. Otherwise, the mass penetration increases with the compressive load increases. On the other hand, tensile stress usually accelerates the mass penetration regardless of its magnitudes, and it becomes even severer with the increase in tensile loads.

© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2020 H. Ye et al., Chloride-Induced Steel Corrosion in Concrete Under Service Loads, https://doi.org/10.1007/978-981-15-4108-7_2

11

12

2 Chloride Ingress in Stressed Concrete

2.2 The Loading Effect Factor 2.2.1 Damage Influence Factor Compared with sound concrete, damaged concrete has numerous micro-cracks. The nature of damage in porous media including concrete, rocks, and soils is substantially reflected by the increase of porosity and enlargement of voids. Some attempts have been made on modifying the ionic diffusion coefficients using an introduced damage variable, among which the one proposed by (Gerard et al. 1998) is referred here, 





D(d) = D0 + Dmax 1 − 1 +

d dcr

n −1  (2.1)

where D0 and Dmax are the chloride diffusion coefficients of sound and completely damaged concrete, respectively. n and d cr are constants (n = 5; d cr = 0.4); d is the damage variable, which varies from 0 (sound) to 1 (completely damaged). There are some models available for the expression of the damage variable. On the basis of Eq. (2.1), the damage influence factor f (d), which quantifies the influence of concrete damage on chloride diffusion coefficient, can be written as    n −1  Dmax d D(d) =1+ 1− 1+ f (d) = D0 D0 dcr

(2.2)

where Dmax /D0 is the ratio of chloride diffusion coefficients of completely damaged and sound concrete. The relation curve between Dmax /D0 and f (d) is shown in Fig. 2.1. It can be seen that when either the stressed concrete is in the elastic state or the damage variable is less than 0.2, the value of the damage influence factor f (d) is 30

Damage influence factor, f (d)

Fig. 2.1 Relationship between damage variable d and f (d)

25 20

D

/D =8 max 0

D

/D =15 max 0

D

/D =20 max 0

D

/D =25 max 0

15 10 5 0 0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

Damage variable (d)

0.8

0.9

1.0

2.2 The Loading Effect Factor

13

almost constant at one. The model merely considers the damaging effect on the ionic diffusion coefficient of concrete after the external load is removed. The model is inadequate to consider the influence of sustained loads with or without the presence of damage on the chloride diffusion coefficients of concrete.

2.2.2 Definition of the Damage Variable Van Brakel et al. (van Brakel and Heertjes 1974) introduced the concepts of tortuosity and constrictivity in porous media and applied them to analyze the diffusion process in soil. The tortuosity and constrictivity are intrinsic properties of porous media. The tortuosity is usually defined as the ratio of actual flow path length to the straight distance between the ends of the flow path (The tortuosity equals 1 for a straight line); while constrictivity refers to the ratio of the diameter of the diffusing particle to the pore diameter, which is usually adopted to express the ink bottle effect. Based on the previous studies (Maekawa et al. 2003), tortuosity of concrete can be correlated to its porosity, as follows:   τ = −1.5 tanh 8.0 φpaste − 0.25 + 2.5

(2.3)

where τ is the tortuosity; φpaste is the effective porosity (i.e., volume fraction of capillary and gel pores) (m3 /m3 ). The constrictivity of concrete is influenced by the peak radius. The constrictivity equals to unit if the pore radius is constant along the pore length. When the pore radius varies along the pore length, the constrictivity can be calculated as (Maekawa et al. 2003)

  peak + 6.2 + 0.405 δ = 0.395 tanh 4.0 log rcp

(2.4)

peak

where δ is the constrictivity; rcp is the porous peak radius (m). Figure 2.2 displays the relationships between the tortuosity and porosity, as well as the constrictivity and peak radius. It can be seen that the tortuosity of a porous media no longer affects the mass transport as the porosity is greater than 0.60; the constrictivity of a porous media is significantly influenced by the peak radius changes from 100 to 1000 nm. The appropriate definition of the damage variable is critical to the reliable quantification of concrete damage effect on the ionic diffusion and other transport processes. As the nature of stress and damage effect on the chloride diffusion in concrete originates from the alteration in pore structure, the damage variable can be conveniently defined as the changes in tortuosity and constrictivity: d=

δstrain /τstrain − δini /τini δulti /τulti − δini /τini

(2.5)

14

2 Chloride Ingress in Stressed Concrete

5.0 4.5

Tortuosity (m/m)

4.0 3.5 3.0 2.5 2.0 1.5 1.0 0.0

0.1

0.2

0.3

0.4 3

0.5

0.6

3

Porosity (m /m )

(a) 1.0 0.9 0.8

Constrictivity

0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0 -7.0

-6.5

-6.0

log(r

peak

-5.5

-5.0

)

(b) Fig. 2.2 a Correlation between tortuosity and porosity; b Correlation between constrictivity and peak radius (unit in m)

2.2 The Loading Effect Factor

15

where τini and δini are the tortuosity and constrictivity of sound concrete, respectively; τstrain and δstrain are the tortuosity and constrictivity of partially damaged concrete at a certain strain level, respectively; τulti and δulti are the tortuosity and constrictivity of completely damaged concrete at ultimate strain, respectively. By adopting Eqs. (2.3) and (2.4), the calculation of damage variable, d, can base on the porosity and peak radius measured from mercury intrusion porosimetry (MIP) results. The assumption in Eq. (2.5) is in fact evidenced and supported by the MIP observation of pore size distribution of concrete with or without damage, as shown in Fig. 2.3. Fig. 2.3 MIP Pore size distribution of damaged concrete; a Pore size distribution of damaged concrete with different uniaxial tensile strain levels; b Pore size distribution of damaged concrete with different uniaxial compressive strain levels

16

2 Chloride Ingress in Stressed Concrete

2.2.3 Derivation of Loading Effect Factor Based on the nature of loading effects on pore characteristics and parameters of concrete, the loading effect factor F can be derived into two components: One is related to the original pore structure alteration caused by sustained loading (denoted as F (1) ); while another is related to the new pores and micro-cracks generated by loading effects (denoted as F (2) ). As loads applied, concrete gains certain values of strain correspondingly to certain loading magnitude, and the concrete strain can be measured by strain gauges attached to the concrete surface. Consequently, the concrete strain and damage variable can be used to reflect the loading level and the loading effect factor, F = F (1) + F (2) = F(ε, d)

(2.6)

The Taylor expansion of F = F(ε, d) at (0, 0) yields ∂ F  ∂ F  (2.7) (0,0) ε + (0,0) d + R1 (ε, d) ∂ε ∂d  where ε is the total strain; d is the damage variable; ∂∂εF (0,0) is the first-order partial derivatives of F w.r.t ε at (0, 0); ∂∂dF (0,0) is the first-order partial derivatives of F w.r.t d at (0, 0); R1 (ε, d) is the Lagrange remainder, R1 (ε, d) = 1 ∂ ∂ 2 (ε ∂ε + d ∂d ) F(θ ε, θ d), 2  (0 < θ < 1). Set F(0, 0) = A, ∂∂εF (0,0) = B, ∂∂dF (0,0) = C, and since under the circumstance of ε = 0 and d = 0, the load effect factor F = 1, thus A = 1 consistently, thus, Eq. (2.7) can be expressed as F(ε, d) = F(0, 0) +

F = 1 + B · ε + C · d + R1 (ε, d)

(2.8)

When the strain value ε is not large enough to cause concrete damage, C · d + R1 (ε, d) can be omitted. However, when the strain value ε becomes large enough to generate damage, F (1) and F (2) in F can be assumed as 

F (1) = 1 + B · ε F (2) = C · d + R1 (ε, d)

(2.9)

On the other hand, Eq. (2.2) can be rewritten as f (d) = 1 +

  n −1 d Dmax Dmax 1+ − D0 D0 dcr

(2.10)

Owing to Eq. (2.5), the damage variable d can be described as a function of strain ε, i.e., d = f (ε). As shown in Eq. (2.10), the last term is associated with the damage variable d, while the first two terms are irrespective of d. By multiplying the second

2.2 The Loading Effect Factor

17

term by a temporary expression of (1 + B1 · ε +C1 · d), where the correlation between B1 , C1 with B and C in Eq. (2.8) will be elaborated later, Eq. (2.10) can be further postulated as   n −1 d Dmax Dmax 1+ (1 + B1 · ε + C1 · d) − D0 D0 dcr    n −1  Dmax Dmax d =1+ · B1 · ε + 1 + C1 · d − 1 + D0 D0 dcr

F(ε, d) = 1 +

(2.11)

To check the boundary condition of F(ε, d) in Eq. (2.11), substituting ε = 0, d = 0 into it gives F(ε, d) = 1, which satisfies the condition for sound concrete. Let B = (Dmax /D0 )· B1 , where B is a parameter related to the concrete material irrespective of damage levels, and can be fitted by the correlation between strain value and chloride diffusion coefficients of undamaged but stressed concrete. It represents the circumstance where no damage occurs in the elastic stage of stressed concrete. This effect is linearly dependent on the elastic strain since tensile load or compressive load could induce elastic deformation in concrete, at which pore structure or porosity changes correspondingly with the strain. Equation (2.11) provides four circumstances to account for the loading effects with different changes of microstructure on chloride diffusion coefficient: · B1 ·ε, it represents the circumstance that there When d = 0, F(ε, d) = 1+ DDmax 0 is no permanent damage (i.e., neither micro- nor macro-cracks) in concrete but chloride diffusion coefficient changes with loads due to elastic deformation in the solid skeleton;    n −1  d 1 + C , it represents · d − 1 + (ii) When ε = 0, F(ε, d) = 1 + DDmax 1 dcr 0 (i)

the circumstance that damage has occurred in concrete, but without a sustained load (e.g., external loads have been removed after triggering permanent cracks); (iii) When ε = 0, d = 0, there is no loading effect and hence F(ε, d) = 1 (i.e., sound concrete); (iv) When ε = 0, d = 0, it represents the case where concrete is in damaged condition with a sustained load. According to the definition of damage variable d in Eq. (2.5), the damage variable d equals to 1 when concrete strain reaches the ultimate strain εulti , and is regarded as completely damaged correspondingly. As such, Eq. (2.11) can be derived as    n −1  Dmax 1 F(εulti , 1) = 1 + B · εulti + 1 + C1 − 1 + D0 dcr

(2.12)

Meanwhile, Eq. (2.10) can be rewritten as f (1) = 1 +

  n −1 1 Dmax Dmax 1+ − D0 D0 dcr

(2.13)

18

2 Chloride Ingress in Stressed Concrete

Let F(εulti , 1) = f (1), then the parameter C1 can be obtained as C1 = −

B · D0 εulti Dmax

(2.14)

The values of the unknown parameters in Eq. (2.11), namely B1 and C1 , can be obtained from experimental tests. The detailed procedure for determining B1 and C1 is available in Fu et al. (2015).

2.3 Effect of Flexural Loading on Chloride Ingress 2.3.1 Experimental Program Two types of concrete were evaluated, including OPC concrete and GGBS concrete (50% OPC replacement ratio). As shown in Fig. 2.4, RC beams with a dimension of 100 mm × 160 mm × 1400 mm were cast. Six HRB335 steel bars were used in each beam, 2 12 with a length of 1350 mm were laid on the bottom as longitudinal steels, while 4 8 with a length of 450 mm were laid on the top as erect steels. The HPB235 stirrups with a diameter of 6 mm and a spacing of 100 mm were laid in the beams. As shown in Fig. 2.5, a self-balancing loading device, which comprises loading cells, screws, and supports, was designed to apply sustaining loads on RC beams during chloride penetration tests. Each specimen demands four springs, four screws, four plats, and two permanent screws in the loading system. In addition, eight strain gauges were pasted along with the beam height at the midspan as indicated in Fig. 2.5. Every spring was calibrated using a universal testing machine, and the correlation between compressive load and spring length was established as shown in Fig. 2.6. Through fitting the test data, the correlation is expressed as F = −0.191 + 0.371L

R 2 = 0.998

(2.15)

Fig. 2.4 Layout of the reinforced concrete beam. The thickness of concrete covers was 15 mm at both the top and bottom of the cross section of the beam (unit: mm)

2.3 Effect of Flexural Loading on Chloride Ingress

19

Fig. 2.5 Four-point bending systems of the RC beam to initiate the transverse cracks and sustain the load during the entire experiment

10

Fig. 2.6 Fitted line for the spring calibration in the loading system

Load (kN)

8

6

4

2

Calibration points Fitted line

0

0

5

10

15

20

25

30

Length (mm)

where F is the magnitude of applied load (kN); L is the compressed length of spring (mm). As demonstrated in Fig. 2.5, a back-to-back four points loading scheme was applied using the designed loading system. Calipers were used to measure the compressed length (L) of springs, thus the magnitude of applied loading can be calculated according to the fitted equation shown in Eq. (2.15). During the loading process, strain values were recorded by the strain gauges. Upon reaching the target value, the permanent screws were tightened at the end of two beams. Also, strain gauges were attached to the concrete prism to measure the peak strain εu for tensile and compressive loads. Within the peak strain εu , different magnitudes of loads were applied on four concrete prisms to generate different degrees of damage. When the four concrete prisms were up to the target loading, the loads were removed and the residual strain ε p was recorded.

20

2 Chloride Ingress in Stressed Concrete

In order to obtain one-dimensional transport of chloride into the loaded concrete, the two side surfaces of beams were coated with paraffin. Two types of exposure environments were evaluated, including immersion and cyclic wet-dry, both of which were performed in a walk-in environmental chamber. The environmental temperature was kept constant at 40 °C and the specimens were exposed to a 5% NaCl solution. The acceleration tests lasted 4 months and the concentration of NaCl solution was adjusted periodically. Another set of specimens were subjected to accelerated penetration of chloride ions for 90 days through a cyclic wet-dry test. Each cycle includes wetting for 18 h through immersion in a 5% NaCl solution and drying for 6 h at 60% RH, both at 50 °C. Concrete powders were drilled in the compressive and tensile zone of beams (see Fig. 2.7). In each position, the concrete powders collected from three holes were mixed. Eight samples were drilled along the depth from 0 to 50 mm. The powder weight of 15 g was used to analyze the chloride content with each sample by automatic potentiometric titration.

Fig. 2.7 Powder drilling in beams; a Maps of drilling positions in a beam; b Beams after drilling

2.3 Effect of Flexural Loading on Chloride Ingress

21

2.3.2 Influence of Flexural Load on Saturated Concrete

Chloride Content (%wt. of concrete)

As shown in Fig. 2.8, the influence of loading on chloride diffusion is more significant to OPC concrete than GGBS concrete. The penetration depths of the chloride threshold concentration (taken as 0.1%) for GGBS concrete specimens in the compressive and tensile zones are about 11 mm and 9–12.5 mm, respectively, which are smaller than those of OPC-counterparts. Owing to the incorporation of GBFS, the chloride binding capacity of concrete is improved. The chloride distribution profiles show a slight difference among specimens under different loading levels but vary noticeably with the change of strain. This finding suggests that strain is a more appropriate indicator than the loading level. The chloride diffusion coefficients of saturated concrete can be obtained by fitting the test data of chloride concentration profiles with Fick’s second law. As shown in 0.35 OPC-IV-Tensile OPC-IV-Compressive OPC-III-Tensile OPC-III-Compressive OPC-II-Tensile OPC-II-Compressive OPC-I-Tensile OPC-I-Compressive Ref.OPC

0.30 0.25 0.20 0.15 0.10 0.05 0.00

0

10

20

30

40

50

Depth (mm)

Chloride Content (%wt. of concrete)

(a) 0.55 0.50 0.45 0.40 0.35 0.30 0.25 0.20 0.15 0.10 0.05 0.00

GGBS-IV-Tensile GGBS-IV-Compressive GGBS-III-Tensile GGBS-III-Compressive GGBS-II-Tensile GGBS-II-Compressive GGBS-I-Tensile GGBS-I-Compressive Ref.GGBS

0

10

20

30

40

50

Depth (mm)

(b) Fig. 2.8 Chloride distribution in tensile and compressive zones of pure bending (midspan) section of the saturated; a OPC concrete; b GGBS concrete

22

2 Chloride Ingress in Stressed Concrete 8 OPC-Compressive OPC-Tensile GGBS-Compressive GGBS-Tensile

7

5

D Cl 10

-11

2

(m /s)

6

4 3 2 1 0

0

10%

16%

22%

28%

Load level of RC beams (w.r.t. loading capacity )

Fig. 2.9 Chloride diffusion coefficient in different concrete specimens

(a)

(b) Fig. 2.10 The relationship between chloride diffusion coefficient and strain level of stressed concrete; a Tensile strain; b Compressive strain

2.3 Effect of Flexural Loading on Chloride Ingress

23

Fig. 2.9, the chloride diffusion coefficient of OPC concrete is higher than that in GGBS concrete. Combining the chloride diffusion coefficient data in Fig. 2.9 and the measured strain value, the relationship between chloride diffusion coefficient and strain is presented in Fig. 2.10. It can be seen that the change in chloride diffusion coefficient and strain value is consistent, which implies that the strain value is a reliable variable in terms of the evaluation of loading effects on chloride diffusion.

2.3.3 Influence of Flexural Load on Non-saturated Concrete Figure 2.11 shows the comparison of chloride profiles in tensile zones and compressive zones at different levels of loads subjected to wet-dry conditions. At a given

Chloride Content (%wt. of concrete)

0.7 10 % fc 16 % fc 22 % fc 28 % fc

0.6 0.5 0.4 0.3 0.2 0.1 0.0

0

10

20

30

40

50

Depth from the exposure surface(mm)

(a) 0.7

Chloride Content (%wt. of concrete)

Fig. 2.11 a Chloride profiles in the compressive zone of pure bending section subjected to cyclic wet-dry; b Chloride profiles in the tensile zone of pure bending section subjected to cyclic wet-dry

10 % fc 16 % fc 22 % fc 28 % fc

0.6 0.5 0.4 0.3 0.2 0.1 0.0

0

10

20

30

40

Depth from the exposure surface(mm)

(b)

50

24

2 Chloride Ingress in Stressed Concrete

level of loading, the chloride concentration at the tensile zone is considerably higher than that in the compressive zone. In the compressive zones, the measured strain has a relationship ε22% > ε28% > ε10% > ε16% , indicating that the chloride concentration tends to be larger with the increase of strains, but not monotonic. It implies that the chloride diffusion coefficient of compressively stressed concrete under a certain strain level is reduced, which may be explained by closure of the micro-cracks under compression. On the other hand, in tensile zones of the pure bending section, this trend is inconspicuous. The tensile strain triggers the propagation of micro-cracks in the weak zone, drastically enlarging the diffusion coefficient by creating a shorter path.

2.4 Fatigue Loads on Moisture Transport 2.4.1 Surface Factor The surface factor is one of the most widely used coefficients to predict the water evaporation at the concrete surface (Akita et al. 1997; Sakata 1983; Shimomurat and Maekawa 1997; Wong et al. 2001). The surface factor is a convective moisture transfer coefficient between air and concrete surface, varying dependent on temperature and relative humidity (RH) at the surface of concrete, wind speed, roughness of concrete surface, and diffusion coefficient of vapor in the air (Almusallam 2001; Bao and Wang 2017; Cussler 1984; Du et al. 2016; Fuller et al. 1966; Ghasemzadeh and Pour-Ghaz 2015; Samaha and Hover 1992; Welty et al. 2008; Yang et al. 2006). The formation of additional void or cracks at the surface of stressed concrete due to mechanical loading would affect the moisture transport at the surface layer of concrete. For instance, the openings of new micro-cracks at the concrete surface may increase the quantity of moisture being lost when exposed to a low RH condition.

2.4.2 Damage Variable of Fatigue-Damaged Concrete Fatigue is defined as the degradation of mechanical properties leading to failure of material or its component under cyclic loading (Meyers and Chawla 2008; Sabir et al. 1998). Continuum damage mechanics give a better understanding of the fatigue damage process by defining a damage variable which represents the deterioration of materials before the crack initiation. Lemaitre et al. described eight different methods to measure damage defined as the effective surface density of micro-cracks and cavities in any plane of a representative volume element (Lemaitre and Dufailly 1987). Damage variable related to the residual strain has also been successfully used to predict cyclic damage evolution (Cheng and Plumtree 1998; Guang-Xu et al. 1994).

2.4 Fatigue Loads on Moisture Transport

25

The damage evolution of concrete was tested to correlate the residual strain variation to the fatigue damage behavior (Li et al. 2011; Xiao et al. 2013). Thus, the damage variable of concrete damaged by uniaxial tensile fatigue loading can be expressed as Dc = εr /εrc

(2.16)

where εrc is the fracture strain of the uncycled concrete and εr is the residual strain after n cyclic loadings.

2.4.3 Experimental Programs Five concrete mixtures were adopted as shown in Table 2.1. Mixtures PC23, PC35, and PC47 denote that the binder consists of OPC with water-to-binder (w/b) of 0.23, 0.35, and 0.47, respectively. Mixtures SL47 and FA47 are obtained by replacing OPC with GGBS and FA with a mass-based replacement ratio of 40% and 30%, respectively. The dimension of the specimens subjected to tensile fatigue loads was 120 mm × 120 mm × 1200 mm (Fig. 2.12). For each specimen, four HRB 335 bars with a diameter of 12 mm and a length of 1400 mm were placed inside the concrete, with the thickness of concrete cover being 20 mm for each side. At the two ends of each specimen, the steel bars with a length of 75 mm stick out. In total, 13 reinforced specimens were prepared (i.e., 7 for PC mixture, 3 for SL mixture, and 3 for FA mixture), as shown in Table 2.2. In addition, 30 150 mm cubic specimens were cast and covered with plastic sheets. After 24 h, all specimens were demolded and cured in a moist room (i.e., 20 °C and 90% RH). The measured 28-day compressive and tensile strength of the cubic specimens are included in Table 2.1. The overview of the implemented tensile fatigue loading system is shown in Fig. 2.13. The tensile fatigue loading was carried out using a MTS electro-hydraulic fatigue test machine after the specimens were moist cured for 28 days. The maximum Table 2.1 Mix proportion (kg/m3 ) and mechanical properties of concrete Mixture ID

w/b

OPC

GGBS

PC23

0.23

580

0

PC35

0.35

455

PC47

0.47

372

SL47

0.47

FA47

0.47

FA

Sand

Gravel

Water

28d compr. strength (MPa)

28d tensile strength (MPa)

0

704

1043

115

32.3

2.29

0

0

691

1138

145

51.8

4.08

0

0

698

1116

175

42.8

2.68

223

149

0

698

1116

175

37.1

2.23

260

0

112

698

1116

175

34.8

2.23

26

2 Chloride Ingress in Stressed Concrete Steel bar

120

Concrete

1200

100

100

120

(a)

(b) Fig. 2.12 a Configuration of specimens for uniaxial tension fatigue testing (unit: in mm) b Specimens after casting

Table 2.2 Fatigue loading parameters and specimen number Mixture

Specimen ID

w/b

Maximum load (kN)

Minimum load (kN)

Number of cycles

PC

PC23-0

0.23

0

0

0

0

PC23-1

0.23

15

6

50,000

5

PC35-0

0.35

0

0

0

0

PC35-1

0.35

15

6

50,000

5

PC47-0

0.47

0

0

0

0

PC47-1

0.47

15

6

50,000

5

PC47-2

0.47

20

9

50,000

5

SL47-0

0.47

0

0

0

0

SL47-1

0.47

15

6

50,000

5

SL47-2

0.47

20

9

50,000

5

FA47-0

0.47

0

0

0

0

FA47-1

0.47

15

6

50,000

5

FA47-2

0.47

20

9

50,000

5

SL

FA

Load frequency (Hz)

Note PC represents the mixtures with 100% OPC; SL represents the mixtures with GGBS; while FA represents the mixtures with fly ash. The number in the acronym represents the loading levels, e.g., PC23-1 means the PC23 mixture specimens were subjected to fatigue loads with maximum fatigue loads of 15 kN and minimum fatigue loads of 6 kN

2.4 Fatigue Loads on Moisture Transport

27

Fig. 2.13 Tensile fatigue loading system

load of MTS is 1000 kN. The tensile fatigue loading frequency, number of loading cycles, and loading amplitude were pre-set using control systems. Table 2.2 shows the fatigue loading parameters used for all specimens. The cyclic loading in the sinusoid waveform is given in Fig. 2.14 with a loading frequency of 5 Hz and number of cycles of 50,000. Two levels of maximum fatigue loads, i.e., 15 kN and 20 kN, were used, which approximately correspond to 25 and 30% of the ultimate tensile load of PC47 specimens. The corresponding minimum loads were 40 and 45% of the maximum loads. As for SL and FA mixture, the same level of the load was applied. In order to quantify the effect of tensile fatigue loading on the mass flux of moisture through the concrete surface, the strain-stress relationship of pure concrete needs to be obtained. The distribution of eight strain gauges was in the two opposite surfaces along the tensile direction for each specimen. After completing the fatigue loading tests, the residual strains were recorded. After the tensile fatigue loading tests, the specimens were saw-cut into cubic specimens with a side length of 120 mm. All cubic specimens were paraffin-coated except one surface, which ensures a one-dimensional mass transport. Then, they

28

2 Chloride Ingress in Stressed Concrete 25

Fig. 2.14 Tensile load time relation curve

Pattern 1:max 15 kN, min 6 kN, 5Hz Pattern 2:max 20 kN, min 9kN, 5Hz

Tensile load ( kN)

20

15

10

5

0 0.0

0.1

0.2

0.3

0.4

Time (s)

were immersed in containers filled with water for 7 days to ensure saturation. At the desired age, the cubic specimens were exposed to desorption on the top surface at a constant temperature and ambient RH in environmental chambers. The cubic specimens were exposed to various artificial environments as shown in Table 2.3, in which temperature varied from 293 to 313 K, wind speed varied from 1 to 4.25 m s−1 , while the ambient RH was constant at 40%. During testing, all specimens were weighed periodically. The mass changes of those specimens were the cause of the cumulative water desorption. A digital capacitance-type sensor that can measure RH and temperature at the same time was used. The scheme of drying experiment setup is shown in Fig. 2.15. The fan was used to control wind speed with the real-time speed measured by wind dial. The air deflector and the wind tunnel guarantee the stability of wind speed and smoothly through the surface of the specimen, as shown in Fig. 2.16.

Table 2.3 Environmental conditions Condition

Specimen ID

Temperature (K)

RH (%)

Wind speed (m s−1 )

A

FA47-0, FA47-1, FA47-2 SL47-0, SL47-1, SL47-2 PC47-0, PC47-1, PC47-2 PC35-0, PC35-1 PC23-0, PC23-1

293

40

1

B

PC47-0, PC47-1

303

40

1

C

PC47-0, PC47-1

313

40

1

D

PC47-0, PC47-1

293

40

2.5

F

PC47-0, PC47-1

293

40

4.25

2.4 Fatigue Loads on Moisture Transport

29

Fig. 2.15 Scheme of the setups for moisture desorption experiments

Fig. 2.16 a Specimens arrangement in the tunnel; b Temperature and humidity sensor; c Placed sensor; d Airflow generating and the wind tunnel; e Wind speed dial

2.4.4 Moisture Transport in Surface Concrete Figure 2.17a shows the cumulative water desorption against the square root of time. For the majority of samples tested, the relation between cumulative water desorption and the square root of time of exposure begins to deviate from regular linearity to another linearity with decreased slope. The results clearly show that the cumulative water desorption in concrete increases with the increasing w/b. At the same w/b, the cumulative water desorption for fatigue-damaged concrete is substantially higher, in comparison to sound concrete. On the premise of fully hydrated cement, the

30 0.7

2 PC23-0(R =0.997)

0.6

2 PC35-0(R =0.996) 2 PC47-0(R =0.999)

2

Moisture loss m(kg/m )

Fig. 2.17 Cumulative water desorption for concrete of unit square meters; a Effect of fatigue loading and w/b on the moisture loss; b Effect of fatigue loading level and SCMs on the moisture loss

2 Chloride Ingress in Stressed Concrete

0.5

2 PC23-1(R =0.981) 2 PC35-1(R =0.998)

0.4

2 PC47-1(R =0.999)

0.3 0.2 0.1 0.0

1.0

1.5

2.0 1/2

2.5

3.0

3.5

2.5

3.0

3.5

1/2

t (h )

(a) 0.7

2

2

SL47-1(R =0.998)

0.5

PC47-0(R =0.999)

2

2

Moisture loss m (kg/m )

PC47-1(R =0.999)

0.6

FA47-1(R =0.999) 2 2

PC47-2(R =0.999)

0.4 0.3 0.2 0.1 0.0

1.0

1.5

2.0 1/2

1/2

t (h )

(b) compactness of concrete is reduced with the increase of w/b ratio and fatigue damage. Accordingly, the resistance of moisture diffusion in concrete is reduced and the surface mass transfer area is enlarged. Figure 2.18a shows a decreasing rate of RH for surface concrete against time, in which the initial RH is 90%. With moisture desorption, the RH decreases rapidly during the early stage of the test. After 6 h exposure, the RH variation of unit time becomes similar for the fatigue damage concrete and sound concrete. Figures 2.17b and 2.18b show the cumulative water desorption and decrease of RH in surface concrete with the various binder types. In Fig. 2.17b, it can be seen that there is a small difference in water loss due to the replacement of OPC by GGBS or FA. At the same w/b, the incorporation of GGBS and FA has minor influence on desorption, at which stage the pozzolanic reaction may have started but still not sufficient to redeem the inverse effect (Liu et al. 2014). From Fig. 2.18b, the RH change rate decreased in concrete with SCMs, suggesting lower surface desorption, in comparison to OPC concrete (Mehta and Monteiro 2017).

2.4 Fatigue Loads on Moisture Transport 8

Decrease of relative humidity (%)

Fig. 2.18 Decrease of RH for surface concrete; a Effect of fatigue loading and w/b on the RH; b Effect of fatigue loading level and SCMs on the RH evolution

31

PC23-0

7

PC35-0

6

PC47-0 PC23-1

5

PC35-1

4

PC47-1

3 2 1 0

0

1

2

3

4

5 6 7 Time (h)

8

9

10 11

(a) Decrease relative humidity (%)

9 PC47-0

8

FA47-0

7

SL47-0

6

FA47-1

PC47-1 SL47-1

5

PC47-2

4

FA47-2 SL47-2

3 2 1 0

0

2

4

6 Time (h)

8

10

(b)

2.4.5 Influence Factors on the Surface Factor The surface factor of concrete under different wind speeds and temperatures was evaluated. The obtained test data were compared and plotted in Figs. 2.19 and 2.20. Figure 2.19 shows that the moisture convection velocity is accelerated by wind speed and surface factor increases with the wind speed. With the increase in wind speed, the influence in factor is gradually reduced. Figure 2.20 shows that the surface factor increases with temperature, likely due to the fact that the temperature variety changes the thermodynamic parameters of the fluid, such as water vapor diffusion coefficient, air density, and air viscosity. It is interesting to see that the surface factor increases linearly with the temperature. On the other hand, there is little effect of fatigue damage on the surface factor.

32

2 Chloride Ingress in Stressed Concrete 0.028

-1

f (m×s )

Fig. 2.19 Influence of wind speed on the surface factor

0.024

PC47-0 js=0.050 PC47-1 js=0.073

0.020

Analytical result

0.016 0.012 0.008 w/b = 0.47 Temperature = 20 °C Humidity = 40 %

0.004 0.000 0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

4.5

5.0

-1

u0 (m×s )

Fig. 2.20 Influence of temperature on the surface factor

0.014 0.012 0.010 0.008 0.006 0.004 0.002

PC47-0 js=0.0502 PC47-1 js=0.0582 Analytical result

0.000 270

280

290

w/c = 0.47 Wind speed = 1 m/s Humidity = 40 %

300

310

320

2.4.6 Surface Mass Transfer Area In order to quantify the effect of tensile fatigue damage on the surface mass transfer area of concrete, the correlation between the fitted surface mass transfer area ϕ0 and damage variable is established. The value of the damage variable Dc is in the range of 0.279–0.475, based on Eq. (2.16). As shown in Fig. 2.21, the magnification factor f d , defined as the ratio of the apparent initial surface mass transfer area ratio ϕ0 for damaged concrete to that of sound concrete at the same time, is approximately parabolically correlated to the measured average damage variable of residual strain. The value of f d is in the range of 1.05–1.60, implying that the tensile fatigue accelerates the initial surface mass transfer area by 1.05–1.60 times when the magnitude of maximum tensile fatigue load is 15 and 20 kN and w/b are between 0.23 and 0.47.

2.5 Fatigue Loads on Chloride Ingress 1.6

Magnification factor ( fd )

Fig. 2.21 Correlation between magnification factor and damage variable in tensile fatigue-damaged concrete

33

1.5 1.4

Derived fd for fitting 2

2

fd=2.37*Dc +1 (R =0.846)

1.3 1.2 1.1 1.0 0.0

0.1

0.2

0.3

0.4

0.5

Dc

2.5 Fatigue Loads on Chloride Ingress 2.5.1 Experimental Program The same kind of fatigue-damaged specimens was adopted for the chloride penetration tests. Table 2.4 shows the fatigue loading parameters used. In particular, for PC mixture, five different levels of maximum fatigue loads, namely 15, 18, 21, 24, and 27 kN, was used, approximately corresponding to the 25, 30, 35, 40, and 45% of the ultimate tensile load of the specimens (denoted as f t ). The ultimate tensile loads of the specimens were calculated to be around 62.3 kN based on the measured tensile strength of concrete, cross-sectional area, elastic modulus ratio between steel and concrete, and the ratio of reinforcement. For SL and FA mixture, only one level of load (a maximal fatigue load of 15 kN) was applied. For the specimens subjected Table 2.4 Fatigue loading protocol Mixture

Mixture IDa

Maximum load (kN)

PC

PC15

15

PC18

18

10.8

50,000

5

PC21

21

12.6

50,000

5

FA SL a The

Minimum load (kN) 6.75

Number of loading cycles

Loading frequency (Hz)

50,000

5

PC24

24

14.4

50,000

5

PC27

27

16.2

50,000

5

FA15

15

6.75

50,000

5

SL15

15

6.75

50,000

5

Note number after the mixture ID represents the maximum loads, for instance, PC27 means the PC mixture specimens which were subjected fatigue loads with maximum fatigue loads of 27 kN and minimum fatigue loads of 16.2 kN

34

2 Chloride Ingress in Stressed Concrete

Fig. 2.22 a The prepared cubic specimens for chloride penetration tests; b Illustration of the powder collection for determining the chloride content in concrete

to a maximum load of 15 kN, the corresponding minimum loads were 45% of the maximum loads, while for other levels of maximal loads, the minimum loads were 60% of the maximal loads. The loading protocol follows constant amplitude sinusoid loading regimes with a loading frequency of 5 Hz and a number of cycles of 50,000. After tensile fatigue loading tests, the specimens were saw-cut into a series of cubic specimens with a side length of 120 mm, as shown in Fig. 2.22a. Given the inhomogeneity of the concrete, the strain values along the longitudinal direction of each specimen would be slightly different. In order to accurately quantify the effect of fatigue damage on the chloride penetration in concrete, the residual strain for each cubic was collected. Therefore, each fatigue loading specimen can be cut into four cubic specimens. Two cubic specimens, corresponding to the strain gauges #1–#4, were used in the chloride penetration tests in saturated concrete (immersion condition), while the other two, corresponding to the strain gauges #5–#8, were for concrete subjected to cyclic wet-dry condition. Prior to chloride penetration tests, the five sides of the cubic specimen were coated with paraffin, exposing only one side (the side which was initially exposed to the atmosphere during fatigue loading test) to chloride attack. For immersion conditions, the specimens were immersed in a container filled with a 5% NaCl solution. To avoid moisture evaporation-induced concentration changes, the container was sealed and stored in a walk-in environmental chamber which was programmed to 20 °C. The chloride content as a function of depth in concrete was measured after immersing for 30, 45, and 60 days. For cyclic wet-dry condition, the specimens were placed in a container which can automatically adjust the water content. For each wet-dry cycle (24 h), the specimens were immersed in a 5% NaCl solution for 6 h, followed by drying at 40% RH in a walk-in environmental chamber for 18 h. The temperature was kept at 20 °C during the entire test. The chloride content as a function of depth in concrete was measured after 30, 45, and 60 cycles.

2.5 Fatigue Loads on Chloride Ingress

35

The concrete powders were drilled at the depths of 2.5, 7.5, 12.5, 17.5, 22.5, 27.5, 35.0, and 45.0 mm away for the exposed surface. At each depth, the concrete powders collected from three holes were mixed to avoid the influence of aggregates, as shown in Fig. 2.22b. The powder with a weight of 10 g (passing through 0.25 mm sieve) was oven-dried at 105 ± 5 °C for two hours before cooling down and conducting the chloride content measurements. The free chloride (water-soluble) concentration was measured, as expressed as the weight percentage of the concrete powders.

2.5.2 Chloride Profiles in Fatigue-Damaged Concrete Figure 2.23 shows the chloride profiles in saturated concrete after immersing for 30, 45, and 60 days. It is reasonable to see that the chloride content in concrete increases with the increasing exposure duration. At the same age, the chloride content at a certain depth for fatigue-damaged concrete is substantially higher, in comparison to that in sound concrete, and tends to increase as the magnitude of loads increases. For instance, at 30 days, the fatigue-damaged concrete has the chloride content of 15–76% higher than that of sound concrete. In addition, when the magnitude of maximum fatigue loads reaches beyond 30–35% f t of the specimen, there is a conspicuous increase of the chloride content. In the range of 25–30% f t , there is no substantial difference in chloride content, which probably suggests that the 30–35% f t is a critical range for the coalescence of micro-cracks in concrete subjected to fatigue loads. When subjected to cyclic wet-dry condition, as shown in Fig. 2.24, the chloride content at the same exposure duration for each mixture is dramatically increased, as compared to that in saturated status. This observation suggests that cyclic wet-dry considerably accelerates chloride penetration in concrete, regardless it is fatiguedamaged or not. The accelerating mechanism of wet-dry cycles is attributed to the capillary suction of drying-induced highly concentrated NaCl solution at the exposed surface during a subsequent wetting period.

2.5.3 Fatigue Loading Effect Factor Based on Fick’s second law, the apparent chloride diffusion coefficient, Dapp,c , for each specimen can be fitted according to the measured free chloride profiles in concrete. Figure 2.25 shows the fitted values for apparent chloride diffusion coefficient for PC mixture as a function of exposure condition (i.e., immersion or wet-dry cycle), duration, and degree of fatigue damage. It can be seen that the apparent chloride diffusion coefficient decreases as a function of exposure duration, regardless of whether it was exposed to immersion or cyclic wet-dry conditions. However, for concrete subjected to wet-dry cycles, the time-dependent evolution of the apparent chloride

36

Free chloride content (weight% concrete)

1.6

Immersion condition

1.4

PC15-1,30d PC18-1,30d PC21-1,30d PC24-1,30d PC27-1,30d PC, 30d

1.2 1.0 0.8 0.6 0.4 0.2 0.0

0

5

10

15

20

25

30

35

40

45

50

Depth (mm)

(a) Free chloride content (weight % concrete)

1.6 Immersion condition

1.4

PC15-2,45d PCO18-2,45d PC21-2,45d PC24-2,45d PC27-2,45d PC, 45d

1.2 1.0 0.8 0.6 0.4 0.2 0.0

0

5

10

15

20

25

30

35

40

45

50

Depth (mm)

(b) 1.6

Free chloride content (weight % concrete)

Fig. 2.23 Free chloride profiles in PC mixture concrete subjected to various levels of fatigue loads in immersion at different ages; a 30 days; b 45 days; c 60 days (Note the number between mixture ID and the exposure duration is the strain gauge number; for example, PC15-2–45 days represents the cubic specimen for chloride penetration tests is cut from the location of strain gauge #2 in PC15 at 45 days)

2 Chloride Ingress in Stressed Concrete

Immersion condition

1.4

PC15-3,60d PC18-3,60d PC21-3,60d PC24-3,60d PC27-3,60d PC, 60d

1.2 1.0 0.8 0.6 0.4 0.2 0.0

0

5

10

15

20

25

30

Depth (mm)

(c)

35

40

45

50

2.5 Fatigue Loads on Chloride Ingress

Free chloride content (weight % concrete)

1.6 Drying-wetting cycles

PC15-5,30d PC18-5,30d PC21-5,30d PC24-5,30d PC27-5,30d PC, 30d

1.4 1.2 1.0 0.8 0.6 0.4 0.2 0.0

0

5

10

15

20

25

30

35

40

45

50

Depth (mm)

(a)

Free chloride content (weight % concrete)

1.6 Drying-wetting cycles

1.4

PC15-6,45d PC18-6,45d PC21-6,45d PC24-6,45d PC27-6,45d PC, 45d

1.2 1.0 0.8 0.6 0.4 0.2 0.0

0

5

10

15

20

25

30

35

40

45

50

Depth (mm)

(b)

1.6

Free chloride content (weight % concrete)

Fig. 2.24 Free chloride profiles in PC mixture concrete subjected to various levels of fatigue loads in wet-dry condition at different ages; a 30 days; b 45 days; c 60 days

37

PC15-7,60d PC18-7,60d PC21-7,60d PC24-7,60d PC27-7,60d PC, 60d

Drying-wetting cycles

1.4 1.2 1.0 0.8 0.6 0.4 0.2 0.0

0

5

10

15

20

25

30

Depth (mm)

(c)

35

40

45

50

2

Fig. 2.25 The fitted values for apparent chloride diffusion coefficient for concrete subjected to various levels of fatigue loads; a Immersion; b Wet-dry cycles

2 Chloride Ingress in Stressed Concrete

Apparent chloride diffusivity (m /s)

38 1.8x10

-10

1.6x10

-10

1.4x10

-10

1.2x10

-10

1.0x10

-10

8.0x10

-11

6.0x10

-11

4.0x10

-11

2.0x10

-11

0.0

30d 45d 60d

PC0

PC15

Immersion condition

PC18

PC21

PC24

PC27

PC24

PC27

Mixture

2

Apparent chloride diffusivity (m /s)

(a) 1.8x10

-10

1.6x10

-10

1.4x10

-10

1.2x10

-10

1.0x10

-10

8.0x10

-11

6.0x10

-11

4.0x10

-11

2.0x10

-11

0.0

30d 45d 60d

PC0

PC15

Drying-wetting cycles

PC18 PC21 Mixture

(b) diffusion coefficient is relatively strong, in comparison to that in immersion condition. This is reasonable since, for saturated concrete under immersion, diffusion is the main transport mechanism for chloride penetration. However, for unsaturated concrete, multi-mechanisms, including diffusion, convection, and capillary absorption, work jointly to facilitate the chloride ingress. Therefore, at the surface layer of concrete, the apparent chloride diffusion coefficient at an early age may be unstable but tends to become stable as time goes, since the chloride diffusion progressively dominates the transport process at a deeper depth. On the other hand, the apparent chloride diffusion coefficient of concrete subjected to wet-dry cycles is about 1.5–2.0 times larger than that of the same mixture subjected to immersion conditions. The difference of apparently chloride diffusion coefficient of concrete when subjected to these two conditions seems to increase as the degree of fatigue damage increases, which may indicate that fatigue damage is more harmful to concrete subjected to wet-dry cycles.

2.5 Fatigue Loads on Chloride Ingress 5.0 Immersion_30d Immersion_45d Immersion_60d Drying-wetting cycles_30d Drying-wetting cycles_45d Drying-wetting cycles_60d

4.5 Magnification factor ( fd )

Fig. 2.26 Correlation between magnification factor and residual strain in fatigue-damaged concrete

39

4.0 3.5

2

3.0

fd=0.01421*(residual strain)+1 (R =0.75)

2.5 2.0 1.5 1.0

0

20

40

60

80

100

120

Resiual strain ( ε )

In order to quantify the effects of tensile fatigue damage on the apparent chloride diffusion coefficient of concrete, the correlation between the fitted apparent chloride diffusion coefficient and measured residual strain is established. As shown in Fig. 2.26, the fatigue loading effect factor f d , which is defined as the ratio of apparent chloride diffusion coefficient for damaged concrete to that of sound concrete at the same age, is approximately linearly correlated to the measured residual strain. The value of f d is in the range of 1.5–3.0, implying that fatigue damage accelerates the chloride penetration by 1.5–3.0 times when the magnitude of maximum fatigue load is between 25 and 45% f t .

2.6 Conclusions In this chapter, the influence of flexural and fatigue damage on the moisture and chloride ingress into concrete subjected to immersion or cyclic wet-dry actions is outlined. This chapter highlights the importance of defining appropriate damage factor to quantify the loading effects. The following conclusion can be drawn. (1) It is the local strain rather than the overall applied load that reflects the chloride diffusion coefficient in stressed concrete under flexural and fatigue loads. Hence, in order to investigate the loading effects on chloride diffusion in concrete, the strain-based method is more reliable than the load-based method. (2) The chloride diffusion coefficient increases with the increase in tensile strain. The generated concrete damage accelerates the increasing rate of the diffusion coefficient. On the contrary, the chloride diffusion coefficient reduces with the increasing compressive strain within a certain range.

40

2 Chloride Ingress in Stressed Concrete

(3) Within the magnitude of fatigue loads studied, the fatigue damage accelerates the chloride diffusion coefficient of concrete by a factor of 1.5–3.0 and surface moisture transfer by a factor of 1.05–1.60 in the initial desorption process. (4) The apparent chloride diffusion coefficient of concrete subjected to wet-dry cycles is about 1.5–2.0 times higher than that under immersion. Fatigue damage is more harmful to concrete subjected to wet-dry cycles. (5) The surface factor of drying concrete increases as the temperature and wind speed increases. The surface factor of concrete is not considerably influenced by fatigue damage. (6) The rate of moisture loss from concrete surface increases with the increasing level of tensile fatigue damage of concrete, when the magnitude of fatigue loads level is below 30% of the ultimate tensile load. Nevertheless, the initial porosity of concrete is the controlling factor influencing the moisture transfer of concrete.

References Akita, H., Fujiwara, T., & Ozaka, Y. (1997). A practical procedure for the analysis of moisture transfer within concrete due to drying. Magazine of Concrete Research, 48(6), 129–137. Almusallam, A. A. (2001). Effect of environmental conditions on the properties of fresh and hardened concrete. Cement and Concrete Composites, 23(4–5), 353–361. Bao, J., & Wang, L. (2017). Effect of short-term sustained uniaxial loadings on water absorption of concrete. Journal of Materials in Civil Engineering, 29(3), 4016234. van Brakel, J., & Heertjes, P. M. (1974). Analysis of diffusion in macroporous media in terms of a porosity, a tortuosity and a constrictivity factor. International Journal of Heat and Mass Transfer, 17(9), 1093–1103. Cheng, G., & Plumtree, A. (1998). A fatigue damage accumulation model based on continuum damage mechanics and ductility exhaustion. International Journal of Fatigue, 20(7), 495–501. Cussler, E. (1984). Diffusion, mass transfer in fluid system. Du, M., Jin, X., Ye, H., Jin, N., & Tian, Y. (2016). A coupled hygro-thermal model of early-age concrete based on micro-pore structure evolution. Construction and Building Materials, 111, 689–698. Fu, C., Jin, X., Ye, H., & Jin, N. (2015). Theoretical and experimental investigation of loading effects on chloride diffusion in saturated concrete. Journal of Advanced Concrete Technology, 13(1), 30–43. Fuller, E., Schettler, P., & Giddings, J. (1966). A new method for prediction of binary gas-phase coefficients. Industrial & Engineering Chemistry, 58. Gerard, B., Pijaudier-Cabot, G., & Laborderie, C. (1998). Coupled diffusion-damage modelling and the implications on failure due to strain localisation. International Journal of Solids and Structures, 35(31–32), 4107–4120. Ghasemzadeh, F., & Pour-Ghaz, M. (2015). Effect of damage on moisture transport in concrete. Journal of Materials in Civil Engineering, 27(9). Gowripalan, N., Sirivivatnanon, V., & Lim, C. C. (2000). Chloride diffusivity of concrete cracked in flexure. Cement and Concrete Research, 30(5), 725–730. Guang-Xu, C., Zhi-Wen, L., & Zhen-Bang, K. (1994). A new damage variable for low-cycle fatigue of metallic materials. Engineering fracture mechanics, 48(2), 281–287. Hoseini, M., Bindiganavile, V., & Banthia, N. (2009). The effect of mechanical stress on permeability of concrete: A review. Cement and Concrete Composites, 31(4), 213–220.

References

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Lemaitre, J., & Dufailly, J. (1987). Damage measurements. Engineering Fracture Mechanics, 28(5– 6), 643–661. Li, W., Sun, W., Jiang, J. (2011). Damage of concrete experiencing flexural fatigue load and closed freeze/thaw cycles simultaneously. Construction and Building Materials, 25(5), 2604–2610. Liu, J., Xing, F., Dong, B., Ma, H., & Pan, D. (2014). Study on water sorptivity of the surface layer of concrete. Materials and Structures, 47(11), 1941–1951. Maekawa, K., Ishida, T., & Kishi, T. (2003). Multi-scale modeling of concrete performance integrated material and structural mechanics. Journal of Advanced Concrete Technology, 1(2), 91–126. Mehta, P. K., & Monteiro, P. J. M. (2017). Concrete microstructure, properties and materials. Meyers, M. A., & Chawla, K. K. (2008). Mechanical behavior of materials. Cambridge university press. Sabir, B. B., Wild, S., & O’Farrell, M. (1998). A water sorptivity test for martar and concrete. 31(8), 568–574. Sakata, K. (1983). A study on moisture diffusion in drying and drying shrinkage of concrete. Cement and Concrete Research, 13(2), 216–224. Samaha, H. R., & Hover, K. C. (1992). Influence of microcracking on the mass transport properties of concrete. Materials Journal, 89(4), 416–424. Shimomurat, T., & Maekawa, K. (1997). Analysis of the drying shrinkage behaviour of concrete using a micromechanical model based on the micropore structure of concrete. Magazine of Concrete Research, 49(181), 303–322. Welty, J., Wicks, C. E., Rorrer, G. L., & Wilson, R. E. (2008). Fundamentals of momentum, heat and mass transfer (5th ed.). Wong, S. F., Wee, T. H., Swaddiwudhipong, S., & Lee, S. L. (2001). Study of water movement in concrete. Magazine of Concrete Research, 53(3), 205–220. Xiao, J., Li, H., & Yang, Z. (2013). Fatigue behavior of recycled aggregate concrete under compression and bending cyclic loadings. Construction and Building Materials, 38, 681–688. Yang, Z., Weiss, W. J., & Olek, J. (2006). Water transport in concrete damaged by tensile loading and freeze–thaw cycling. Journal of materials in civil engineering, 18(3), 424–434.

Chapter 3

Chloride Ingress in Cracked Concrete

Abstract This chapter focuses on the influence of macroscopic cracks on the moisture transfer and chloride ingress in pre-cracked concrete. The influence of geometrical characteristics of cracks on the moisture influential depth and chloride diffusion coefficient of concrete is discussed.

3.1 Introduction For concrete with macro-cracks, the crack patterns and characteristic parameters (e.g., crack width, length, roughness, and geometry) are accessible and measurable. As such, establishing the relationship between the crack parameters and their impacts on the transport properties of cracked concrete is intriguing.

3.2 Moisture Influential Depth The moisture state within concrete exposed to a natural environment is in a nonequilibrium condition, following the changes in external conditions (Andrade et al. 1999). Therefore, it is reasonable to assume that the moisture state in concrete cracks follows the change of external environments as well. However, due to the variation of environmental changing frequency, the moisture influential depth may be different for concrete with various crack patterns.

3.2.1 Experimental Program Figure 3.1 shows the configuration of the pre-cracked concrete specimen with a rectangular cross section of 150 mm × 75 mm and a span of 400 mm. The specimens are prefabricated with cracks of natural shape with designated width and depth in the © The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2020 H. Ye et al., Chloride-Induced Steel Corrosion in Concrete Under Service Loads, https://doi.org/10.1007/978-981-15-4108-7_3

43

44

Fig. 3.1 Configuration of prefabricated specimens

3 Chloride Ingress in Cracked Concrete

3.2 Moisture Influential Depth Table 3.1 Crack characteristics of prefabricated concrete specimens

45 Crack depth (mm)

Crack width (mm) Specimen 1

Specimen 2

Specimen 3

90–105

0.02

0.10

0.12

75–90

0.04

0.12

0.16

60–75

0.06

0.16

0.20

45–60

0.08

0.16

0.20

30–45

0.10

0.20

0.28

15–30

0.10

0.20

0.30

0–15 (reserved crack)

1.10

1.10

1.10

designated direction. The details of the crack pattern along the crack path are shown in Table 3.1. The specimens were dried for 15 days in an environment chamber at 20 ± 3 °C and 60% RH, in order to eliminate the influence of internal moisture variability on the measured RH in the cracked zone. The RH and temperature sensors were waterproofed using a vapor-permeable cover to ensure the accuracy of measurement in high humid conditions (Fig. 3.2). The environment chamber was capable of simulating rainfall and fog environments using the shower nozzle as shown in Fig. 3.3. Five kinds of environmental conditions were simulated, including periodic RH change at constant temperature, periodic temperature change, artificial fog, artificial rainfall, and the drying period after rainfall precipitation. For each RH cycle, external RH changes from 95% to 70%, and then back to 95% at 5% interval. For each temperature cycle, it changes from 40 °C to 15 °C, then back to 40 °C at 5 °C interval. For the drying period after rainfall precipitation, the external atmospheric condition maintains at 25 ± 3 °C and 40% RH. In order to monitor the external RH and temperature, a sensor was placed inside the environmental chamber.

Fig. 3.2 Arrangement of sensors along the crack path of cracked specimens

46

3 Chloride Ingress in Cracked Concrete

Fig. 3.3 Arrangement of specimens in the environmental chamber

3.2.2 Cyclic RH Change Figure 3.4 shows the time-dependent evolution of RH in cracks subjected to a cyclic change of external RH. For each cycle, the moisture status in concrete is synchronous with the wet-dry cycles. It shows that the RHs in different cracked concretes under the same periodic environmental change appear to be different and influenced by the crack width. The RH change amplitude of the points far from the surface is smaller than the points near to the surface. Among specimens, the RH at almost all points tends to be higher with a wider crack width. The phenomena are caused by the fact that moisture adsorption/desorption of concrete with a larger crack width is faster than that with a smaller crack width since the larger crack width has a more open space with a much larger interactive area and faster exchange rate with the external environment. The moisture influential depth increases as crack width increases; however, it is unable to predict the accurate value of moisture influential depth, as the results depend on the exact wet-dry cycles and exposure duration. In a real atmospheric environment, the temperature changes periodically as well, which leads to uncertainty in determining the moisture influential depth. However, it is reasonable to propose a possible range of influential depth of moisture for cracked concrete with different crack widths. It assumes that the point with the longest distance from the exposed surface (XY plane), which changes sensitively with the external periodic environments, is the moisture influential depth. Nevertheless, this distance needs to be subtracted by the reversed crack depth which is 15 mm as shown in Tables 3.2 and 3.1. In the specimens with 0.1 mm surface crack, the approximately moisture influential depth is 30 mm. Furthermore, it needs to take account of the errors of sensor installation which is about 10 mm. As such, for concrete with surface crack widths of 0.1, 0.2, and 0.3 mm, the corresponding moisture influential depth is 25–35 mm, 50–60 mm, and 65–75 mm, respectively.

3.2 Moisture Influential Depth 100

At the depth of 100mm

At the depth of 85mm

Relative Himidity (%)

95 90

At the depth of 70mm

At the depth of 55mm

85 80 75 70

External humidity at constant temperature

65

At the depth of 40mm

crack width 0.1mm at surface specimens 1

60 0

6 Time (day)

3

9

(a) 100

At depth of 85mm At depth of 100mm

Relative Humidity (%)

95 90

At depth of 70mm

At depth of 55mm

85 80 75

At depth of 40mm

70 External Humidity at constant temperature

65 0

3

crack width 0.2mm at surface specimen 2

6 Time (day)

9

(b) At depth of 85mm

At depth of 55mm

100

At depth of 70mm

95

Relative Humidity (%)

Fig. 3.4 Time-dependent changing of RH in cracks of concrete subjected to a cyclic change of external artificial RH. a specimen 1; b specimen 2; c specimens 3

47

90 85 At depth of 25mm

80

At depth of 40mm

75 70

External Humidity at constant temperature

65 0

3

crack width 0.3mm at surface specimen 3

6 Time (day)

(c)

9

48

3 Chloride Ingress in Cracked Concrete

Table 3.2 Crack widths of specimens Crack depth(mm)

Crack width (mm) (OPC concrete) 0.05

Crack width (mm) (GGBS concrete)

0.10

0.15

0.20

0.02

0.02

115–135

0.05

0.10

0.15

0.20

0.04

0.02

0.02

0.01

0.02

0.02

95–115

0.02

75–95

0.02

0.02

0.04

0.09

0.03

0.04

0.04

0.10

55–75

0.03

0.06

0.10

0.15

0.04

0.07

0.10

0.14

35–55

0.05

0.08

0.15

0.20

0.05

0.08

0.14

0.20

15–35

0.05

0.10

0.15

0.20

0.05

0.10

0.15

0.20

0–15 (reserved crack)

0.85

0.90

0.95

1.00

0.85

0.90

0.95

1.00

3.2.3 Cyclic Temperature Change Figure 3.5 shows the time-dependent change of RH in cracks subjected to a cyclic change of temperature. In the beginning, the RHs in both the environment and concrete crack are at a high level. The highest atmospheric temperature approximately corresponds to the lowest atmospheric RHs in cracks. The initial RH in the crack path is high; however, with the change of external temperature as well as external RH, the RHs of all points in specimen 3 get much closer to external RH than those in specimen 2. It can be well explained by the same reasons that larger crack width is affected by the external environment easier and the corresponding moisture influential depth is larger.

3.2.4 Fog Environment Figure 3.6 shows the time-dependent change of RH in cracks subjected to artificial fog at a constant temperature. It can be seen that the RH at the external environment is generally smaller than that in cracks. The RHs in the external environment and crack increase quickly and then keep almost unchanged. There is no significant difference between the three specimens since the RHs are all at a high level.

3.2.5 Rainfall Environment Figure 3.7 shows the time-dependent change of RH in cracks subjected to artificial rainfall and the subsequent drying. During the rainfall period, environmental RH increases rapidly. Then, the RHs in both concrete crack and external environment keep approximately high at 100% RH. During the drying period, the RHs in external

3.2 Moisture Influential Depth 100

at depth of 100mm

at depth of 85mm

90 at depth of 70mm

Relative Humidity (%)

80 70

at depth of 40mm

60

at depth of 55mm

50 40

External Humidity with the changing temperature

30

External Temperature

20 10

0

crack width 0.1 at surface specimen 1

3 Time (day)

(a) at depth of 85mm

Relative Humidity(%)

100

at depth of 100mm

80

at depth of 70mm

60

at depth of 55mm External Humidity with changing temperature

at depth of 40mm

40 External Temperature

20 0

crack width 0.2 at surface specimen 2

3 Time (day)

(b) 100

Relative Humidity (%)

Fig. 3.5 Time-dependent RH change in cracks of concrete subjected to a cyclic change of external artificial temperature. a specimen 1; b specimen 2; c specimen 3

49

at depth of 85mm

at depth of 70mm

at depth of 55mm

80 at depth of 25mm

60 External Humidity with changing temperature

at depth of 40mm

40 External Temperature

20 0

crack width 0.3 at surface specimen 3

3 Time (day)

(c)

50 100

at depth of 85mm

at depth of 100mm

Relative Humidity (%)

90 80

at depth of 55mm

70

at depth of 70mm

at depth of 40mm

External Humidity

60 50 40

20

crack width 0.1 at surface specimen 1

Exrernal Temperature

30 0

1

Time (day)

(a) 100

at depth of 55mm

at depth of 100mm

Relative Humidity (%)

90 80

at depth of 85mm

at depth of 70mm

70

External Humidity

60

at depth of 40mm

50 40 Exrernal Temperature

30 20

crack width 0.2 mm at surface specimen 2

0

1

Time (day)

(b) 100

at depth of 85mm

90

Relative Humidity (%)

Fig. 3.6 Time-dependent changing of RH in cracks of concrete subjected to simulated fog environment. a specimen 1; b specimen 2; c specimens 3

3 Chloride Ingress in Cracked Concrete

at depth of 55mm

at depth of 70mm

at depth of 25mm

80 70

External Humidity

60

at depth of 40mm

50 40 Exrernal Temperature

30 20

0

crack width 0.3 at surface specimen 3

Time (day)

(c)

1

3.2 Moisture Influential Depth

100

at depth of 85mm

Relative Humidity (%)

90

at depth of 100mm at depth of 55mm

80 at depth of 40mm

at depth of 75mm

70

External Humidity

60 50 40

External Temperature

30 20

0

1

crack width 0.1 at surface specimen 1

2

Time (day)

(a) at depth of 100mm

100

at depth of 85mm

Relative Humidity (%)

90

at depth of 40mm

at depth of 70mm

80

at depth of 55mm

70 60

External Humidity

50 40

20

crack width 0.1 at surface specimen 2

External Temperature

30 0

1

Time (day)

2

(b) at depth of 55mm

100

at depth of 70mm

90

Relative Humidity (%)

Fig. 3.7 Time-dependent changing of RH in cracks of concrete subjected to artificial rainfall and the subsequent drying. a specimen 1; b specimen 2; c specimens 3

51

80

at depth of 25mm at depth of 40mm

70 at depth of 85mm

60

External Humidity

50 40 30

crack width 0.3 at surface specimen 3

External Temperature

20 0

1

Time (day)

(c)

2

52

3 Chloride Ingress in Cracked Concrete

environments and crack change relatively slow. It shows that the external environmental RH is much sensitive to the drying and decreases more than that in cracks. In addition, this phenomenon is much pronounced for the specimen with larger surface crack width and the points closer to the exposure surface. During the drying phase, the liquid phase of moisture spontaneously vaporizes from the surface of the concrete. The amount of such evaporation progressively increases when the crack has a larger interface area with the external atmospheric environment, while decreases as the decreasing supply of moisture from the inside of the sound concrete zone.

3.3 Chloride Diffusion in Cracked Concrete in Saturated States 3.3.1 Experimental Program The same kind of cracked concrete specimens was adopted. The specimens were subjected to accelerated chloride penetration under immersion in a 5.0% NaCl solution at 25 ± 3 °C for 30 and 60 days. The specimens with surface crack width of 0.05, 0.10, 0.15, and 0.20 were prepared both for OPC concrete and GGBS concrete. The details of crack patterns along the crack path at the XZ plane are shown in Table 3.2, where the crack width is the width along the crack path at different crack depths.

3.3.2 Chloride Diffusion Coefficient Figure 3.8 shows the electron probe micro-analysis (EMPA) of chloride profiles for both OPC and GGBS concrete subjected to immersion for 30 days and 60 days. For both types of concrete, chloride concentration in cracked concrete increases with immersion ages, although the existence of crack geometrically disturbs the chloride ingress path and diverges the chloride concentration by transporting vertically into the crack surface. On the other hand, the aggregate embedded at the crack surface affects chloride profiles to a large extent. The aggregate at the cracked zone significantly affects chloride ingress and chloride profiles by increasing the tortuosity of the transport path. Alternatively, the presence of aggregate particles may locally facilitate the chloride movement due to the existence of interfacial transition zone (ITZ) around aggregates. The EMPA line scanning results of chloride concentration for GGBS concrete are shown in Fig. 3.9, and the EMPA point scanning results for OPC and GGBS concrete with different crack widths are shown in Fig. 3.10. It can be seen that the chloride concentration at the crack surface tends to increase with the increase of crack width since more chloride ions can be carried through the wider openings. The chloride

3.3 Chloride Diffusion in Cracked Concrete in Saturated States

53

Fig. 3.8 EMPA mapping of chloride concentration of a OPC concrete subjected to immersion for 30 days b OPC concrete subjected to immersion for 60 days c GGBS concrete subjected to immersion for 30 days d GGBS concrete subjected to immersion for 60 days

concentration of OPC concrete is larger than that in GGBS concrete at the same depth with the same crack width as the GGBS refines pore structures of concrete. The chloride profiles perpendicular to the crack surface, particularly at the depth of 45 and 65 mm for GGBS concrete are shown in Fig. 3.11. It can be seen that within the range of approximately 5–10 mm from the crack surface, chloride concentrates remarkably. The chloride concentration drops quickly from the crack surface to the inside of concrete, and the higher concentration of chloride ions at the crack surface, the higher concentration can be observed at the same distance. However, because of the existence of aggregates along the crack surface and in the cracked zone, the obtained chloride concentration at crack zone may have some measurement errors.

54

3 Chloride Ingress in Cracked Concrete

Fig. 3.9 a EMPA Line scanning of chloride concentration for GGBS concrete subjected to immersion for 30 days b Area Line scanning of chloride concentration for GGBS concrete subjected to immersion for 60 days

On the other hand, the point along the crack surface serves as a new boundary vertically into a cracked zone except for the area influenced by aggregate particles. Some researchers have tried to simplify concrete crack merely as a plat-shape space and calculate the chloride diffusion coefficient regardless of concrete mixes. However, it can be seen that the mix proportions should not be ignored in calculating the chloride diffusion coefficient for cracked concrete.

3.4 Chloride Penetration into Cracked Concrete Under Wet-Dry Cycles 3.4.1 Experimental Program Tables 3.3 and 3.4 show the crack characteristics with a surface crack width of 0.1 and 0.2 mm. The cracked zone is defined as shown in Fig. 3.12. The thickness of the cracked zone was 5 mm on each side from the crack surface, based on the roughness of the crack surface and the average particle size of the aggregate, the same proposed by Kato et al. (2005). The concrete samples were taken from the cracked surface at an interval of 5 mm on the cracked zone and 10 mm for longer distance. The chloride ions content in each sample was determined based on EMPA.

Fig. 3.10 a Point EMPA scanning results of chloride concentration in the crack surface for OPC concrete subjected to immersion for 60d b Point EMPA scanning results of chloride concentration in the crack surface for GGBS concrete subjected to immersion for 60d

Chloride Content (% weight of concrete)

3.4 Chloride Penetration into Cracked Concrete Under Wet-Dry …

55

0.6

0.20mm crack width 0.15mm crack width 0.10mm crack width 0.05mm crack width

0.5 0.4 0.3 0.2 0.1 0.0

0

20

40

60

80

100

120

140

Depth from exposed surface (mm)

Chloride Content (% weight of concrete)

(a) 0.6

0.20mm crack width 0.15mm crack width 0.10mm crack width 0.05mm crack width

0.5 0.4 0.3 0.2 0.1 0.0 0

20

40

60

80

100

120

140

Depth from exposed surface(mm)

(b)

3.4.2 Chloride Profiles in Cracks The chloride ions profiles in the crack surface of the sliced specimens after 90 days are shown in Fig. 3.13. In sound concrete, there exists an ascent stage on chloride ions profiles because of the convection effect, resulting in maximum chloride concentration at the skin layer; this depth is always regarded as convection zone (Engelund et al. 2000). It can be seen that there exists an ascent point at the skin layer of cracked concrete as well. It can be explained by the reason similar to that of sound concrete. During the cyclic wet-dry action, the concentration at the skin layer exists a state when concentration is higher than that of surface, resulting in the diffusion of chloride ions at the skin layer, and as the number of wet-dry cycles increases, the depth of convection zone moves deeper. The depth of the convection zone is influenced by the crack characteristics in cracked concrete.

Fig. 3.11 a Chloride distribution perpendicular to the crack surface of GGBS concrete at surface crack width 0.1 mm b Chloride distribution perpendicular to the crack surface of GGBS concrete at surface crack width 0.15 mm

3 Chloride Ingress in Cracked Concrete

Chloride Content(% weight of concrete)

56

0.5 0.4 0.3

at depth of 65 mm from exposed surface

at depth of 45 mm from exposed surface

0.2 0.1 0.0

-20

-10

0

10

Distance from crack surface (mm)

20

Chloride Content(% weight of concrete)

(a)

0.5 0.4

at depth of 45 mm from exposed surface at depth of 65 mm from exposed surface

0.3 0.2 0.1 0.0

-20

-10

0

10

20

Distance from crack surface (mm)

(b)

Table 3.3 Characteristics of crack for specimen 1

Crack morphology specimen 1

Crack depth (mm)

Crack width (mm)

115

0

95

0.02

75

0.04

55

0.06

35

0.08

15

0.1

0 (reserved crack)

1.1

3.4 Chloride Penetration into Cracked Concrete Under Wet-Dry … Table 3.4 Characteristics of crack for specimen 2

Crack morphology specimen 2

57 Crack depth (mm)

Crack width (mm)

115

0

95

0.04

75

0.1

55

0.1

35

0.2

15

0.2

0 (reserved crack)

1.2

Fig. 3.12 Definition of the cracked zone

Figure 3.14 shows the chloride profiles both in the vertically and parallel directions in the cracked zone and deeper zone from 5 to 30 mm. The crack has a complex shape and the method of simplifying the crack characteristic also results in some errors. In addition, because of the concentration difference of every point among the cracked zone, every point along the crack surface serves as a new boundary for chloride ingress vertically into the cracked zone. From the point of selecting the thickness of the cracked zone, it can be seen that within the distance from the crack surface to about 5 mm from each side, the concentration of chloride ions changes more dramatically than that in a deeper zone. It also can be seen that at the same distance away from the crack surface, the higher chloride concentration at the crack surface, the higher concentration inside the crack zone. These phenomena can be explained by that the mechanism of chloride transport in the cracked zone is a diffusion-dominated process. It is assumed that if the concentration of the chloride ions solution in the crack is uniform, the chloride ions content in the cracked zone would not change along with the depth from the exposed surface. By comparing the chloride ions profiles in different crack widths at the surface, it also can be found that the chloride ions concentration tends to increase both in the

Fig. 3.13 a Chloride ions profiles in the cracked zone for the specimen with a surface crack width of 0.1 mm. b Chloride ions profiles in the cracked zone for the specimen with a surface crack width of 0.2 mm

3 Chloride Ingress in Cracked Concrete Chloride Content (%wt. of concrete)

58 0.6 0.5 0.4 0.3 0.2

0.1mm surface crack width

0.1 0

20

40

60

80

100

120

140

120

140

Depth from exposed surface (mm)

Chloride Content (%wt. of concrete)

(a) 0.6 0.5 0.4 0.3 0.2 0.2mm surface crack width

0.1 0

20

40

60

80

100

Depth from exposed surface (mm)

(b)

vertically and parallel directions in the cracked zone, as the crack widths increase. However, this tendency is limited to the points near the exposed surface and fades away at the crack pit, it should be known that the crack widths along the cracked zone must have some narrow points, not the assumed “V” shape.

3.5 Conclusions In this chapter, the influence of macroscopic cracks on the internal environments and chloride profiles in pre-cracked concrete with various crack patterns is discussed. The following conclusions can be drawn: (1) The concentration of chloride ions in the cracked zone tends to decrease with the increase of crack depths. The larger the crack width, the deeper the convection zone is.

Fig. 3.14 a Chloride ions profiles for the specimen with a surface crack width of 0.1 mm. b Chloride ions profiles for the specimen with a surface crack width of 0.2 mm

59

Chloride Content (%wt. of concrete)

3.5 Conclusions 0.55 0.50

0~15mm 15~35mm 35~55mm 55~75mm 75~95mm 95~115mm 115~135mm

0.45 0.40 0.35 0.30 0.25 0.20 0.15 0.10 0.05

-30

-20

-10

0

10

20

30

Distance from the crack surface (mm) Chloride Content (%wt. of concrete)

(a)

0.55 0.50

0~15mm 15~35mm 35~55mm 55~75mm 75~95mm 95~115mm 115~135mm

0.45 0.40 0.35 0.30 0.25 0.20 0.15 0.10 0.05 -30

-20

-10

0

10

20

30

Distance from the crack surface (mm)

(b) (2) At the same distance away from the crack surface, the higher chloride concentration at the crack surface, the higher concentration inside the crack zone. (3) For both OPC concrete and GGBS concrete, the larger in crack width results in higher chloride concentration at the crack surface. The chloride diffusion coefficient of cracked GGBS concrete is much more sensitive to cracks than that of OPC concrete. (4) The influential depth of moisture in concrete crack increases with the increase in crack width. For concrete specimens with 0.1, 0.2, and 0.3 mm crack widths at the surface, the moisture influential depth ranges approximately 25–35 mm, 50–60 mm, and 65–75 mm, respectively. At the drying stage after rainfall, the internal RH in the concrete crack with larger crack width tends to change more sensitively to the change of external environment.

60

3 Chloride Ingress in Cracked Concrete

References Andrade, C., Sarría, J., & Alonso, C. (1999). Relative humidity in the interior of concrete exposed to natural and artificial weathering. Cement and Concrete Research, 29(8), 1249–1259. Engelund, S., Edvardsen, C., & Mohr, L. (2000). General guidelines for durability design and redesign. In Report R15 of EU-Brite EuRam III project BE95-1347 DuraCrete. Probabilistic performance based durability design of concrete structures. Kato, E., Kato, Y., & Uomoto, T. (2005). Development of simulation model of chloride ion transportation in cracked concrete. Journal of Advanced Concrete Technology, 3(1), 85–94.

Chapter 4

Influence of Environmental Condition on Chloride Ingress into Loaded Concrete

Abstract In this chapter, the effect of cyclic wet-dry in the salt-fog condition which mimics the atmospheric zone environment on the chloride ingress into concrete is discussed. Also, the influence of carbonation on the chloride penetration in concrete under service load is discussed. The coupling effects of service load, chloride attack, and carbonation in RC beams are highlighted.

4.1 Introduction The environmental condition to which the concrete is exposed to plays an important role in affecting the chloride penetration process. The chloride penetration and steel corrosion are most accelerated for those exposed to tidal and splash zones since the cyclic wet-dry action accelerates chloride ingress and supplies oxygen and moisture for facilitating corrosion development. The atmospheric zone induces a cyclic wetdry effect due to seasonal variation in ambient RH and temperature, depending on the geographic locations. The chloride-contaminated concrete is also exposed to atmospheric carbonation or carbonate-containing water, both of which causes chloride redistribution.

4.2 Cyclic Wet-Dry Salt-Fog Environment 4.2.1 Experimental Program Mortar mixtures of different binder types and w/b ratios were studied, as shown in Table 4.1. Cubic specimens of 100 mm were exposed to simulated wet-dry saltfog environments. To realize one-dimension chloride ingress, only one surface of each specimen was exposed to simulated environments, while other surfaces were tightly waxed. As illustrated in Fig. 4.1a, the salt-fog wetting condition was realized by spraying a 5% NaCl solution in closed salt-fog chambers. The amount of salt-spray deposition, © The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2020 H. Ye et al., Chloride-Induced Steel Corrosion in Concrete Under Service Loads, https://doi.org/10.1007/978-981-15-4108-7_4

61

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4 Influence of Environmental Condition on Chloride …

Table 4.1 Mixing proportion of mortar (kg/m3 ) Mixture IDa

OPC

Sandb

OPC_42

489

0

Water

w/b

0

917

205

0.42

OPC_47

477

OPC_52

466

0

0

894

224

0.47

0

0

874

242

SL_47

381

0.52

95

0

893

224

0.47

FA_47 SF_47

380

0

94

890

223

0.47

309

94

71

890

223

0.47

GGBS

Fly ash

a SF represents the combined use of GGBS and FA, and the number following the abbreviation

Note represents the w/b ratio b The sand was in an oven-dried condition

Fig. 4.1 Salt-fog chambers during; a salt-fog spraying (wetting period); b drying period

which is defined as the volume of deposited NaCl solution on the unit exposed surface per hour (mL/(cm2 h)), was controlled at 1.5 mL/(cm2 h). The salt-fog drying condition was realized by opening the salt-fog chambers to the atmosphere, rendering the RH in salt-fog chambers to drop to around 60–80% (see Fig. 4.1b). For each saltfog wet-dry cycle, the wetting (i.e., salt-fog spraying) lasted for 24 h, followed by drying for 48 h (i.e., one salt-fog wet-dry cycles lasted for 3d). The temperature during the entire experiment remained at 35 ± 0.5 °C. In order to investigate the effects of exposure orientation to salt-fog environments on chloride profiles in the mortar, three different exposure angels (i.e., 0°, 60°, and 30°) to the horizontal were implemented, as illustrated in Fig. 4.2. The shaded surface is the exposure surface, while other surfaces were waxed. It should be noted that the exposure orientation with an angle of 90° to the horizontal was not included in this study due to its relatively low chloride content. In order to monitor the response of internal RH in mortar during salt-fog wetdry cycles, a series of RH sensors were embedded in the mortar. Different from the cubic specimens for chloride penetration tests, prism specimens with dimensions of 100 mm × 100 mm × 525 mm were used for internal RH monitoring. As illustrated

4.2 Cyclic Wet-Dry Salt-Fog Environment

A

63

B

C

Fig. 4.2 Illustration of specimen placement orientation in the salt-fog chamber; A an angle of 0° to the horizontal; B an angle of 60° to the horizontal; C an angle of 30° to the horizontal

in Fig. 4.3, in each specimen, five sensors were buried with the spacing between adjacent sensors of 80 mm, and depths of 5, 10, 20, 35, and 50 mm away from exposure surface. The surfaces of the specimens except the exposure surface were waxed to ensure one-dimension moisture transfer. After being cured at the moist room for 28 days, the specimens were exposed to the same salt-fog wet-dry cycles as those used for chloride penetration tests. The chloride concentration in mortar as a function of penetration depth was measured for each specimen after 15, 30, and 45 salt-fog wet-dry cycles. The mortar powders were drilled at the penetration depths of 3, 6, 9, 12, 15, 20, 25, 30, 35, 40, 45, and 50 mm. The mortar powders collected from three holes in each certain depth were mixed together to reduce the random errors (see Fu et al. 2016 for details). Rapid chloride test was implemented to measure the free (water-soluble) and total (acid-soluble) chloride content in the collected mortar powders. In particular, 1.5 g powder samples were dissolved in a 10 ml extraction liquid, shaken for 5 min, and then stood for 24 h. Thereafter, a calibrated electrode was submerged into the solution and the voltage was recorded, which was further transformed into the wateror acid-soluble chloride content (expressed as weight % Cl− by mortar) by means of the calibration curve. The adopted method was reasonably accurate to obtain the evolution of chloride profiles due to a number of variables, at least for comparative studies. Based on the free chloride concentration profiles, the depth of the convection zone can be obtained by measuring the depth where maximal concentration occurs.

4.2.2 Internal RH Evolution in Concrete Under Salt-Fog Environments Figure 4.4 shows the evolution of internal RH in the mortar during exposure to cyclic wet-dry salt-fog environments. It can be seen that the internal RH in mortar changes with the change of environmental RH in the fog-salt chamber. It indicates that the

64

4 Influence of Environmental Condition on Chloride …

(a)

(b)

(c) Fig. 4.3 Arrangements of RH sensors in the mortar (unit: in mm) a Top view; b Side view; c During specimen preparation

implemented salt-fog environment can successfully create cyclic wet-dry microclimates inside the specimens. Regardless of the mixture, the RH near the exposure surface shows a broader range of variance and lower mean value, in comparison to that in the deeper region. For instance, the RH in the depth of 5 mm ranges from about 85–98% RH, while the RH in the depth of 50 mm ranges from about 95–100% RH. It implies that the surface mortar tends to be more sensitive to the evolution of environments and the RH fluctuation diminishes as the depth increases. At such, the moisture status in mortar exposed to cyclic wet-dry fog-salt environments is in a non-equilibrium state. The evolution of internal RH in mortar is affected by the w/c ratio and incorporation of SCMs. With the increase of the w/c ratio, the evolution of RH becomes more sensitive to the variance of environments and the RH fluctuation range increases. For instance, in the depth of 50 mm, the RH in the mixture OPC_42 ranges from 98 to 100% RH, while it ranges from 94 to 100% RH for mixture OPC_52. This can be attributed to the coarser pore structure of mixture with a larger w/c ratio (see Fig. 4.5a for pore size distribution), which enlarges the moisture diffusion

4.2 Cyclic Wet-Dry Salt-Fog Environment

RH (%)

100 98 96 94 92 90 88 86 84 82 80 78 76 74 72 70

5 mm 10 mm 20 mm 35 mm 50 mm 0

3

OPC_42

6

9

12

15

18

21

Time (d)

(a)

RH (%)

100 98 96 94 92 90 88 86 84 82 80 78 76 74 72 70

5 mm 10 mm 20 mm 35 mm 50 mm 0

3

OPC_47

6

9

12

15

18

21

Time (d)

(b) 100 98 96 94 92 90 88 86 84 82 80 78 76 74 72 70

RH (%)

Fig. 4.4 Evolution of RH in the internal mortar exposed to the salt-fog wet-dry cycles a OPC_42; b OPC_47; c OPC_52; d SL_47; e FA_47; f SF_47

65

5 mm 10 mm 20 mm 35 mm 50 mm 0

3

OPC_52

6

9

12

Time (d )

(c)

15

18

21

66

4 Influence of Environmental Condition on Chloride …

RH (%)

Fig. 4.4 (continued)

100 98 96 94 92 90 88 86 84 82 80 78 76 74 72 70

5 mm 10 mm 20 mm 35 mm 50 mm 0

3

SL_47

6

9

12

15

18

15

18

15

18

21

Time (d )

RH (%)

(d)

100 98 96 94 92 90 88 86 84 82 80 78 76 74 72 70

5 mm 10 mm 20 mm 35 mm 50 mm 0

3

FA_47

6

9

12

21

Time (d)

RH (%)

(e)

100 98 96 94 92 90 88 86 84 82 80 78 76 74 72 70

5 mm 10 mm 20 mm 35 mm 50 mm

0

3

SF_47

6

9

12

Time (d)

(f)

21

4.2 Cyclic Wet-Dry Salt-Fog Environment

67

0.08

0.10 OPC_42 OPC_47 OPC_52

0.07

dV/dlogD (ml/g)

0.06

0.09 0.08 0.07

0.05

0.06

0.04

0.05

0.03

0.04 0.03

0.02

0.02

0.01 0.00

0.01 1

10

100

1000

10000

Cumulative intrusion (ml/g)

Fig. 4.5 Pore size distribution of mortar a OPC with various w/c ratios; b OPC with/without SCMs

0.00 100000

Diamater (nm)

(a)

0.08

0.08 0.07

dV/dlogD/(ml/g)

0.06 0.05

0.07 0.06 0.05

0.04

0.04

0.03

0.03

0.02

0.02

0.01

0.01

0.00

1

10

100

1000

10000

Cumulative intrusion (ml/g)

OPC_47 SL_47 FA_47 SF_47

0.00 100000

Diamater (nm)

(b) coefficient and hence increases the moisture influential depth (Ye et al. 2013). On the other hand, it is found that there exists a conspicuously delayed response of internal RH for FA and SF mixtures in the depth deeper than 5–10 mm. The delayed internal RH response may be an indication of the reduced moisture transfer coefficient due to the refined pore structure. As evidenced in Fig. 4.5b, the mixture with SCMs has a larger number of pores with diameters smaller than 7–10 nm, in comparison to pure OPC mixture. The RH measurements in the mortar interior represent an indirect methodology to measure the water content in material pores. However, considering different material porosities, the same RH value can represent different degrees of saturation in pores, which means different mass transport characteristics. At such, based on the evolution of internal RH, it is reasonable to deduce that the mass transport coefficient of mortar declines with decreasing w/c ratio and addition of SCMs.

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4 Influence of Environmental Condition on Chloride …

4.2.3 Influence of Exposure Duration on Chloride Profiles Figure 4.6 shows the free chloride concentration profiles in mortar exposed to wetdry salt-fog for 15, 30, and 45 cycles. It can be seen that the free chloride content considerably increases with increasing exposure duration, regardless of the mixture. Similar to being exposed to wet-dry seawater environments, the free chloride concentration profiles in mortar exposed to salt-fog environments also present a convection zone with a maximum chloride content shown near the exposure surface. The depth of the convection zone ranges from 5 to 12 mm, which is close to the value reported for concrete in wet-dry seawater condition (5–15 mm, Ye et al. 2012). Moreover, the depth of the convection zone increases with increasing exposure duration. It should be noticed that although the depth of convection zones between these two environments is close, it does not necessarily mean the cause of the convection zone is identical. The cause of convection zone can be a result of wet-dry actions, carbonation, or both (Ye et al. 2016).

4.2.4 Influence of Binder Type on Chloride Profiles The effects of w/c ratio and incorporation of SCMs on chloride resistance of mortar are clearly demonstrated in Figs. 4.6 and 4.7. With the increasing w/c ratio, the chloride content in the deep depth tends to increase, as well as the depth of the convection zone. It may be attributed to the coarser pore structure in the mortar with a larger w/c ratio, as shown in Fig. 4.5a, which enlarges the porosity and pore size that allows more extensive moisture transfer and chloride ions ingress. This argument is somehow supported by the more extensive RH variation in mixture with a larger w/c ratio. On the other hand, the incorporation of SCMs can noticeably reduce the free chloride content in the deep depth but tends to increase it within the skin layer. In other words, there is a free chloride accumulation effect in the mortar with SCMs. After 15 and 30 salt-fog cycles, it can be seen that the SF mixture has the highest free chloride content in the skin layer and lowest free chloride content in the deep depth. However, after 45 cycles, the free chloride content in the deep depth in SF mixture excesses those of SL and FA mixtures. It suggests that the accumulated free chloride ions at the skin layer of mortar tend to increase the chloride concentration gradient that facilitates the movement of free chloride ions inwards. It implies that in short term, the addition of SCMs is beneficial for chloride resistance at deep depth for mortar exposed to the salt-fog environment due to the free chloride accumulation effect; however, it may impose a potential threat in the long term. The significant amount of accumulated free chloride ions at SF surface may be attributed to the densest pore structure but lowest carbonation resistance due to portlandite deficiency (Lee et al. 2013; Ye et al. 2016). In addition, it is interesting to notice that the FA mixture seems to have the strongest chloride resistance among all mixtures after 45 salt-fog cycles, which may be attributed to the reduced moisture transfer diffusion

1.0 OPC_42_15 cycles OPC_42_30 cycles OPC_42_45 cycles

0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0

0

5

10

15

20

25

30

35

40

45

50

Depth (mm)

Free chloride content (weight % by mortar)

(a) 1.0 OPC_47_15 cycles OPC_47_30 cycles OPC_47_45 cycles

0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0

0

5

10

15

20

25

30

35

40

45

50

Depth (mm)

Free chloride content (weight % by mortar)

Fig. 4.6 Free chloride concentration profiles in mortar exposed to cyclic wet-dry salt-fog environments; a OPC_42; b OPC_47; c OPC_52; d SL_47; e FA_47; f SF_47

69 Free chloride content (weight % by mortar)

4.2 Cyclic Wet-Dry Salt-Fog Environment

(b) 1.0 OPC_52_15 cycles OPC_52_30 cycles OPC_52_45 cycles

0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0

0

5

10

15

20

25

30

Depth (mm)

(c)

35

40

45

50

Free chloride content (weight % by mortar)

1.0 SL_47_15 cycles SL_47_30 cycles SL_47_45 cycles

0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0

0

5

10

15

20

25

30

35

40

45

50

Depth (mm)

(d) Free chloride content (weight % by mortar)

Fig. 4.6 (continued)

4 Influence of Environmental Condition on Chloride …

1.0 FA_47_15 cycles FA_47_30 cycles FA_47_45 cycles

0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0

0

5

10

15

20

25

30

35

40

45

50

Depth (mm)

(e) Free chloride content (weight % by mortar)

70

1.0 SF_47_15 cycles SF_47_30 cycles SF_47_45 cycles

0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0

0

5

10

15

20

25

30

Depth (mm)

(f)

35

40

45

50

71 1.0 OPC_42_30 cycles OPC_47_30 cycles OPC_52_30 cycles

0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0

0

5

10

15

20

25

30

35

40

45

50

Depth (mm) Free chloride content (weight % by mortar)

(a)

1.0 OPC_47_15 cycles SL_47_15 cycles FA_47_15 cycles SF_47_15 cycles

0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0

0

5

10

15

20

25

30

35

40

45

50

Depth (mm)

(b) Free chloride content (weight % by mortar)

Fig. 4.7 Free chloride concentration profiles in mortar exposed to cyclic wet-dry salt-fog environments; a OPC with various w/c ratios after 30 salt-fog cycles; b OPC with/without SCMs after 15 cycles; c OPC with/without SCMs after 30 cycles; d OPC with/without SCMs after 45 cycles

Free chloride content (weight % by mortar)

4.2 Cyclic Wet-Dry Salt-Fog Environment

1.0 OPC_47_30 cycles SL_47_30 cycles FA_47_30 cycles SF_47_30 cycles

0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0

0

5

10

15

20

25

30

Depth (mm)

(c)

35

40

45

50

Fig. 4.7 (continued)

4 Influence of Environmental Condition on Chloride … Free chloride content (weight % by mortar)

72

1.0 OPC_47_45 cycles SL_47_45 cycles FA_47_45 cycles SF_47_45 cycles

0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0

0

5

10

15

20

25

30

35

40

45

50

Depth (mm)

(d)

coefficient (as reflected in delayed RH response in Fig. 4.4), and strongest chloride binding capacity.

4.2.5 Influence of Exposure Orientation on Chloride Profiles Figure 4.8 shows the free chloride concentration profiles in OPC_47 mixture with three different exposure orientations after 30 and 45 salt-fog wet-dry cycles. It can be seen that the chloride concentration in specimens with the Type A exposure orientation is considerably higher than those of Type B and C. It suggests that the chloride profiles in mortar under salt-fog environments are significantly affected by the exposure orientations. In particular, the chloride concentration tends to be higher when the horizontally projected area with respect to the salt-fog deposition direction is larger. However, it is interesting to notice that the depth of the convection zone is not affected by the exposure orientation. This may imply that the amount of chloride concentration at the exposure surface hardly affects the occurrence of the convection zone.

4.3 Cyclic Carbonation and Wet-Dry Condition 4.3.1 Experimental Program In order to investigate the influence of SCMs on the chloride penetration resistance, two types of concrete mixture proportions with different SCMs content were used, as shown in Table 4.2. Mixture PC is pure OPC-based concrete and Mixture FS has considerable amounts of fly ash and GGBS, while both of them have the same w/b ratio and volume fraction of aggregates.

4.3 Cyclic Carbonation and Wet-Dry Condition Free chloride content (weight % by mortar)

Fig. 4.8 Free chloride concentration profiles in OPC_47 mixture with three different exposure orientations after; a 30 salt-fog cycles; b 45 salt-fog cycles

73

1.0 30 cycles

0.9

OPC_47_Type A orientation OPC_47_Type B orientation OPC_47_Type C orientation

0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0

0

5

10

15

20

25

30

35

40

45

50

Depth (mm) Free chloride content (weight % by mortar)

(a) 1.0 45 cycles

0.9

OPC_47_Type A orientation OPC_47_Type B orientation OPC_47_Type C orientation

0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1 0.0

0

5

10

15

20

25

30

35

40

45

50

Depth (mm)

(b) Table 4.2 Mixing proportion of concrete (kg/m3 ) Mixture ID

Cement

PC

433

FS

281.4

Fly ash

Slag

Fine aggregate

Coarse aggregate

Superplasticizer

Water

0

0

709

1063

0.866

195

65

86.6

709

1063

0.866

195

The configuration of the sample was a prism with dimensions of 100 mm × 100 mm × 460 mm. All samples were cast and demodeled after 24 h, followed by the moist curing at 100% RH and 20 ± 5 °C for 28 days. After moisture curing, every two samples (on PC and one FS) were grouped and were loaded using a back-to-back four-point loading scheme as shown in Fig. 4.9a.

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4 Influence of Environmental Condition on Chloride …

Fig. 4.9 a Configuration of a sample with loading system. b Setup for chloride penetration test

For each mixture, 39 samples were cast in total, 3 of which was used for testing the ultimate loading capability, 12 samples were remained un-loaded, 12 samples were loaded with 30% of the ultimate loading capability, and 12 samples were loaded with 60% of the ultimate loading capability. The ultimate loading capability for PC and FS mixtures was measured to be 13.2 kN and 12.7 kN, respectively.

4.3 Cyclic Carbonation and Wet-Dry Condition

75

Table 4.3 Exposure conditions for one wet-dry-carbonation cycle Exposure ID

Experimental procedure for one testing cycle

I (Control)

15% NaCl solution for four days

Oven-drying for two days

Natural air environment for two days

II

15% NaCl solution for four days

Oven-drying for two days

Accelerated carbonation for two days

III

15% NaCl solution for four days

Oven-drying for two days

Accelerated carbonation for four days

After curing and loading, all samples were experiencing three different types of exposure as shown in Table 4.3. In particular, the samples were first immersed into a highly concentrated 15% NaCl solution for four days (as shown in Fig. 4.9b). Afterward, they were surface-dried before putting into an oven which was programmed to 60 °C for two days. Then, they were selectively moved to an accelerated carbonation chamber which was programmed to a CO2 concentration of 20 ± 3%, temperature of 20 ± 5 °C, and 70 ± 5% RH. More specifically, one-third of samples (with same mixture proportion and loading magnitude) were exposed to natural air as control (i.e., exposure condition I), while one-third were acceleratedly carbonated for two days (i.e., exposure condition II), and another one-third were acceleratedly carbonated for four days (i.e., exposure condition III). In addition, in order to realize one-dimension chloride penetration, merely the two opposite surfaces (i.e., the two faces experiencing pure flexural loadings as shown in Fig. 4.9) were exposed to chloride attack, while other surfaces were coated with high melting point paraffin. The chloride concentration in concrete as a function of depth was measured for each sample after 2, 4, 6, 8, and 10 wet-dry-carbonation cycles (i.e., one cycle is as indicated in Table 4.3). The concrete powders were drilled with an interval of 2 mm for the first 10 mm away from the exposed surface, then with an interval of 5 mm in the range of 10–30 mm. In each position, the concrete powders collected from three holes were mixed to avoid the influence of aggregates. The powder with a weight of 15 g (passing through 0.63 mm sieve) was oven-dried at 105 ± 5 °C for two hours before cooling down and conducting the chloride content measurements. A rapid chloride ion concentration determination method of automatic potentiometric titration was used to precisely measure the free (water-soluble) chloride content in concrete powders. The carbonation depth in concrete as a function of depth was measured for each sample after 2, 4, 6, 8, and 10 wet-dry-carbonation cycles. The depths of carbonation were determined by spraying on a freshly broken surface with phenolphthalein. Given that the carbonated zone is uncolored (concrete color) and the non-carbonated portion is purple, the depth of carbonation for a certain point was measured by callipering the width of the uncolored region vertically to the exposed surface. Therefore, the carbonation depth for each sample can be determined by averaging the depth of carbonation at 10 different points along the edge of the broken surface.

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4 Influence of Environmental Condition on Chloride …

4.3.2 Influence of Carbonation on Chloride Profiles Figure 4.10 shows the water-soluble (free) chloride profiles in stress-free concrete exposed to different wet-dry-carbonation exposure conditions. It can be seen that the free chloride content at each location increases with the increasing number of cycles, regardless of mixture type and exposed condition. According to the previous studies, cyclic wet-dry action can result in a convective zone, in which the concentration of chloride reaches a local maximum at the near surface (Nilsson 2000; Ye et al. 2012, 2013, 2015). The primary reason for the presence of convective zone is that the moisture influencing depth is limited at the surface of exposure, typically ranging between 5 and 15 mm for sound concrete (Gang et al. 2015; Ye et al. 2012, 2013, 2015) In the range of convective zone, the drying period can result in a high chloride concentrated pore solution, which can be acceleratedly transported inwards due to capillary suction during the subsequent wetting period. Beyond the convective zone, the moisture status is less sensitive to the evolution of external environments (Ye et al. 2013), and hence the diffusion process dominates the chloride penetration. For the free chloride profiles in PC and FS mixtures without accelerated carbonation (i.e., control condition), the presence of convective zone (in the depth of 4–6 mm) is merely conspicuous for FS mixture. Additionally, the free chloride content for the FS mixture is considerably higher than that of PC mixture near the exposed surface, then becomes less for interior regions. For example, the free chloride contents for FS-10I-00 are about 1.26% and 0.16% at a depth of approximately 3.0 mm and 17.5 mm, respectively, while the free chloride contents for PC-10I-00 are 0.89% and 0.35% at these two respective locations. These findings indicate that the incorporation of SCMs makes the free chloride ion accumulate at the exposed surface, which actually reduces the interior chloride content. This phenomenon is likely due to a smaller effective chloride diffusion coefficient in FS concrete due to the modified pore structure and chloride binding capability. As a consequence, the chloride penetration process in FS concrete is much slower compared with that of PC mixture, resulting in a gradual accumulation at the surface during wet-dry cycles. Carbonation shows a significant influence on the free chloride profiles in both PC and FS mixtures. In the case of PC mixture, it can be seen that the free chloride content at the surface (approximately