Figure 7.3 Multimeter (FLUKE Model No. 179 True RMS Multimeter). ...... Bridges Relative to Reinforcement Corrosion, Volume 8: Procedure Manual, 1992.
ACCELERATED QUANTIFICATION OF CRITICAL PARAMETERS FOR PREDICTING THE SERVICE LIFE AND LIFE CYCLE COSTS OF CHLORIDE-LADEN REINFORCED CONCRETE STRUCTURES
A Thesis by RADHAKRISHNA PILLAI GOPALAKRISHNAN
Submitted to the Office of Graduate Studies of Texas A&M University in partial fulfillment of the requirements for the degree of MASTER OF SCIENCE
August 2003
Major Subject: Civil Engineering
ACCELERATED QUANTIFICATION OF CRITICAL PARAMETERS FOR PREDICTING THE SERVICE LIFE AND LIFE CYCLE COSTS OF CHLORIDE-LADEN REINFORCED CONCRETE STRUCTURES
A Thesis by RADHAKRISHNA PILLAI GOPALAKRISHNAN
Submitted to the Office of Graduate Studies of Texas A&M University in partial fulfillment of the requirements for the degree of MASTER OF SCIENCE
Approved as to style and content by:
David Trejo (Chair of Committee)
Joseph M. Bracci (Member)
Richard B. Griffin (Member)
Paul N. Roschke (Head of Department) August 2003
Major Subject: Civil Engineering
iii
ABSTRACT
Accelerated Quantification of Critical Parameters for Predicting the Service Life and Life Cycle Costs of Chloride-Laden Reinforced Concrete Structures. (August 2003) Radhakrishna Pillai Gopalakrishnan, B.E., University of Allahabad, Allahabad, India Chair of Advisory Committee: Dr. David Trejo
The use of corrosion resistant steels (instead of conventional carbon steels) and/or high performance concrete can increase the overall service life and can reduce the life cycle cost (LCC) of reinforced concrete (RC) structures exposed to chloride environments. At present, no accelerated standardized test procedures are available to directly evaluate critical parameters affecting the service life of RC systems and current test methods can take years or decades to indirectly evaluate these critical parameters for high performance construction materials. This prevents the engineers, designers, and owners from using new high performance materials, especially, the corrosion resistant steel reinforcement. This thesis evaluates the Accelerated Chloride Threshold (ACT) test procedure developed to determine the critical chloride threshold value of uncoated steel reinforcement embedded in cementitious materials. Using the ACT test procedure, the critical chloride threshold values of the ASTM A706, ASTM A615, microcomposite, SS304, and SS316LN reinforcement types were determined to be 0.2 kg/m3 (0.3 lb/yd3),
iv
0.5 kg/m3 (0.9 lb/yd3), 4.6 kg/m3 (7.7 lb/yd3), 5.0 kg/m3 (8.5 lb/yd3), and 10.8 kg/m3 (18.1 lb/yd3), respectively.
Using these values, the time to corrosion initiation of
chloride-laden RC systems can be determined. The Accelerated Cracking and Spalling Threshold (ACST) test procedure has been developed to determine the amount of steel corrosion required to cause cracking and spalling of concrete cover. From preliminary experimental data, the critical cracking and spalling threshold thickness for a 19 mm (0.75 inch) concrete cover with 0.45, 0.55, and 0.65 water-cement ratios has been determined to be 20.64, 16.85, and 37.46 mils, respectively. Preliminary results indicate that for a cover depth of 19 mm (0.75 inch) the critical
cracking
and
10 [ −2.4 + (12.5 × w / c ) − 11.6 × ( w / c )
spalling
2 −1
]
threshold
value
(mils)
is
equal
to
and can be used to determine the time of corrosion
propagation in chloride-laden RC systems. A parametric study with different steel reinforcement, water-cement ratios, and chloride exposure conditions indicated that the use of corrosion resistant steels will increase the overall service life and can reduce the LCC of RC structures exposed to severe chloride environments.
v
ACKNOWLEDGEMENTS
The author takes this opportunity to acknowledge the excellent academic guidance and financial assistance offered by Dr. David Trejo during the entire research project. Without his helping hands this thesis would not have been in its current shape. Dr. Trejo, thank you for awarding me the title, "Nocturnal Researcher". The timely support offered by Dr. Joseph M. Bracci and Dr. Richard B. Griffin is also appreciated. The author offers a special acknowledgement of appreciation to the Texas Engineering Experiment Station and MMFX Technologies, Inc., for funding the research project. The author acknowledges all of the technical assistance he received from Mr. Francisco Aguiniga, Ms. Victoria Salgado, Mr. Alan Bell, Mr. Michael Esfeller, Mr. Ceki Halmen, Mr. Fayez Mautassem, and other members of the Trejo Research Group. The author acknowledges Mr. Partha, Mr. Vivek, and other inmates of the basement of Civil Engineering building, especially those who love to work during the nighttime, for all the help I received. Also acknowledged is the help offered by Mr. Scott Cronauer, Ms. Laura Raleigh and other employees in the Department of Civil Engineering. I greatly appreciate all of the favors and advice I received from my friends including (in the chronological order I initially met them) Vipul, Aninda, Parasuram, Abhijit, Shyam, Rajesh, Bharani, Vinodh, Balaji, Bharath, Amitabh, Karun, Julian, Elango, Pratheesh, etc. (the list continues...) during the course of my graduate studies at TAMU. I am grateful to Sunil chettan, Leena chechi and their daughter, Pooja mol, for all kinds of favors, especially the tasty food, they offered me.
vi
I offer special thanks to Achan, Amma, Sandhya, and Tushi for their patience and silent support without which this endeavor would not have been accomplished. Support from other family members is also appreciated.
vii
TABLE OF CONTENTS Page 1
THE INTRODUCTION ................................................................................................ 1 1.1 BACKGROUND.................................................................................................. 1 1.1.1 Chloride-induced corrosion in concrete: Causes......................................... 1 1.1.2 Chloride-induced corrosion in concrete: Remedies .................................... 2 1.1.2.1 High performance cementitious materials................................................ 2 1.1.2.2 High performance steel reinforcement ..................................................... 3 1.1.3 Critical parameters for service life prediction and life cycle cost analysis . 4 1.2 PROBLEM STATEMENT AND RESEARCH OBJECTIVES........................... 5 1.3 THESIS ORGANIZATION ................................................................................. 6
2 BASICS OF ELECTROCHEMICAL CORROSION................................................... 9 2.1 INTRODUCTION................................................................................................ 9 2.2 FORMS OF CORROSION ................................................................................ 10 2.2.1 General corrosion ...................................................................................... 10 2.2.2 Localized corrosion ................................................................................... 11 2.3 MECHANISMS OF CORROSION ................................................................... 11 2.4 THERMODYNAMICS OF CORROSION........................................................ 14 2.4.1 Electrochemical potential of corrosion reactions ...................................... 14 2.4.1.1 Activity and Gibbs free energy .............................................................. 15 2.4.1.2 The fundamental work-energy relationships.......................................... 16 2.5 KINETICS OF CORROSION............................................................................ 20 2.5.1 Corrosion rate ............................................................................................ 20 2.5.1.1 Average corrosion rate ........................................................................... 21 2.5.1.2 Instantaneous corrosion rate................................................................... 22 2.6 PROTECTIVE SURFACE BARRIERS ............................................................ 23 3 MECHANISMS OF CHLORIDE-INDUCED CORROSION IN CONCRETE ........ 25 3.1 CHLORIDE PENETRATION IN UNCRACKED CONCRETE....................... 26 3.1.1 Diffusion of chloride ions in concrete....................................................... 27 3.1.1.1 Effect of water-binder ratio .................................................................... 28 3.1.1.2 Effect of cement type and supplementary cementitious materials ......... 29 3.1.1.3 Effect of aggregates................................................................................ 31 3.1.1.4 Effect of compaction and consolidation................................................. 33 3.1.1.5 Effect of initial curing conditions........................................................... 34 3.1.1.6 Effect of environmental conditions ........................................................ 35 3.1.1.7 Effect of chloride exposure conditions and time of exposure ................ 37
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Page 3.1.2 Mathematical models for diffusion based chloride transport in concrete . 39 3.2 CHLORIDE PENETRATION IN CRACKED CONCRETE ............................ 43 3.3 INITIATION OF CHLORIDE-INDUCED CORROSION................................ 45 3.3.1.1 The passive film in concrete................................................................... 46 3.3.2 Chloride-induced breakdown of passive film ........................................... 48 3.3.3 The corrosion reactions after the breakdown of the protective layers ...... 50 3.3.4 Critical chloride threshold value ............................................................... 52 3.3.5 Factors influencing the critical chloride threshold value .......................... 58 3.3.5.1 Steel characteristics ................................................................................ 58 3.3.5.2 Cementitious material and interfacial transition zone characteristics .... 60 3.4 PROPAGATION OF CHLORIDE-INDUCED CORROSION ......................... 61 3.5 CORROSION-INDUCED CRACKING OR SPALLING OF CONCRETE COVER .............................................................................................................. 64 3.5.1 Critical amount of corrosion products resulting in cracking and spalling 67 3.5.2 Cracking and spalling threshold thickness ................................................ 70 4 SERVICE LIFE AND LIFE CYCLE COST OF RC STRUCTURES EXPOSED TO CHLORIDE ENVIRONMENTS.......................................................................... 73 4.1 SERVICE LIFE OF RC STRUCTURES............................................................ 73 4.1.1 Definitions and influencing factors ........................................................... 73 4.1.2 Various time phases and prediction of service life ................................... 75 4.1.3 The chloride-induced corrosion initiation phase ....................................... 76 4.1.4 The chloride-induced corrosion propagation phase .................................. 79 4.1.5 The repair and rehabilitation phase ........................................................... 82 4.1.6 Methodology for predicting service life of RC structures exposed to chloride environments ............................................................................... 83 4.2 LIFE CYCLE COST OF RC STRUCTURES .................................................... 85 4.2.1 Definition and factors contributing to the life-cycle cost.......................... 85 4.2.2 Life cycle cost analysis ............................................................................. 86 5 CURRENT TEST METHODS TO PREDICT SERVICE-LIFE OF RC STRUCTURES EXPOSED TO CHLORIDE ENVIRONMENTS ............................ 94 5.1 ACCELERATED METHODS FOR CHLORIDE PENETRATION................. 94 5.1.1 Chloride penetration by cyclic wet-dry exposure ..................................... 94 5.1.2 Electrically accelerated chloride penetration ............................................ 95 5.2 CORROSION RATE MEASUREMENT BY MASS LOSS TESTS ................. 97 5.3 ELECTROCHEMICAL METHODS FOR CORROSION MONITORING...... 98 5.3.1 Half-cell potential measurements.............................................................. 99 5.3.2 Polarization resistance measurement techniques .................................... 101
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Page 5.4
CHEMICAL METHODS FOR CHLORIDE CONTENT ANALYSIS ........... 107
6 RESEARCH SIGNIFICANCE ................................................................................. 110 7 EXPERIMENTAL PROGRAM AND PRELIMINARY TESTS............................. 112 7.1 RESEARCH OBJECTIVES............................................................................. 112 7.2 THE ACCELERATED CHLORIDE THRESHOLD (ACT) TEST ................. 112 7.2.1 The ACT test methodology ..................................................................... 113 7.2.1.1 The general test methodology .............................................................. 113 7.2.1.2 The ACT test layout ............................................................................. 114 7.2.1.3 The accelerated chloride transport system ........................................... 116 7.2.1.4 The corrosion initiation detection system ............................................ 116 7.2.1.5 Quantification of the critical chloride concentration............................ 118 7.2.2 Evaluation and engineering refinement of the ACT test......................... 118 7.2.2.1 Type of potential gradient (voltage) source and electrical timer.......... 118 7.2.2.2 Steel potential variations due to applied potential gradient ................. 121 7.2.2.3 Chloride migration rate and pH variations due to applied potential gradient................................................................................................. 126 7.2.2.4 Time to formation of a stable passive film........................................... 130 7.2.2.5 Time for attaining a stabilized polarization resistance......................... 131 7.2.2.6 Reference electrode, Haber-Lugin probe and Haber-Lugin probe electrolyte ............................................................................................. 133 7.2.2.7 Voltage source - distribution box assembly ......................................... 135 7.2.2.8 Definition of parameters for electrochemical testing ........................... 136 7.2.2.9 Mortar dust collection and modified chloride analysis method ........... 137 7.2.3 Materials and experimental design: ACT tests ....................................... 141 7.3 THE ACCELERATED CRACKING AND SPALLING THRESHOLD (CST) TEST ................................................................................................................ 146 7.3.1 The general test methodology ................................................................. 146 7.3.2 The ACST test layout, and procedure ..................................................... 146 7.3.3 Materials and experimental design: ACST tests ..................................... 152 8 RESULTS AND DISCUSSIONS ............................................................................. 156 8.1 CRITICAL CHLORIDE THRESHOLD VALUES ......................................... 156 8.1.1 ASTM A706 type reinforcement............................................................. 158 8.1.2 ASTM A615 type reinforcement............................................................. 162 8.1.3 Microcomposite steel reinforcement ....................................................... 168 8.1.4 Stainless steel 304 reinforcement ............................................................ 174 8.1.5 Stainless steel 316LN reinforcement....................................................... 178 8.1.6 Summary of critical chloride threshold values........................................ 182
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Page 8.2 THE DURATION OF THE CORROSION INITIATION PHASE .................. 184 8.3 CRITICAL CRACKING AND SPALLING THRESHOLD THICKNESS..... 205 8.4 THE DURATION OF CORROSION PROPAGATION PHASE .................... 209 8.5 OVERALL SERVICE LIFE AND LIFE CYCLE COST COMPARISON...... 211 8.5.1 Overall service life .................................................................................. 211 8.5.2 Life cycle cost comparison...................................................................... 212 9 CONCLUSIONS AND FUTURE RECOMMENDATIONS ................................... 219 9.1 9.2
RESEARCH CONCLUSIONS ........................................................................ 219 RECOMMENDATIONS FOR FUTURE RESEARCH................................... 221
REFERENCES................................................................................................................223 APPENDIX A..................................................................................................................241 APPENDIX B ..................................................................................................................257 APPENDIX C ..................................................................................................................266 APPENDIX D ..................................................................................................................279 APPENDIX E ..................................................................................................................281 APPENDIX F ..................................................................................................................283 APPENDIX G ..................................................................................................................296 APPENDIX H ..................................................................................................................298 APPENDIX I ..................................................................................................................300 VITA................................................................................................................................302
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LIST OF FIGURES Page Figure 2.1 General and localized corrosion..................................................................... 10 Figure 2.2 The variation of Ecell as a function of pH....................................................... 19 Figure 2.3 The variation of Ecell versus common logarithm of O2 concentration. .......... 19 Figure 3.1 A schematic of the general diffusion process. ............................................... 27 Figure 3.2 The interfacial transition zones (ITZ) around aggregates in concrete. .......... 32 Figure 3.3 The adsorption-displacement mechanism of passive film breakdown (Stansbury and Buchanan 2000). .................................................................. 48 Figure 3.4 Typical mechanism of propagation of chloride-induced corrosion in concrete (Adapted from Ahmed 2003). ........................................................ 62 Figure 3.5 The free expansion phase............................................................................... 66 Figure 3.6 The stress initiation phase. ............................................................................. 66 Figure 3.7 The cracking or spalling phase....................................................................... 67 Figure 3.8 Expansive pressure on surrounding concrete due to formation of rusty products (Liu 1996)....................................................................................... 68 Figure 3.9 The cracking, spalling and delamination of concrete cover........................... 70 Figure 4.1 Some important factors influencing the service life or RC structures exposed to chloride environments. ............................................................... 74 Figure 4.2 Different phases of overall service life in terms of reinforcement corrosion in RC structures (adapted after Trejo 2002). ................................ 75 Figure 4.3 A flow chart for predicting the overall service life of RC structures exposed to chloride environments. ............................................................... 84 Figure 4.4 Some important factors contributing to the life cycle cost of RC structures. 86 Figure 4.5 A typical cash flow diagram. ......................................................................... 89
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Page Figure 4.6 A flow chart for determining the life cycle cost (LCC) of RC structures exposed to chloride environments. ............................................................... 91 Figure 4.7 A simple flowchart for selecting the most durable and cost effective material combination or alternative for RC structures.................................. 92 Figure 5.1 A typical polarization curve......................................................................... 102 Figure 5.2 Randle's circuit for steel reinforcement in concrete (adapted from Millard et al. 2001). .................................................................................... 104 Figure 5.3 A simple schematic of the guard ring arrangement for corrosion rate measurement on steel reinforcement (adapted from Song 2000). .............. 106 Figure 7.1 ACT test layout. ........................................................................................... 115 Figure 7.2 Potential gradient source (DC Power Supply system, Agilent Technologies Model No. E3611A). ............................................................ 119 Figure 7.3 Multimeter (FLUKE Model No. 179 True RMS Multimeter)..................... 120 Figure 7.4 Electrical timer (General Electric Company (Model No. PM613US))........ 121 Figure 7.5 Equipotential and equicurrent lines in the upper portion of the ACT specimen (adapted from Trejo and Pillai 2003a)........................................ 122 Figure 7.6 OCP variation before, during, and after the application of potential gradient. ...................................................................................................... 123 Figure 7.7 The potential as a function of the distance from the anode, during the application of the external potential gradient of 20 Volts (a). .................... 124 Figure 7.8 The potential as a function of the distance from the anode, during the application of the external potential gradient of 20 Volts (b)..................... 125 Figure 7.9 Mortar regions used for chloride migration and pH study with different levels of applied potential gradient. ............................................................ 127 Figure 7.10 Chloride concentrations near the steel specimen surface for different levels of applied potential gradients (Trejo and Pillai 2003a). ................... 128 Figure 7.11 pH values at the anode-mortar interface for different levels of applied potential gradients (Trejo and Pillai 2003a)................................................ 129
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Page Figure 7.12 Variation of OCP with time of exposure showing the formation of stable passive film....................................................................................... 130 Figure 7.13 Variation of inverse R p after the voltage source is switched off. ............. 132 Figure 7.14 Accumet® Standard-Size Prefilled Saturated Calomel Electrode. ............. 133 Figure 7.15 Reference electrode, Haber-Lugin probe and Haber-Lugin probe electrolyte system........................................................................................ 134 Figure 7.16 Front view of voltage source - distribution box assembly. ........................ 135 Figure 7.17 Side view of voltage source-distribution box assembly. ........................... 136 Figure 7.18 Profile Grinder supplied by Germann Instruments, Inc. (Model No. Metabo D-72622)........................................................................................ 138 Figure 7.19 Mortar dust collection system. ................................................................... 139 Figure 7.20 Accelerated CST test layout....................................................................... 147 Figure 7.21 Steel reinforcement for the accelerated CST test....................................... 148 Figure 7.22 Regions for crack width monitoring. ........................................................ 149 Figure 7.23 Crack detection microscope (ELE International, Inc., Model No. 35-2505)...................................................................................................... 150 Figure 7.24 A typical corroded top reinforcement from the ACST test indicating that only the top portion is corroding.......................................................... 151 Figure 7.25 A schematic showing the corroding area of the reinforcement. ................ 151 Figure 7.26 The compressive strength as a function of curing period for concrete with different water-cement ratios (w/c)..................................................... 154 Figure 8.1 Inverse polarization resistance values of ASTM A706 "as received" ACT samples as a function of time of applied potential gradient (after Trejo and Pillai 2003a)...................................................................... 161 Figure 8.2 Inverse polarization resistance values of ASTM A706 "polished" ACT samples as a function of time of applied potential gradient........................ 161
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Page Figure 8.3 Inverse polarization resistance values of ASTM A615 "as received" ACT samples as a function of time of applied potential gradient (Sample #1 through #9) (after Trejo and Pillai 2003a)............................... 166 Figure 8.4 Inverse polarization resistance values of ASTM A615 "as received" ACT samples as a function of time of applied potential gradient (Sample #10 through #18) (after Trejo and Pillai 2003a).......................... 166 Figure 8.5 Inverse polarization resistance values of ASTM A615 "polished" ACT samples as a function of time of applied potential gradient (Sample #1 through #8). ............................................................................. 167 Figure 8.6 Inverse polarization resistance values of ASTM A615 "polished" ACT samples as a function of time of applied potential gradient (Sample #9 through #17). ........................................................................... 167 Figure 8.7 Inverse polarization resistance values of microcomposite "as received" ACT samples as a function of time of applied potential gradient (after Trejo and Pillai 2003b)...................................................................... 171 Figure 8.8 Inverse polarization resistance values of microcomposite "polished" ACT samples as a function of time of applied potential gradient (Sample #1 through #9). ............................................................................. 172 Figure 8.9 Inverse polarization resistance values of microcomposite "polished" ACT samples as a function of time of applied potential gradient (Sample #10 through #17). ......................................................................... 172 Figure 8.10 Inverse polarization resistance values of stainless steel 304 "as received" ACT samples as a function of time of applied potential gradient (after Trejo and Pillai 2003b)...................................................................... 177 Figure 8.11 Inverse polarization resistance values of stainless steel 304 "polished" ACT samples as a function of time of applied potential gradient............... 177 Figure 8.12 Inverse polarization resistance values of stainless steel 316LN "as received" ACT samples as a function of time of applied potential gradient (after Trejo and Pillai 2003b)........................................................ 181 Figure 8.13 Inverse polarization resistance values of stainless steel 316LN "polished" ACT samples as a function of time of applied potential gradient. ....................................................................................... 181
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Page Figure 8.14 A box plot showing critical chloride threshold values of different steel reinforcement types tested using ACT test procedure (after Trejo and Pillai 2003a and Trejo and Pillai 2003b). ................................................... 183 Figure 8.15 Surface chloride concentration as a function of time of exposure, with k calculated using surface chloride concentration at the end of first year.................................................................................................. 189 Figure 8.16 Surface chloride concentration as a function of time of exposure, with k calculated using surface chloride concentration at the end of second year. ............................................................................................ 190 Figure 8.17 Surface chloride concentration as a function of time of exposure, with k calculated using surface chloride concentration at the end of fifth year. ................................................................................................ 191 Figure 8.18 Surface chloride concentration as a function of time of exposure, with k calculated using surface chloride concentration at the end of tenth year. ............................................................................................... 192 Figure 8.19 Time of initiation using the Life-365 (2000) software for surface chloride concentration at the end of two years equal to 1.0 kg/m3 (1.7 lb/yd3). ................................................................................................. 196 Figure 8.20 Time of initiation using the SRC method for surface chloride concentration at the end of two years equal to 1.0 kg/m3 (1.7 lb/yd3)........ 198 Figure 8.21 Time of initiation using the Life-365 (2000) software for surface chloride concentration at the end of two years equal to 2.5 kg/m3 (4.2 lb/yd3). ................................................................................................. 199 Figure 8.22 Time of initiation using SRC method for surface chloride concentration at the end of two years equal to 2.5 kg/m3 (4.2 lb/yd3)........ 200 Figure 8.23 Time of initiation using the Life-365 (2000) software for surface chloride concentration at the end of two years equal to 5.0 kg/m3 (8.3 lb/yd3). ................................................................................................. 202 Figure 8.24 Time of initiation using SRC method for surface chloride concentration at the end of two years equal to 5.0 kg/m3 (8.3 lb/yd3)........ 203 Figure 8.25 Conceptual diagram showing the relationship between the time to corrosion initiation and other material and environmental characteristics. 204
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Page Figure 8.26 Critical cracking and spalling threshold thickness, Tcrit, steel, as a function of the water-cement ratio.............................................................. 207 Figure 8.27 Inverse logarithm of Tcrit , steel as a function of water-cement ratio............. 207
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LIST OF TABLES Page Table 2.1 Cathodic reactions in presence of acidic, neutral, and alkaline electrolytes ... 13 Table 3.1 Activation energies of chloride diffusion process in ordinary portland cement pastes................................................................................................... 36 Table 3.2 Critical chloride threshold values, obtained from literature, for conventional carbon steels .............................................................................. 56 Table 3.3 Critical chloride threshold values, obtained from literature, for nonconventional carbon steels ........................................................................ 57 Table 3.4 Critical cracking and spalling threshold thickness values, Tcrit , steel obtained from the literature............................................................................................ 71 Table 4.1 Life expectancy for various repair and rehabilitation methods (Sprinkel et al. 1991 and Koch et al. 2001)..................................................... 83 Table 4.2 Construction costs of various construction materials (Darwin et al. 2002 and Trejo 2003) ............................................................................................... 87 Table 4.3 Costs related to various repair and rehabilitation methods or works (Sprinkel et al. 1991, Koch et al. 2001, and Darwin et al. 2002) .................... 88 Table 5.1 Heat generation due to applied voltage (after El-Belbol and Buenfeld 1989) ................................................................................................ 97 Table 5.2 Half-cell potential Vs probability of corrosion occurrence (ASTM C876-91 1999) ................................................................................. 100 Table 7.1 Average chloride concentration gradients across 2 mm depth at the surface of steel specimen............................................................................... 140 Table 7.2 Chemical composition of Type I cement used in the ACT test program (Trejo and Pillai 2003a)................................................................................. 141 Table 7.3 Chemical composition (as weight %) of steel reinforcement used in ACT tests (after Trejo and Pillai 2003a and Trejo and Pillai 2003b) ........... 143 Table 7.4 The sample population for the ACT test program (after Trejo and Pillai 2003a and Trejo and Pillai 2003b)....................................................... 145
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Page Table 7.5 Mixture proportion of the concrete used for ACST test program ................. 153 Table 7.6 Air content and slump of concrete used for the ACST test program ............ 153 Table 7.7 Average compressive strength, MPa (psi), of the concretes with different water-cement ratios used for ACST tests ....................................... 154 Table 7.8 The ACST test matrix.................................................................................... 155 Table 8.1 Critical chloride threshold values of ASTM A706 "as received" steel reinforcement (after Trejo and Pillai 2003a)................................................. 158 Table 8.2 Critical chloride threshold values of ASTM A706 "polished" steel reinforcement ................................................................................................ 159 Table 8.3 Critical chloride threshold values of ASTM A615 "as received" steel reinforcement (after Trejo and Pillai 2003a)................................................. 163 Table 8.4 Critical chloride threshold values of ASTM A615 "polished" steel reinforcement ................................................................................................ 164 Table 8.5 Critical chloride threshold values of microcomposite "as received" steel reinforcement (after Trejo and Pillai 2003b)................................................. 169 Table 8.6 Critical chloride threshold values of microcomposite "polished" steel reinforcement ................................................................................................ 170 Table 8.7 Critical chloride threshold values of stainless steel 304 "as received" steel reinforcement (after Trejo and Pillai 2003b) ........................................ 174 Table 8.8 Critical chloride threshold values of stainless steel 304 "polished" steel reinforcement ................................................................................................ 175 Table 8.9 Critical chloride threshold values of stainless steel 316LN "as received" steel reinforcement (after Trejo and Pillai 2003b) ........................................ 178 Table 8.10 Critical chloride threshold values of stainless steel 316LN "polished" steel reinforcement ........................................................................................ 179 Table 8.11 Mean and standard deviation of critical chloride threshold values of all the steel reinforcement types tested with ACT test procedure (after Trejo and Pillai 2003a and Trejo and Pillai 2003b)............................. 182 Table 8.12 The variables for parametric study .............................................................. 185
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Page Table 8.13 Chloride diffusion coefficients used for the parametric study .................... 186 Table 8.14 Surface chloride concentrations, due to de-icing salt application, for the bridge decks in different cities in North America (Life-365 2000)......... 188 Table 8.15 Time to corrosion initiation (years) calculated using Life-365 (2000) software (decay constant, m =0.4) ................................................................ 194 Table 8.16 Time of initiation (years) calculated using the SRC method ...................... 195 Table 8.17 Experimentally obtained critical cracking and spalling threshold thickness values for a concrete cover of 19 mm (0.75 inches) ..................... 205 Table 8.18 Assumed critical cracking and spalling threshold thickness values for a concrete cover of 65 mm (2.5 inches) ........................................................ 208 Table 8.19 Corrosion rate values for various steel reinforcement................................. 210 Table 8.20 Duration of propagation phase for different steel reinforcement and water-cement ratios ....................................................................................... 211 Table 8.21 Cost of a new 8.5 inch bridge deck reinforced with different types of steels (after Darwin et al. 2002 and Trejo 2003)........................................... 213 Table 8.22 Life cycle cost (in $/m2) of RC structures exposed to a low chloride environment corresponding to a surface chloride concentration at the end of two years equal to 1 kg/m3 (1.7 lb/yd3); design life = 100 years ....... 214 Table 8.23 Life cycle cost (in $/m2) of RC structures exposed to a moderate chloride environment corresponding to a surface chloride concentration at the end of two years equal to 2.5 kg/m3 (4.2 lb/yd3); design life = 100 years .............................................................. 215 Table 8.24 Life cycle cost (in $/m2) of RC structures exposed to a high chloride environment corresponding to a surface chloride concentration at the end of two years equal to 5.0 kg/m3 (8.3 lb/yd3); design life = 100 years .............................................................. 216
1
1 1.1
THE INTRODUCTION1
BACKGROUND Premature deterioration of reinforced concrete (RC) structures resulting from
exposure to aggressive environments is a serious challenge facing civil engineers, designers, contractors, and owners. Highway bridges, marine structures, and parking garages are typical examples of structures facing premature deterioration. The two main causes of structural damage to RC structures are degradation of the cementitious material and corrosion of the embedded steel reinforcement. Corrosion of steel reinforcement in bridge structures has been recognized as the largest overall maintenance cost in the United States infrastructure. The annual direct cost of corrosion for highway bridges is estimated to be $8.3 billion (Koch et al. 2001). Life cycle cost analyses estimate that the indirect cost to the user due to traffic delays and lost productivity is more than 10 times the direct cost of corrosion (Koch et al. 2001). New technologies, which are yet to be utilized, may help to reduce this huge economic loss. 1.1.1
Chloride-induced corrosion in concrete: Causes The most common cause of initiation and propagation of steel reinforcement
corrosion in RC structures is the presence of chlorides, mostly from seawater and deicing salts (e.g., sodium chloride, calcium chloride, and magnesium chloride). potential source of chlorides are admixtures containing chlorides. described chloride ions as "a specific and unique destroyer". This document follows the style and format of ACI Materials Journal.
Another
Verbeck (1975)
2
1.1.2
Chloride-induced corrosion in concrete: Remedies Two of the several strategies to improve the resistance of RC structures against
chloride-induced corrosion are using high performance cementitious materials, and high performance steel reinforcement. High performance cementitious materials slow the transport rate of chloride ions towards the steel reinforcement, thereby delaying the onset of corrosion. High performance steel reinforcement resist corrosion activity by requiring higher concentrations for activation of corrosion, thereby extending the service life of the structure. 1.1.2.1 High performance cementitious materials The use of high performance cementitious materials can improve the resistance of RC structures against chloride-induced corrosion and the resulting premature structural failure in two different ways: • •
by retarding the rate of chloride ingress in concrete, and by retarding the rate of corrosion of reinforcement.
Retarding the chloride ingress rate in concrete can increase the time required to attain sufficient chloride concentrations at the reinforcement level to initiate corrosion. This retardation in chloride transport rate can be achieved by densifying the microstructure of concrete. Dense microstructures can be achieved by various methods. Some of these methods include using dense, durable concrete with supplementary cementitious materials (e.g., fly ash, slag, silica fume, and metakaolin) (Dehghanian and Arjemandi 1997, Thomas, and Bamforth 1999, Thomas and Matthews 2003). It has been well documented that chloride ions can penetrate faster through cracked concrete than
3
through uncracked concrete (Wang et al. 1997). Hence, keeping the concrete free of cracks can delay the chloride ingress rate. The lower the corrosion reaction rate, the higher will be the time to corrode a specific amount of steel.
Corrosion inhibitors can effectively retard the corrosion
reaction rate by altering the chemical mechanisms in concrete (Trepanier et al. 2001, Saricimen et al. 2002, Al-Amoudi et al. 2003). The use of corrosion inhibitors can be effective in reducing the corrosion reaction rate even in concrete with high chloride contamination levels. Also, some corrosion inhibitors have been reported to retard the chloride ingress rate resulting in a delayed initiation of corrosion (Kondratova et al. 2003). Formation of corrosion products cause expansive stresses on the concrete cover. When these expansive stresses exceed the tensile strength capacity of the concrete, cracking and spalling of concrete cover occurs. Balabanic et al. (1996) and many others have reported that a reduction in water-cement ratio will result in an increased effective tensile strength capacity of concrete cover. It is well documented that increased cover depth will not only result in longer time requirement for chloride ions to reach the embedded reinforcement to start corrosion but will also require more corrosion products to cause cracking and spalling of the concrete cover (Balabanic et al. 1996). 1.1.2.2 High performance steel reinforcement The steel industry is manufacturing various types of reinforcing steels, each with unique corrosion and strength performance characteristics. It is well documented that the high performance steel reinforcement (i.e., steel with improved corrosion resistance) can
4
significantly improve the corrosion resistance of RC structural elements (Trejo et al. 2000).
The initial material cost may be higher for corrosion resistant steels when
compared to conventional carbon steels. But, because of improved resistance to chlorideinduced corrosion, some high performance reinforcing steels may be more cost effective, based on average life cycle costs. Moreover, for the same repair method, the repair frequency will be less for corrosion resistant steels than that for conventional carbon steels. This is again attributed towards the faster corrosion of conventional carbon steels when compared with corrosion resistant steels. Knudsen et al. (1998) reported that for discount rates below 7% (often used by bridge designers while selecting rehabilitation strategies), repairs using stainless steel are more economical than that using conventional carbon steel or cathodic protection. Thus, the service life can be increased and life cycle cost (LCC) may be reduced if high performance cementitious materials as well as high performance steel reinforcement are used in RC structures exposed to corrosive environments.
Key
material parameters to determine the service life and LCC are needed to assist engineers in selecting optimal strategies for selecting materials. 1.1.3
Critical parameters for service life prediction and life cycle cost analysis Corrosion of steel reinforcement causes the concrete surface to crack and spall,
resulting in reduced service life times. Three key parameters, the minimum chloride concentration required to initiate corrosion of steel reinforcement, the amount of steel reinforcement corrosion required to trigger surface cracking and spalling of the concrete
5
cover, and environmental exposure conditions are needed to predict the service life of RC structures. Available standard test methods for determining the corrosion characteristics of steel reinforcement embedded in concrete do not specifically evaluate these key parameters and can take years or decades to complete, making these methods uneconomical and impractical. Standardized short-term test methods to evaluate the corrosion characteristics of steel reinforcement embedded in cementitious materials are not yet available. This lack of reliable quantitative data makes decision makers hesitant towards using the new durable corrosion resistant steel reinforcement. Thus there is an urgent need to develop short-term test methods that provide quantitative data on the critical chloride threshold value for steel reinforcement embedded in cementitious materials and the amount of corrosion required to crack or spall the concrete cover, especially when the steel industry is producing various types of reinforcing steels, each having unique corrosion performance characteristics. This quantitative data, required to predict the service life and life cycle cost will assist designers in making better decisions in selecting cost effective construction materials during the design stage for RC structures. 1.2
PROBLEM STATEMENT AND RESEARCH OBJECTIVES The purpose of this research program is to study, using accelerated test
procedures, the influence of steel reinforcement types, water-cement ratios, and cover depths on the overall serviceability and life-cycle cost of RC structural systems exposed to chloride environments.
6
Various objectives of this study are: • • • • •
to evaluate and perform engineering refinement of the Accelerated Chloride Threshold (ACT) test methodology originally developed by Trejo and Miller (2002), to quantitatively determine the critical chloride threshold values of different uncoated steel reinforcement types embedded in a standard cementitious material using the ACT test methodology, to develop an Accelerated Cracking and Spalling Threshold (ACST) test method, to quantitatively determine the critical cracking and spalling threshold thickness of concrete cover using the ACST test methodology, and to study the effect of the quality of both the steel reinforcement and concrete cover on the service life and life cycle cost of RC structures exposed chloride environments.
Recommendations on selecting durable construction materials for reduced life cycle cost of RC structures will be presented. 1.3
THESIS ORGANIZATION This thesis includes 9 sections and several subsections. Section 1 introduces the
background to the magnitude of the problems associated with corrosion-induced deterioration of RC structures exposed to chloride environments. An introduction on how this premature deterioration and the resulting economic loss can be curbed or controlled is provided. The urgent need for developing standardized short-term test methods for efficient, reliable, and quantitative determination of critical parameters for predicting service life of RC structures exposed to chloride environments is emphasized. Section 2 is comprised of a brief review of basic principles and mechanisms of electrochemical corrosion of metals in aqueous solution environment. Thermodynamic and kinetic principles are discussed.
7
Section 3 provides a comprehensive review of the principles and mechanisms of chloride-induced corrosion of steel reinforcement embedded in concrete. Mechanisms such as diffusion based transport of chloride ions in concrete, and the formation and breakdown of protective layers on the embedded steel reinforcement are presented. A review of critical chloride threshold values for different steel reinforcement types, cracking and spalling threshold thickness of for various concrete design parameters and other issues is provided. Section 4 presents mathematical models for predicting the service life of RC structures exposed to chloride environments. A brief review of life cycle cost analysis models is also provided. Section 5 presents a discussion on different electrical, electrochemical and chemical test methodologies available for determining critical service life parameters of RC structures exposed chloride environments. Section 6 emphasizes the significance and necessity for the development of shortterm test methodologies required to determine the critical chloride threshold level of steel reinforcement and critical cracking and spalling threshold thickness for concrete cover. The quantitative information on these parameters can be used for the prediction of service life and life cycle cost of RC structures exposed to chloride environments. Section 7 presents the experimental program followed in this thesis for determining the critical chloride threshold values and cracking and spalling threshold thickness values of uncoated steel reinforcement embedded in cementitious materials. This section also includes a description and evaluation of the new accelerated test
8
methods used in the experimental program to evaluate the corrosion performance of steel in cementitious materials. Section 8 provides a detailed discussion on the results of the testing program. These results include the critical chloride threshold values and cracking and spalling threshold thickness values obtained from the experimental programs explained in section 7. Finally a parametric study on the service life and life cycle costs of RC structures with different construction materials are provided. Section 9 provides conclusions and recommendations for future research.
9
2 2.1
BASICS OF ELECTROCHEMICAL CORROSION
INTRODUCTION Corrosion is defined as the deterioration of a substance (usually a metal) or its
properties because of a reaction with its environment (NACE 1971). Corrosion can also be defined as extractive metallurgy in reverse (Fontana and Greene 1978). Many metals undergo corrosion and form corrosion products. Corrosion of metals depends on various physical and chemical characteristics of the metal. These characteristics include the chemical composition, the amount of stored energy, the electrical potential, the atomic or molecular structure, and other characteristics. Almost all metals and their alloys can corrode. But, the kinetics of corrosion of these metals and their alloys when exposed to similar or identical corrosive environments may vary depending on their specific material characteristics.
Also,
corrosion is dependent on the physical and chemical characteristics of the environment. All environments are not equally corrosive. Inorganic materials are more corrosive than organic materials. Extraction of every metal from its ore consumes some amount of energy. This energy is stored in the metal. The amount of this stored energy is different for different metals. The process of corrosion of a metal is a chemical reaction in which a metallic compound that has a similar chemical composition as of its ore, is being produced. This process releases energy that is equal to the amount of energy that was required to extract or produce the metal from its ore. The higher the energy required for the extraction of the
10
pure base metal from its ore, the higher will be the resistance of the pure base metal against corrosion. 2.2
FORMS OF CORROSION Different forms of corrosion have been identified mainly based on the pattern of
corrosion, the location of metals or their access to the corrosive environment, nature of metals or alloys involved and their relative location, and other parameters. Uniform corrosion, pitting corrosion, crevice corrosion, galvanic corrosion, intergranular corrosion, stress corrosion cracking, and hydrogen embrittlement are some typical forms of corrosion. The metallic corrosion can also be classified into two main categories: general corrosion and localized corrosion. A simplified schematic of these two types of corrosion is shown in Figure 2.1.
Corrosion products Base metal General corrosion
Localized corrosion
Figure 2.1 General and localized corrosion.
2.2.1
General corrosion General corrosion occurs when all parts of the corroding metal are exposed
equally to the corrosive environment. There are numerous anode and cathode areas on a metal surface and these areas are very minute making it almost inseparable (NACE
11
1971). These anode and cathode areas may shift from time to time. This time dependent process of switching back and forth from an anode to a cathode, and vice versa, may occur throughout the reinforcement surface area resulting in general corrosion. In the case of general corrosion, corrosion occurs continuously over substantial areas of the metal surface with metal loss fairly evenly distributed over the exposed area (Vassie 1984). Uniform corrosion is an example of general corrosion. 2.2.2
Localized corrosion Localized corrosion occurs when there is only localized access to the corrosive
environment. Material inhomogeneity and surface geometry (i.e., corners) are other causes for localized corrosion.
The presence of chlorides or other halides in the
electrolyte adjacent to the metal surface results in local breakdown of protective layers on the exposed metallic surface.
This local breakdown of the protective layers and
continued corrosion thereafter at these locations results in localized corrosion. Pitting corrosion and crevice corrosion are typical examples of localized corrosion. 2.3
MECHANISMS OF CORROSION Material deterioration occurs due to the actual atomic, molecular, or ionic
transport processes that take place at the material-environment interface. In other words, the flow of electric charge in a closed circuit causes corrosion. There is no gain or loss of electric charge during the electrochemical corrosion. The flow of electric charge can occur across two regions on the same metal surface, across two regions on two different metal surfaces, or across one region on a metal surface and another region in a metal
12
recipient. The two regions between which the flow of electric charge occurs are referred to as the anode and the cathode. This flow of electric charge between the anode and cathode occurs only when there is an electrochemical potential difference and there is a conductive medium connecting these two regions. This conductive medium, known as the electrolyte, can be a fluid, a granular solid, or a more complex solid-fluid combination. Saturated or partially saturated concretes are examples of such electrolytes with complex solid-fluid combinations. For electrolyte systems with constant resistance, higher potential differences will result in higher current flows and vice versa. Therefore, three necessary components of an electrochemical corrosion cell are the anode, the cathode, and the electrolyte. The anode is the area or the region on the metal surface where oxidation, or the release of valence electrons from metal atoms, takes place. The general form of a typical anodic reaction is as follows:
M → M n + + ne −
(2.1)
Here, the metal atom is oxidized and becomes a positively charged ion. For instance, iron atoms ( Fe ) can be oxidized to ferrous ions ( Fe 2+ ) by releasing two electrons. These electrons released from the metal atoms in the anode region migrate through the metal towards the cathode region. At the same time, the positively charged metal ions (i.e., Fe2+ and Fe3+ ) are released from their lattice sites and combine with negatively charged ions (i.e., Cl − and (OH ) − ) present in the electrolyte forming the corrosion products. There can be several intermediate reaction steps before the formation
13
of the final corrosion product. The corrosion or deterioration of the metal actually occurs at the anodic region on the metallic surface. The cathode is the area or the region on the metal surface where the consumption of electrons takes place. Table 2.1 shows typical cathodic reactions as a function of pH and oxygen availability in the electrolyte.
Table 2.1 Cathodic reactions in presence of acidic, neutral, and alkaline electrolytes Oxygen Content Low availability
Acidic electrolyte 2 H + + 2e− → H 2 ↑
Neutral and alkaline electrolyte 2 H 2O + 2e− → H 2 ↑ +2(OH ) −
High availability
O2 + 4 H + + 4e− → 2 H 2O
O2 + 2 H 2O + 4e− → 4(OH ) −
There are cases where an already oxidized metal ion undergoes a further reduction reactions (i.e., consumes more electrons) and becomes a more negatively charged metal ion. A general form of such reduction reactions is as follows: M m + + (m − n)e− → M n + ( where, m > n)
(2.2)
For instance, ferric ions (Fe3+) can be reduced to form ferrous ions (Fe2+) as follows: Fe3+ + e − → Fe 2+
(2.3)
Also, there are cases where an already oxidized metal ion undergoes a reduction reaction (i.e., consumes more electrons) and becomes a metal atom. A general form of such reduction reactions is as follows:
M n + + ne − → M
(2.4)
14
Depending upon the chemical constituents of the metal and the electrolyte, the corrosion process is a combination of one or more types of anodic and cathodic reactions. For instance, the corrosion of an alloy can include the oxidation and reduction of various metallic components of the alloy. 2.4
THERMODYNAMICS OF CORROSION
In Subsection 2.3, a general introduction to the electrochemical reactions associated with a metal-solution system was presented.
Thermodynamic principles
cannot predict the rate of electrochemical reactions, which is governed by the laws of kinetics discussed in Subsection 2.5. In this subsection some thermodynamic principles that can only predict the occurrence of electrochemical reactions, will be presented. 2.4.1
Electrochemical potential of corrosion reactions
Every electrochemical reaction has a unique potential, referred to as the electrochemical potential, and is a measure of the tendency of the corrosion reaction to occur. Every electrochemical reaction can be split into various half-cell reactions that occur at the anodic and the cathodic regions of the corrosion cell. The algebraic sum of the electrochemical potentials of individual half-cell reactions occurring at the anode, Ea , and the cathode, Ec , is equal to the electrochemical potential of the cell reaction, Ecell , and can be mathematically expressed as follows: Ecell = Ea + Ec
(2.5)
The absolute value of the half-cell reaction potential is impossible to measure. Only the half-cell potential difference can be measured against some standard reference
15
electrode potentials. The half-cell potential difference, under reversible conditions, forms the basis for predicting the tendency of a reaction to occur (Fontana and Greene, 1978). The electromotive force (EMF) series, available in many electrochemistry textbooks (NACE 1971, Uhlig and Revie 1985, and Jones 1996), is an orderly listing of electrochemical potentials of various half-cell reactions with reference to the potential of the hydrogen half-cell reaction: 2 H + + 2e − → H 2 ↑
(2.6)
For simplicity, the potential of this reaction, under reversible conditions, is arbitrarily defined as zero. The EMF of a corrosion cell is dependent on the free energy change of the anodic and cathodic reactions involved. The free energy change is in turn dependent on the activity (or concentration) of the reactants and products involved in the anodic and cathodic reactions. Following is an illustration on how free energy change and activity can be used in determining the EMF of corrosion cells. 2.4.1.1 Activity and Gibbs free energy
The activity of a reactant or a product in a chemical reaction is defined as its concentration in the chemical system and environment. The activity is expressed in terms of pressure (atmospheres) for gaseous systems, and concentration (gram equivalents) for aqueous solution systems (Jones 1996). In a standard state, the activity of reactants and products are defined as unity, whereas in non-standard states they tend to deviate from unity causing reactions to occur. Activity of the reactants and products continuously
16
change in a system. Hence, it is important to know the effect of changes in activity of the reactants and products. Free energy, also known as Gibbs free energy, of a system is the energy available from the system to do useful work. The free energy of a thermodynamic system is constant if both pressure and temperature of the system are constant and there is no other work done on or by the system other than the product of the pressure and volume change (Pauling 1988). This free energy change and electrochemical potential of corrosion reactions can be related by a fundamental work-energy relationship. 2.4.1.2 The fundamental work-energy relationships
There occurs a change of Gibbs free energy, ∆G , associated with any chemical reactions when the products exhibit a higher or lower energy than the reactants. The amount of work done on or by a system when it undergoes a reversible change in state at constant temperature and pressure is equal to the change in free energy of the system. This fundamental work-energy relationship can be mathematically expressed as follows: ∆G ° = − nFE °
(2.7)
∆G = −nFE
(2.8)
where, ∆G ° and ∆G are Gibbs free energy of the reaction at the standard and non-standard state, respectively, ∆E ° and ∆E are the EMFs of the reaction at the standard and non-standard state, respectively, n is the number of electrons (or electrochemical equivalents) exchanged in the reaction, and F is Faraday’s constant. To illustrate the use of Gibbs free energy and activity of reactants and products in determining the EMF of a corrosion cell, the following half-cell reactions are considered.
17
Fe → Fe 2+ + 2e− (anode)
(2.9)
O2 + 2 H 2O + 4e − → 4(OH ) − (cathode)
(2.10)
For a reaction with the reactants at activity ar and products at activity a p the Gibbs free energy change from standard to non-standard state, G − G° , can be mathematically expressed as follows (Pauling 1988): ⎡ ap ⎤ ⎡ ap ⎤ G − G° = RT ln ⎢ ⎥ = 2.303RT log ⎢ ⎥ ⎣ ar ⎦ ⎣ ar ⎦
(2.11)
where R is the universal gas constant (8.314 Jmoles-1K-1) and T is the absolute temperature. For a reaction with the reactants at activity ar and products at activity a p the EMF change from the standard to non-standard state, E − E ° , can be determined by substituting work-energy relationships into Equation (2.12) as follows: ⎛ 2.303RT ⎡a ⎤⎞ E − E ° = ± ⎜⎜ log ⎢ p ⎥ ⎟⎟ ⎣ ar ⎦ ⎠ ⎝ nF
(2.12)
On the right hand side of the equation, the sign is negative when electrons are released (anode reaction) and positive when electrons are consumed (cathode reaction). Assume an absolute temperature equal to 298oK and the activity of both the iron molecule and the water molecule in the electrolytic solution is equal to unity. The EMFs of the anode reaction ( Eao ) and the cathode reaction ( Eco ) at the standard state are +0.441V versus SHE and -0.401V versus SHE, respectively. The activity of ferrous ions is related to the concentration of ( Fe(OH ) 2 ) and ( H 2O) in the electrolyte adjacent to the steel.
18
Therefore, the EMF change of the anode reaction, Ea − Eao , during the change from the standard to the non-standard state is:
⎛ 2.303RT ⎡ ( Fe 2+ ) ⎤ ⎞ Ea − Eao = − ⎜ log ⎢ ⎥⎟ ⎣ ( Fe) ⎦ ⎠ ⎝ nF
(2.13)
Ea = +0.441 − 0.0296 log[( Fe 2+ )]
(2.14)
and the EMF change of the cathode reaction, Ec − Eco , during the change from the standard to the non-standard state is: ⎛ 2.303RT ⎡ ([OH ]− ) 4 ⎤ ⎞ Ec − E = + ⎜⎜ log ⎢ 1 2 ⎥⎟ ⎟ ⎣ (O2 ) .( H 2O) ⎦ ⎠ ⎝ nF
(2.15)
Ec = +1.229 − 0.0148log[(O2 )] + 0.0591 pH
(2.16)
o c
Substituting Equations (2.15) and (2.17) in Equation (2.6) and solving, the EMF of the corrosion cell involving Equation (2.10) and (2.11) is as follows: Ecell = 1.67 − 0.0296 log[ Fe 2+ ] − 0.0148log[O2 ] + 0.0591( pH )
(2.17)
The negative and positive signs of the coefficients of third and fourth terms, respectively, in Equation (2.18) indicates the following: •
•
At constant activities of Fe2+ and O2 , Ecell becomes more negative (i.e., higher tendency to corrode) as the pH decreases as shown in Figure 2.2. At constant activity of Fe2+ and at constant pH, Ecell becomes more negative (i.e. higher tendency to corrode) as the activity of the O2 increases as shown in Figure 2.3.
19
-0.2 -13
O concentration = 1x10 M 2
-0.3
2+
-13
Fe concentration = 1x10 M
E
cell
(Volts)
-0.4
-0.5
-0.6
-0.7 7
8
9
10
11
12
13
14
pH
Figure 2.2 The variation of Ecell as a function of pH.
-0.2 pH = 13 2+
-13
Fe concentration = 1x10 M
-0.4
E
cell
(Volts)
-0.3
-0.5
-0.6
-0.7 -16
-12
-8
-4
0
4
8
Log [O ] 2
Figure 2.3 The variation of Ecell versus common logarithm of O2 concentration.
20
The EMF of any corrosion cell with any combination of half-cell reactions can be obtained using the procedure mentioned in this subsection and can be used to predict corrosion tendencies. Due to the change of the corrosion cell reaction system from standard to nonstandard state, a change in Gibbs free energy of the system occurs. A positive value of Gibbs free energy change will result in a negative value of EMF change. Gibbs free energy change and EMF change of any half-cell reactions can be used to predict their tendencies to occur. A more positive value of Gibbs free energy change (more negative EMF) indicates a higher tendency of the corrosion reaction to occur, where as a more negative value of Gibbs free energy change indicates a lower tendency of the corrosion reaction to occur. Any half-cell corrosion reaction with a higher EMF (more positive value) is less active than another half-cell corrosion reaction with a lower EMF (more negative value). 2.5
KINETICS OF CORROSION
Electrochemical kinetics are the rate of the electrochemical reactions that take place at the electrode-electrolyte interface. In this subsection the definitions of two parameters, corrosion rate and current density, are presented. 2.5.1
Corrosion rate
Corrosion rates are very important in predicting or determining the amount of electrode that will corrode over a certain period of time. To determine this rate, Faraday's law can be used as follows:
21
m=
Ita nF
(2.18)
where m (grams) is the mass of the reacting species with an atomic number, a , I (Amps) is the current passed, t (seconds) is the time, n is the number of electrons (or
electrochemical equivalents) exchanged in the reaction, and F is Faraday’s constant (96,490 coulombs per equivalent). This relationship indicates the following (Pauling 1988): • •
The weight of a substance produced by a cathode or an anode reaction in electrolysis is directly proportional to the quantity of electricity passed through the cell. The weights of different substances produced by the same quantity of electricity are proportional to the equivalent weights of the substances.
The current passing, I (Amps), can be obtained by rearranging the Equation 2.19. The current density, i (Amp/m2 or Amp/ft2), is defined as the current passing across a unit area and can be determined as follows: i=
I nFm = A Ata
(2.19)
2.5.1.1 Average corrosion rate
The average corrosion rate, ravg , over a period of time can be determined as follows:
ravg =
m ⎛ kg g ⎞ ⎜ 2 or ⎟ tA ⎝ m s cm 2 s ⎠
(2.20)
where, m is the mass of the reacting species or product corroded, t is the time of exposure, and A is the area of exposure. Any system of units (i.e., SI or English) can be used, provided the equation is dimensionally correct. Here, the mass loss of the corroded
22
steel can be determined with destructive testing methods. Average corrosion rates can also be expressed in terms of the thickness of metal lost (or depth of corrosion) per unit of time (e.g., mils/year). Once the average corrosion rate from a sample specimen is known, the amount of steel corrosion at a future time can be predicted or determined, assuming a relationship between the corrosion rate and time of exposure. 2.5.1.2 Instantaneous corrosion rate
Instantaneous corrosion activity at a metal surface is a measure of the instantaneous current passing across the area of exposure. Current density is defined as the current passing across a unit exposure area. The instantaneous current density, determined using Equation (2.20), can be used to nondestructively determine the instantaneous corrosion rate, rins , as follows:
rins =
ia nF
(2.21)
where i (Amp/m2 or Amp/ft2) is the current density, a is the atomic number of the reacting species, n is the number of electrons (or electrochemical equivalents) exchanged in the corrosion reaction, and F is Faraday’s constant (96,490 coulombs per equivalent). Any system of units (i.e., SI or English) can be used. Assume that the current measured over two areas with different magnitudes are same. Current density is the ratio of current passed to area. Hence, the current density over the smaller area will be higher than that at the larger area, if same quantity of current is passing. This indicates that high corrosion activity may be ongoing at the smaller area than that at the larger area with same current passing.
Hence, measurement and
23
monitoring instantaneous current density values, or instantaneous corrosion rates, could be helpful in detecting and identifying the phase changes of corrosion activities on metals. 2.6
PROTECTIVE SURFACE BARRIERS
There are two types of protective layers that can prevent metal from corrosion. These protective layers can be formed before embedment in an electrolyte (or concrete), and after embedment in electrolyte (or concrete). Apart from the passive film formed in concrete during the curing period, there may be other thin films formed on the steel surface during the manufacturing process. During the cooling processes during steel manufacturing, a thin porous layer of metal oxides is formed on the steel surface. This layer, which is known as mill scale, can behave like a physical barrier against the ingress of chloride ions to the steel base. But, insufficient literature is available on the influence of this mill scale on corrosion characteristics of steel reinforcement. In addition to the mill scale, another film can form on the steel surface. This film, o
known as the passive film, is a very thin ( 10 A ) protective layer of iron oxides. There are two broadly accepted definitions for metal passivity. The first definition indicates that “... a metal is passive if it substantially resists corrosion in a given environment resulting from marked anodic polarization” (Uhlig 1978). The second definition states that “... a metal is passive if it substantially resists corrosion in a given environment despite a marked thermodynamic tendency to react” (Uhlig 1978).
24
There are two theories on passivity: the oxide film theory and adsorption theory. A passive film can act as a diffusion-barrier against the oxidizing agents, preventing the agents from reaching and reacting with the base metal. This is known as oxide film theory of passivity. The adsorption theory states that as oxygen molecules are adsorbed on the metal surface, water molecules are displaced creating a shortage of water molecules at the metal surface. Because, water (or moisture) is essential for corrosion process at the cathode and electrical neutrality must be achieved, the dissolution rate of positively charged metal ions formed during the anodic reaction, M → M n + + ne− , is retarded because of the unavailability of sufficient water molecules at the cathode. Whether the passive film prevents the movement of oxidizing agents towards the base metal or prevents corrosion by limiting the availability of water to the cathode, it has been well established that the passive films forms a protective film that resists corrosion.
25
3
MECHANISMS OF CHLORIDE-INDUCED CORROSION IN CONCRETE When RC structures are exposed to chloride environments, chloride ions migrate
through the concrete cover towards the embedded steel reinforcement.
It is well
documented that the reinforcement starts corroding when the chloride concentration in the concrete adjacent to the reinforcement attains a certain chloride threshold concentration. Once initiated, the corrosion process continues, provided both oxygen and moisture are present in sufficient quantity at the steel-cementitious material interface. Continued formation of corrosion products generates expansive stresses in concrete. These expansive stresses cause the concrete cover to crack and spall, adversely affecting the strength and serviceability of RC structures. To provide the reader with a thorough understanding of the corrosion process of steel reinforcement in cementitious materials, a review of the following topics is provided in this section.
• • • • •
Mechanism of chloride migration in concrete Mechanism of chloride-induced corrosion in concrete Chloride concentration threshold parameters Mechanism of corrosion-induced cracking and spalling of concrete cover Cracking or spalling threshold parameters
The nature and characteristics of chloride environments can be varying from location to location. For marine structures, the main source of chlorides is seawater. The submerged portions of the marine structures are under continuous exposure to chlorides whereas the splash zone areas on the structures experience a cyclic wet-dry exposure to
26
chlorides. When exposed to seawater, the chlorides slowly penetrate into concrete and can cause reinforcement corrosion. For bridge superstructures, the major source of chlorides is from de-icing/anti-icing salts that are applied seasonally.
The rate of
application of de-icing/anti-icing salts varies from region to region depending on the climatic conditions. Once the de-icing/anti-icing salts are applied, the chloride ions slowly penetrate into the concrete and can cause reinforcement corrosion. The material characteristics of concrete influence the rate of chloride transport into concrete. Concrete can be classified into two categories: uncracked concrete and cracked concrete. The mechanisms of chloride transport are different for these two types of concrete leading to different rates of chloride transport. 3.1
CHLORIDE PENETRATION IN UNCRACKED CONCRETE
The penetration rate of chloride ions in concrete is a dynamic process that is influenced by various physical and chemical parameters. Diffusion, permeation, and sorption are some of the chloride transport mechanisms in concrete. The presence of hydrostatic pressure causing chloride permeation in concrete occurs in rarer cases. Even so, Boddy et al. (1999) reported that failure to consider the minor chloride penetration mechanisms such as sorption, permeation, wicking, and chloride binding could result in erroneous determination of chloride penetration depth and concentration. At the same time, many researchers including Atkinson and Nickerson (1984), Goto and Roy (1981), Feldman et al. (1993), Yu and Page (1991), Feldman et al. (1994), Gj ∅ rv et al. (1994), and Dehghanian and Arjemandi (1997), agreed that diffusion is the most important
27
mechanism contributing to chloride penetration in concrete. Hence, only diffusion based chloride penetration mechanisms in concrete will be discussed in this thesis. Subsection 3.1.1 provides a brief review of the chloride diffusion mechanism in concrete. In Subsection 3.1.2 some diffusion based chloride transport models to predict time dependent chloride concentrations at specified locations within the concrete matrix is provided. 3.1.1
Diffusion of chloride ions in concrete
Diffusion is defined as the movement of species or particles from one region of higher concentration or density to another region of lower concentration or density. Figure 3.1 shows a schematic of a general diffusion process.
Diffusant
Diffusing medium
Direction of Diffusion
High ionic concentration
Low ionic concentration
Figure 3.1 A schematic of the general diffusion process.
Chloride ions accumulate on the concrete surface when de-icing salts are applied to the concrete surface, when exposed to seawater, or by other means. Over the course of time these chloride ions diffuse into regions within the concrete where the chloride
28
concentration is much lower than that at the concrete surface. Accurate determination of the chloride diffusion coefficient or diffusivity in concrete is very important in predicting chloride ingress behavior or chloride profiles in concrete. The diffusion of chlorides in concrete is a very slow and complex process.
In general, the chloride diffusion
coefficient in concrete is on the order of magnitude of 10-12 m2/s, but has been documented to vary significantly depending on various factors. Following is a discussion on some of these influencing factors. 3.1.1.1 Effect of water-binder ratio
The water-cement ratio (or water-binder ratio) has a significant effect on the concrete pore structure and its formation (Soroka 1979). An increase in water-cement ratio can result in a increased porosity. Increased porosity could result in increased interconnectivity of pores in concrete. Increased interconnectivity of pores will result in an increase in the ingress rate of chloride ions into concrete. It has been reported that an increase in the water-cement ratio from 0.4 to 0.6 resulted in a 2.2 times increase in the chloride diffusion coefficient in concrete (Sugiyama et al. 1996).
Dehghanian and
Arjemandi (1997) reported that chloride diffusivity was minimum for slag blended cement concrete with 0.45 water-cement ratio and increased with water-cement ratios greater than 0.55. These findings clearly indicate that a concrete with higher watercement ratio can have higher chloride diffusion coefficients. In 2002, Thoft-Christensen reported that higher water-cement ratios could lead to less binding of chloride ions. Less binding of chloride ions result in a higher amounts of available free chlorides to either diffuse into the concrete or to react with the steel
29
reinforcement. In general, the chloride ingress rate increases with increasing watercement ratio for the same cement content and other ingredients. Complex chemical reactions are involved during the process of cement hydration. Page et al. (1981) and Goto and Roy (1981) reported that the ionic diffusion rates in hardened cement pastes could be influenced if there exists an electric double layer region between the walls of micro pores in concrete and the pore solution. Later this view was proved correct when it was found that the calcium silicate hydrate (CSH) gel formed during the cement hydration phase exhibited negatively charged surface particles, implying the existence of an electric double layer in the vicinity of the concrete pore walls (Nagele 1987 and Chatterji and Kawamura 1992). The water-cement ratio has a great influence in the pattern and formation of these CSH gel structures in concrete. Hence, the water-cement ratio controls the structure and formation of these electric double layers, which influences the diffusion of chloride ions in concrete. In field concreting, workability is an important factor that must be considered. Concrete workability typically decreases with a decrease in the water-cement ratio. Engineers and designers must consider workability issues and how these relate to concrete quality. The workability of concrete with low water-cement ratios can be increased by using chemical admixtures, plasticizers, or by improving the aggregate grading. 3.1.1.2 Effect of cement type and supplementary cementitious materials
The main constituents of portland cement are dicalcium silicate ( C2 S ), dicalcium silicate ( C3 S ), tricalcium aluminate ( C3 A ), tetracalcium aluminoferrite ( C4 AF ). During
30
hydration of cement many complex compounds are formed.
The pore structure of
hardened concrete depends on the microstructure of these hydration products.
The
hydration products have different chemical characteristics leading to varying reaction mechanisms with penetrating chlorides. Chloride ions can exist in both free and bound (chemically or physically) states and these states can change from with time. In 1977, Traetteberg reported that chloride intrusion is not affected by either tricalcium aluminate ( C3 A ) or the combination of C3 A and tetracalcium aluminoferrite ( C4 AF ) present in the cement paste. It is generally accepted that only free chlorides can diffuse. Hansson and Sorenson (1990) proved that Traetteberg's (1977) assumption was flawed by finding that the increase in C3 A content of cement could effectively reduce the amount of free chlorides available to diffuse into concrete. This reduction in the amount of free chlorides available is attributed to the fact that an increase in C3 A results in an increase in the amount of bound chlorides in concrete (Rasheeduzzafar et al. 1990 and Rasheeduzzafar 1992). This reduced amount of free chlorides, coupled with the fact that only free chlorides can diffuse, will effectively result in a reduced amount of chloride diffusion. The addition of industrial waste products such as fly ash (FA), silica fume (SF), slag, and other supplementary cementitious materials in concrete can alter the concrete microstructure, porosity, and other characteristics. The changes in microstructure and chemistry of the cement paste caused by the addition of supplementary cementitious materials properties can influence the chloride ingress rate in concrete.
Thus, the
31
addition of supplementary cementitious materials can influence the chloride ingress rate in concrete. Khatib and Mangat (2002) found that replacement of cement with 22% FA and 9% SF in concrete significantly increased the resistance against chloride penetration, especially at depths greater than 20 mm and 10 mm. Thirty percent or more replacement of cement by pulverized fuel ash (PFA) was found to significantly reduce both the chloride penetration and penetration depth (Bai et al. 2003, Thomas and Matthews 2003). Bai et al. (2003) also found that in blends of portland cement-PFA-Metakaolin, increasing metakaolin content will reduce the chloride penetration and penetration depth. In general, it can be concluded that the chloride ingress rate for concretes with blended cement is less than that for concrete with conventional cement. Hence, the use of supplementary cementitious materials can delay the increase of chloride ion concentration at the reinforcement level for uncracked concrete. 3.1.1.3 Effect of aggregates
The presence of fine and coarse aggregates tends to modify the microstructure of concrete and influences the chloride ingress behavior. The presence of aggregates can both retard and enhance the chloride transport rate. It is relatively difficult for the chloride ions to diffuse through the dense aggregates. As such, the majority of chlorides are transported through the cement paste phase. The tortuosity of the path followed by an external species (i.e., chloride ions) increases as the length of the actual path followed while traveling through the material increases. As aggregates are introduced to the cement paste, the tortuosity of the concrete
32
matrix increases. This is due to the fact that the chloride ions have to follow a longer path around the dense aggregates.
The dense aggregates act as a physical barrier,
hindering the process of chloride diffusion. Contrary to the above discussion, the presence of an interfacial transition zone (ITZ), a very thin layer with a typical thickness of 20-30 µm (Delagrave et al.1997) that forms between the hydrated cement paste and the aggregate surface, may accelerate the migration process. Figure 3.2 shows a schematic of ITZ.
Fine Aggregate Coarse Aggregate ITZ
Figure 3.2 The interfacial transition zones (ITZ) around aggregates in concrete.
The ITZs are typically, for higher water-cement ratio, more porous than the bulk cement paste and are most often interconnected with each other. The water-cement ratio has a significant influence on the porosity of ITZ (Soroka 1979). Highly porous and interconnected ITZs serve as a continuous pathway for chloride ions and can increase the chloride diffusion coefficient in concrete. Delagrave et al. (1997) found that the chloride
33
diffusion coefficient is influenced more by the tortuosity of concrete matrix than by the interconnectivity of ITZs. Delagrave et al. (1997), Yang and Su (2002), and other researchers have provided guidance and mathematical models for determining the influence of aggregates on the chloride diffusion coefficient in concrete. For instance, Xi and Bazant (1999) used a composite sphere model, provided by Christensen (1979), to determine the effective diffusivity of saturated concrete. This relationship does not consider the effect of ITZs instead considers the volume fraction of aggregates and the diffusion coefficients of constituents of concrete. If the volume fraction of aggregates, gi , aggregate diffusivity,
Di , and cement paste diffusivity, Dm , are known, then the effective diffusivity of chlorides in concrete, Deff , can be determined as follows (Xi and Bazant 1999 and Christensen 1979):
gi ⎪⎧ ⎪⎫ Deff = Dm ⎨1 + ⎬ ⎩⎪ [ (1 − gi ) / 3] + [ Dm /( Di − Dm ) ] ⎭⎪
(3.1)
It is clear that the presence of aggregates and other ingredients influences the chloride diffusion behavior in concrete. At the same time, it is not only the presence of but also the degree of compaction or consolidation of aggregates and other ingredients that influence the chloride diffusion behavior in concrete. 3.1.1.4 Effect of compaction and consolidation
Dense aggregates alone can act as a wall or physical barrier in stopping or retarding the ingress of chloride ions. Moreover, well-compacted concrete will have
34
closely packed aggregate-cement systems with a minimum amount of interconnected air voids. These closely packed aggregates can cause a wall-effect or act as a physical barrier against the diffusing chloride ions.
Clear (1976) found that inadequate
consolidation of low water-cement ratio, low slump concrete can lead to easier penetration of aggressive elements into concrete.
In general, higher degrees of
compaction and consolidation of concrete result in a slower chloride ingress rates. 3.1.1.5 Effect of initial curing conditions
The relative humidity, water saturation, and temperature can also influence the chloride transport rate in hardened concrete during the first few months after casting. Patel et al. (1985) and Khatib and Mangat (2002) found that the initial curing regime has a large impact on the chloride migration behavior in concrete during early exposure periods. During early exposure periods, the chloride ions can penetrate into surface layers of the concrete. If the concrete surface is dry or improperly cured during this period, the degree of hydration will be lower at the concrete surface layer, resulting in a high degree of porosity near the surface. This increased porosity will lead to a high degree of chloride penetration at or near the exposed surface of the concrete. Because the penetration rate is a function of the chloride concentration at the concrete surface, it is important to minimize chloride exposure at early ages. Khatib and Mangat (2002) also found that the effect of initial curing on the chloride migration property is less for aged (i.e., after few months of casting) concrete. The degree of water saturation and humidity during early exposure periods are also important factors influencing chloride diffusivity. Dehghanian and Arjemandi (1997) found that after 9 and 18 days, water immersion
35
curing of concrete resulted in a lower chloride diffusivity values than concretes exposed to curing at 100% relative humidity. It is not only the initial curing conditions but also the environmental conditions that influence the chloride diffusion rate in concrete. 3.1.1.6 Effect of environmental conditions
The relative humidity, water saturation, and temperature can also influence the chloride diffusion rate in aged (i.e., after first few months) concrete. Climent et al. (2002) reported that a reduction in the percentage of water saturation from approximately 80% to 30% could result in a reduction of the chloride diffusion coefficient by approximately two orders of magnitude. It was also reported that the rate at which the chloride diffusion is decreased is less when the water saturation level in concrete dropped from 80% to 45% compared with water saturation levels between 45% and 30%. Ionic transport in any medium can be influenced by the resistivity of the medium. Moisture content and chemical composition can cause a variation in concrete resistivity from approximately 101 to 105 Ω m (Gj ∅ rv 1977). Also, high degrees of pore saturation and high water-cement ratios can reduce the resistivity of concrete. Climent et al. (2002) reported that the chloride diffusion coefficients in concrete are inversely proportional to the resistivity of the concrete.
In completely and partially saturated concretes the
diffusion coefficient, D , can be empirically calculated using electrical resistivity measurements of the concrete, ρ , as follows (Climent et al. 2002):
D = 6.026 × 10−10 × ρ −0.9997 where ρ ranges from 100 Ωm to 10,000 Ωm , respectively.
(3.2)
36
Goto and Roy (1981) reported that increasing temperatures would result in higher diffusion coefficients in concrete. Higher temperatures increase the thermal vibrations of diffusants (i.e., chloride ions) as explained by Arrhenius' law. Using this, the chloride diffusion coefficient in concrete, D (t , T ) , at time, t , and temperature, T , can be determined as follows (Saetta et al. 1993, Xi and Bazant, 1999; Boddy et al., 1999, and Thoft-Christensen, 2002): m ⎡U ⎛ 1 ⎛ tref ⎞ 1 ⎞⎤ exp . D(t , T ) = Dref ⎜ − ⎢ ⎜ ⎟⎥ ⎟ ⎢⎣ R ⎜⎝ Tref T ⎟⎠ ⎥⎦ ⎝ t ⎠
(3.3)
where Dref is the diffusion coefficient at the reference time, tref , and at the reference temperature, Tref , with an activation energy, U . The R term is the universal gas constant and m is a constant based on the activation energy and is dependent on the water-cement ratio, temperature, and the cement type (Collepardi et al. 1972, and Page et al. 1981). Table 3.1 shows the activation energy values for some cement pastes with different water-cement ratios.
Table 3.1 Activation energies of chloride diffusion process in ordinary portland cement pastes
Water-cement ratio 0.4 0.4 0.5 0.6
Temperature, T , o C ,oT 25, 77 27, 80.6 25, 77 25. 77
Activation Energy, U , (kJ/mol) 41.8 ± 4.0 50.24 44.6 ± 4.3 32.0 ± 2.4
Reference (Page et al. 1981) (Goto and Roy 1981) (Page et al. 1981) (Page et al. 1981)
37
3.1.1.7 Effect of chloride exposure conditions and time of exposure
The chloride diffusivity in concrete varies with the chemical-physical properties of the source of chlorides. Chlorides in cementitious materials exist in the form of free and chemically or physically bound chlorides. Free chlorides are mobile and can diffuse deeper into concrete.
Sodium chloride ( NaCl ), calcium chloride ( CaCl2 ), and
magnesium chloride ( MgCl2 ) are typical de-icing/anti-icing salts. Al-Hussaini et al. (1990) found that more free chlorides were available when the source of chlorides was NaCl than CaCl2 . This may be because of the probability of formation of more calcium
compounds, which bind more chlorides resulting in a lower amount of free chlorides than that in the case of NaCl . Also, the probability of the formation of sodium compounds is less when compared with that of calcium compounds. The researchers also found that the amount of free chlorides increased with increasing amount of NaCl applied. Luping and Nilsson's (1993) report, that the chemical binding of penetrating chlorides will effectively reduce the amount of free chlorides and thus slow the diffusion process and vice versa, could support this finding. Considering the effect of ionic interaction, the effect of retardation on the drift velocities of chlorides and the effect of the electrical double layer on the solid surface and ionic clouds, Zhang and Gj ∅ rv (1996) ranked some typical salt solutions on the basis of their chloride diffusivity in concrete as follows:
DLiCl < DNaCl < DKCl < DCaCl2
(3.4)
38
The above ranking was based on salt solutions that were very dilute when compared to the salt solutions used for testing of chloride diffusivity in concrete. Unfortunately, the research did not provide quantitative information on the concentration of these solutions. Although some parts of marine and bridge structures are submersed in water, there are structural components (i.e., splash zones, and parts of piers near or above water level) that experience wet/dry cycles.
Traetteberg (1977) found that concrete
carbonation, which occurs during the dry period of the wet/dry cycles, reduces the chloride binding in concrete. This effectively results in an increased amount of available free chlorides that can penetrate or diffuse into concrete. Hong and Hooton (1999) reported that the depth of chloride penetration in concrete under wet/dry cycles is related to the square root of number of wet/dry cycles. This implies that the depth of chloride penetration in concrete under wet/dry cycles could be related to the square root of the exposure period, which can be related to the number of wet/dry cycles. Concrete structures experiencing wet/dry cyclic exposure occurs mostly under partially saturated conditions. Climent et al. (2002) suggested another formulation for determining the chloride diffusion coefficient, D (m2/s), in partially saturated, semiinfinite, porous concrete exposed to an instantaneous plane and limited source of chloride ions as follows: ⎛ m ln(C ) = ln ⎜ ⎝ π Dt
⎞ ⎛ x2 ⎞ ⎟ − ⎜ 4 Dt ⎟ ⎠ ⎝ ⎠
(3.5)
where m (kg/m2, lb/ft2 or lb/yd2) is the amount of chlorides applied and C (kg/m3, lb/ft3 or lb/yd3)is the chloride concentration value obtained by chemical analysis
39
of core samples collected from a depth, x (meter, feet or yard), after time of exposure, t (second). This represents the case where de-icing/anti-icing salts are applied during the winter season followed by a summer season when no application is required, but the already applied chlorides are deposited or absorbed into the concrete surface and later diffuse into the concrete. 3.1.2
Mathematical models for diffusion based chloride transport in concrete
The migration process of chloride ions occurs at a slow rate, mainly by diffusion. Fick’s second law (Crank 1975) can be used to model both steady and non-steady state migration of chloride ions in concrete, which is a porous media, as follows:
∂C ∂ 2C =D 2 ∂t ∂x
(3.6)
where D is the diffusion coefficient, C is the chloride concentration, x is the depth of penetration, and t is the time. By changing boundary conditions one can achieve many particular solutions for this differential equation. Theoretical and experimental results obtained in heat transfer studies can be directly applied to diffusion phenomena and vice versa (Frank-Kamenetskii 1955). Carslaw and Jaeger (1947) provided solutions with various boundary conditions for the flow of heat in a semi-infinite solid media obtained using the Laplace transformation method. These solutions can be used for modeling chloride migration in concrete with various boundary conditions. For the following formulations some assumptions are made.
•
Concrete is a semi-infinite, porous, homogeneous, and isotropic material,
40
• • •
No reactions occur between the concrete and the diffusing species (chlorides), C ( x, t ) = 0 when t = 0; 0 < x 24
Reference
Pfeifer 1997 Hurley and Scully 2002
As is evident from Tables 3.2 and 3.3 there exists a large scatter among various chloride threshold levels reported for similar type of steels. Alonso et al. (2000) reported that the scatter observed in the critical chloride threshold values reported in the literature could be attributed to the experimental difficulties and inaccuracies associated with the measurement of actual chloride concentrations or
Cl − , especially at the thin steel(OH ) −
cementitious material interface. Hussain et al. (1995) reported, "... Due to the difficulties for measurement of
Cl − and dependency of corrosion initiation on numerous factors, (OH ) −
there is no single value of threshold
Cl − which is accepted universally". Alonso et al. (OH ) −
(2000) reported that the large number of influencing factors is the reason for the scatter in the critical chloride threshold value for the depassivation of steel reinforcement embedded in cementitious materials. Some of these influencing factors are discussed in the next subsection.
58
3.3.5
Factors influencing the critical chloride threshold value
As stated before, the critical chloride threshold value of steel reinforcement embedded in concrete is influenced by several factors including characteristics of the steel reinforcement, the cementitious material, and the steel-cementitious material interface. 3.3.5.1 Steel characteristics
Steels with different chemical compositions are available on the market. The presence of nickel, chromium, molybdenum and other alloying elements in steel can result in increased resistance against corrosion. The critical chloride threshold value could also be influenced by the mill scale and other protective films present on the steel surface. In addition, the homogeneity or uniformity of films or scales can influence the critical chloride threshold values. On the same reinforcement, the thickness and density of the mill scale can vary from one location to another. This can result in localized corrosion attack. Steels can also have various types of microstructure, including single or multiple phases of ferrite, martensite, austenite, or pearlite. These microstructures have different corrosion resistant properties. Microstructural design of steel bars can be achieved by proper selection of alloying elements and manufacturing procedures (especially cooling procedures). It is not only the cooling procedure but also the cooling rate that influences the microstructure selection (Pan et al. 1998).
The presence of ferritic-martensite
structures with no carbides can increase the corrosion resistance property (Trejo et al. 2000). Maslehuddin et al. (2002) studied the effect of steel manufacturing processes on
59
the corrosion-resistance of air-cooled and water-cooled steel bars. Cooling procedures during steel manufacturing lead to the formation of thin oxide layers (i.e., mill scale) that influenced the corrosion characteristics at the steel surface. Maslehuddin et al. (2002) found that the water-quenched steel bars have higher corrosion resistance than air-cooled steel bars. This property is attributed to the fact that a loosely adherent mill scale layer is formed during the air-cooling of steel bars while a relatively well adherent mill scale layer is formed during water-cooling of steel bars. The surface condition of the steel can have a major influence on the chloride threshold value for corrosion initiation (Mammoliti et al. 1996).
Steels with good
corrosion resistant microstructures may still be highly susceptible to localized and intergranular corrosion because of the possible irregularities on the steel surface. Pan et al. (1998) found that laser surface remelting or hardening can increase the corrosion resistance of structural steel. Pantelis et al. (2002) confirmed that localized corrosion resistance could be increased with laser surface hardening of structural steel. When exposed to the atmosphere and before embedding in concrete a rust film can form on the surface of the steel reinforcement.
Hanson and Sorensen (1990)
compared the corrosion performance of pre-rusted and as-rolled steel bars and concluded that the presence of a rust film on the steel reinforcement has a positive effect on its corrosion resistance when embedded in concrete. The rust film is believed to adhere to the steel substrate and acts as a physical barrier that retards the chloride ingress to the metal substrate.
60
3.3.5.2 Cementitious material and interfacial transition zone characteristics
Apart from the steel characteristics, various characteristics of the cementitious material and the interfacial transition zone (ITZ) influence the critical chloride threshold value of the embedded steel reinforcement. Lewis and Copenhagen (1959) rated the concrete permeability as the most important single factor affecting the corrosion. Diffusivity of concrete does not directly affect the critical chloride threshold value but it certainly influences the time for the chlorides to reach the steel reinforcement embedded in concrete. Both diffusivity and permeability of concrete is greatly influenced by the water-cement ratio, curing regime, mixture proportioning and constituents, and other factors. The chemistry and pH of the interstitial pore solution significantly affects the chloride threshold level of steel reinforcement embedded in cementitious materials. Glass and Buenfeld (1997) reported that the most important inhibitive property of the cement appears to be its effective buffering capacity that restricts any local decrease in the pH. Suryavanshi et al. (1996) reported that chloride binding could also increase the pH due to the release of hydroxyl ions into the interstitial solution during the chloride binding. It is also well documented that the C3 A and C 4 AF content of the cement paste results in the binding of chlorides by the formation of calcium chloroaluminate (Friedel's salt) (Rasheeduzzafar et al. 1990, Rasheeduzzafar 1992). Chloride binding reduces the mobility of chlorides. It should be noted that the chloride binding does not influence the
61
critical chloride threshold value. It only increases the time to attain the threshold chloride concentrations in concrete. Hussain et al. (1995) reported that an increase in C3 A content in cement from 2.43% to 14% could increase the critical chloride threshold value by 2.85 times. The authors also reported, "The alkali content of cement has marginal effect whereas presence of sulfates along with chlorides has moderate effect on the threshold chloride content." The presence of corrosion inhibitors such as calcium nitrite in concrete can increase the stability of the passive layer and thereby increases the critical chloride threshold level of embedded steel in concrete (Berke et al. 1988). Thomas (1996) reported that chloride threshold levels decreased with increasing fly ash contents. 3.4
PROPAGATION OF CHLORIDE-INDUCED CORROSION
In most cases, the chlorides from the atmosphere and other sources penetrate into the concrete. Once corrosion is initiated, the chloride ions, the concentration of which keep increasing, act as a catalyst and the process of corrosion propagates. Some of the several factors that influence the propagation of chloride-induced corrosion are explained below. Liu and Weyers (1998) reported that corrosion rate is a function of the ohmic resistance of the concrete cover, the chloride content, the temperature, and the corrosion time.
The effect of the cement composition (Rasheeduzzafar et al. 1990), mineral
admixtures (Hope and Ip 1987, Al-Amoudi et al. 1993, Andióna et al. 2001, Pal et al., 2002), material composition (Lorentz and French 1995), loading conditions (Yoon et al. 2000), and environmental conditions (Balabanic et al. 1996, Andrade et al. 2002, Pech-
62
Canul and Castro 2002) on the corrosion rate have been investigated and reported throughout the literature. This information can be used as guidance for designers and engineers to estimate the time from corrosion initiation until the time of cracking or spalling. Corrosion will not occur, even on already depassivated steel, if oxygen is not present (Ahmad 2003). The conversion of ferrous chloride into ferrous hydroxide, and other insoluble hydrous iron oxides (rust) occurs only in the presence of both moisture and oxygen (Hope and Ip 1987). Hence, it can be concluded that both moisture and oxygen in sufficient quantity is necessary for the corrosion propagation.
A typical
mechanism of propagation of chloride-induced corrosion in concrete is shown in Figure 3.4. Concrete surface
Corrosion products Passive film
ClFe
O2
H2O
Fe + 2e2+
Anode Fe2++ 2(OH)-
Cl-
O2 2(OH) 1 2 O2 + H2O + 2e 2e-
Fe(OH)2
H2O 2(OH)- Base steel
Cathode Concrete
Figure 3.4 Typical mechanism of propagation of chloride-induced corrosion in concrete (Adapted from Ahmed 2003).
63
Polder and Peelen (2002) reported that the resistivity of the concrete and the corrosion rate of the steel reinforcement after corrosion initiation are related. Lower electrical resistivity values of concrete and steel-concrete interface can facilitate faster corrosion current flow.
As the chloride concentration in the pore water solution
increases, the concrete resistivity decreases, leading to faster corrosion rates (Hope and Ip 1987). The time after corrosion initiation has a significant effect on corrosion rate during early stages of corrosion initiation. The corrosion rate decreases rapidly during the first year after initiation and then attains a constant value. This may be attributed to two facts: reduction in anode/cathode area ratio and reduced flow rate (due to a slower diffusion rate through rust layers) of iron ions away from the steel surface (Liu and Weyers 1998). High temperature and high humidity will increase the electrode reaction rates if
(
the situation is conducive for corrosion (Uhlig and Revie 1985). Z ivica et al. (1997) reported that the ambient temperature above 40 oC can result in reduced corrosion rates. The authors attributed this behavior to the fact that the oxygen solubility in the pore solution is considerably decreased in these conditions. The continuous propagation of corrosion of steel reinforcement in concrete eventually results in serious structural damages. A discussion on corrosion-induced structural damage is provided in the following subsection.
64
3.5
CORROSION-INDUCED CRACKING OR SPALLING OF CONCRETE COVER
The corrosion propagation of reinforcement embedded in concrete leads to the structural failure of RC structures. The two important causes of structural failure are the loss of effective steel cross section and cracking or spalling of the concrete cover. Weyers et al. (1994) concluded that "... it is likely that the end of functional service life for concrete bridge decks is reached when the percentage of the worst traffic lane surface area that is spalled, delaminated, and patched ranges from approximately 9% to 14%." Corrosion produces corrosion products around the steel surface. It has been reported that the volume of these corrosion products are approximately 4 to 6 times larger in volume than the corroding base steel (Berke et al. 1988, Mehta 1986).
If
interconnected voids are present, these corrosion products can diffuse into voids around the steel reinforcement. The total available volume of these voids is dependent on the water-cement ratio, the surface area of reinforcement, the degree of hydration, and the degree of consolidation (Liu 1996). At some point, these corrosion products completely fill these voids and start exerting expansive-tensile stresses on concrete cover causing cracking or spalling (Berke et al. 1988). • • •
Critical parameters that can be used in describing this corrosioninduced cracking mechanism are (Liu 1996): the total amount of corrosion products formed at any point of time, WT , the total amount of corrosion products needed to fill the individual and interconnected and pores and voids in the vicinity of steel reinforcement, WP ,
65
• •
the critical amount of corrosion products or rust required to initiate cracking or spalling of concrete cover, Wcrit , the critical base steel thickness lost or corroded (corresponding to Wcrit ), Tcrit , steel .
This mechanism of corrosion-induced cracking can be divided into four phases. These phases include • • • •
the corrosion initiation phase, the free expansion phase, the stress initiation phase, and the cracking or spalling phase.
In the corrosion initiation phase, the steel corrosion initiates, and the amount of corrosion products is very low and does not generate expansive stresses on the surrounding concrete. The ITZ around the steel surface is typically porous (Belaid et al. 2001) and can easily accommodate the corrosion products formed, provided the condition in Equation 3.19 is satisfied. WT
WP
(3.17)
In the free expansion phase (Figure 3.5), the corrosion products diffuse into voids of the ITZ around steel surface, and individual or interconnected voids in the concrete. At this stage, the total amount of corrosion products, WT , is less than or equal to the amount of uncompressed corrosion products required to fill the pores and voids in the vicinity of steel reinforcement, WP , as follows: WT ≤ WP
(3.18)
Hence, there are no significant stresses exerted on the surrounding concrete to cause cracking or spalling.
66
"Uncompressed" corrosion products fills the interfacial pores Base steel Surrounding concrete
Figure 3.5 The free expansion phase.
In the free expansion phase, the interfacial zone and other voids in the concrete become fully occupied with uncompressed corrosion products.
As the corrosion
propagation continues, the cumulative volume of corrosion products increase. These corrosion products become compressed and begin exerting expansive-tensile stresses on the surrounding concrete (Berke et al. 1988). Wcrit > WT > WP
(3.19)
A schematic of stress initiation phase is given in Figure 3.6.
"Slightly compressed" corrosion products in the interfacial pores Base steel Surrounding concrete
Figure 3.6 The stress initiation phase.
67
The cracking or spalling phase follows the stress initiation phase. At some point in time, the expansive stresses, generated by excessive formation of corrosion products, cross the tensile strength capacity of the concrete. At this point, WT > Wcrit
(3.20)
This causes the concrete to crack or spall, an indication of loss of structural integrity. Figure 3.7 shows the schematic of cracking or spalling phase.
"Highly compressed" corrosion products causing cracks in concrete Base steel Radial cracks Surrounding concrete
Figure 3.7 The cracking or spalling phase.
Also, as shown in Figure 3.7, the cross section of the steel reinforcement is reduced, reducing the tensile capacity of the structural element. 3.5.1
Critical amount of corrosion products resulting in cracking and spalling
The critical amount of corrosion products required to cause cracking, Wcrit , can be obtained as follows (Liu 1996): ⎛ W ⎞ Wcrit = ρ rust ⎜ π [ d s + d o ] D + st ⎟ ρ st ⎠ ⎝
(3.21)
68
where
ρ rust = density of corrosion products ρ st = density of steel Wst = mass of steel corroded D = diameter of steel reinforcement d o = thickness of porous zone around the steel/concrete interface; d o
D,
d s = radial displacement of concrete under pressure, which is assumed to be equal to the thickness of corrosion products to generate the critical tensile stresses; d s
D
Assuming that the concrete is a homogeneous elastic material and a crack occurs just over the reinforcement or at the steel-concrete interface, the term d s can be obtained as follows.
Expansive pressures develop on the surrounding concrete due to the
formation of corrosion products as shown in Figure 3.8.
P do ds D
P P
P
P P
P
+
P
d
P Figure 3.8 Expansive pressure on surrounding concrete due to formation of rusty products (Liu 1996).
69
The algebraic sum of opposite forces acting at the concrete-corrosion product interface gives the formulation for determining d s : ds =
⎞ d ⋅ ft ′⎛ a 2 + b 2 + υc ⎟ ⎜ 2 2 Eef ⎝ b − a ⎠
(3.22)
where d is cover depth, ft ′ is tensile strength of concrete, υc is Poisson's ratio of the concrete, Eef is effective elastic modulus of concrete, a and b are inner and outer radii of the thick-walled concrete, respectively.
The terms a and b can be obtained as
follows:
( D + 2d o ) 2
(3.23)
⎛ D + 2d o ⎞ b = d +⎜ ⎟ 2 ⎠ ⎝
(3.24)
a=
Bazant (1979b) reported that factors such as bar size, bar spacing, and concrete cover determine the type of failures such as cracking, spalling, and delamination (see Figure 3.9).
70
cover depth
(a) Cracking
cover depth
(b) Spalling
cover depth
(c) Delamination Figure 3.9 The cracking, spalling and delamination of concrete cover.
3.5.2
Cracking and spalling threshold thickness
The cracking or spalling threshold thickness can be defined as the minimum amount of steel corrosion required to trigger the cracking or spalling of the concrete cover. This parameter can be expressed in terms of the critical bar thickness loss, Tcrit , steel , corresponding to the amount of corrosion products required to cause cracking and spalling, Wcrit . The term Tcrit , steel is a function of the interconnectivity of the concrete pores, the tensile strength capacity of concrete, the concrete cover depth, the surface area and shape
71
of the reinforcement, and other parameters (Clear 1992, Andrade et al. 1993). The watercement ratio along with degree of consolidation influences both the interconnectivity of concrete pores (Soroka 1979, Sugiyama et al. 1996, Pfeifer 1997) and the tensile strength capacity of concrete (Neville 1998). The material quality of the concrete can also be influenced by proper mixture proportioning, degree of consolidation, curing regime, and other factors. Rodriguez et al. (1994) reported that crack formation is a function of both the reduction of bar radius and concrete cover to bar diameter ratio. Table 3.4 clearly shows that a large scatter (ranging between 0.1 and 29 mils) exist among the critical cracking and spalling threshold thickness reported in the literature.
Table 3.4 Critical cracking and spalling threshold thickness values, Tcrit , steel obtained
from the literature Tcrit , steel
Reference
(mils)
Other test details
Spellman and Stratfull (1968) Pfeifer 1997 (from Iding)
concrete tensile stress = 1000 psi; creep factor = 3
0.3
Bazant (1979b)
Interpretation of Equations
0.59
Pfeifer 1997 (from Iding)
0.59-1.5
Rodriguez et al. (1994)
0.63-1.26
Hladky et al. (1989)
0.98
Pfeifer (1997)
3.7
Clear (1992) Spellman and Stratfull (1968)
concrete tensile stress = 1000 psi (6.89 MPa); no creep concrete tensile stress = 537psi (3.70 MPa); at surface crack appearance; 16mm diameter bar Interpretation of results at surface crack appearance ; 16mm diameter bar Interpretation of results
0.1 0.2
up to 29
1mil = 0.001 inch.
Laboratory experiment
Field experiment
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Pfeifer (1997) reported that a 96-week in-concrete test with conventional carbon steels and stainless steels (Type 304 and 316) resulted, on average, in corrosion rates of 0.036 mm/year (1.4 mils/year) and 0.000051 mm/year (0.002 mils/year), respectively. The Tcrit , steel for a concrete with same material and strength properties can increase as the cover depth increases. This is because of the fact that the effective tensile strength of the concrete cover increases with increase in cover depth. Insufficient literature is available on the effect of water-cement ratio, which influences the porosity of concrete, on the Tcrit , steel . Quantitative and reliable information on the Tcrit , steel for a concrete cover along with that on the corrosion rate can be used to determine the duration of the corrosion propagation phase. In this section, various mechanisms associated with the transport of chlorides into concrete, the initiation and propagation of the chloride-induced corrosion in concrete, and the corrosion-induced cracking or spalling of concrete cover were discussed. The two critical service life parameters discussed were the critical chloride threshold value of the steel reinforcement and the critical cracking or spalling threshold thickness of the concrete cover. Quantitative and reliable information on these and other parameters can be useful to predict the duration of initiation and propagation phases of chloride induced corrosion in concrete. Knowledge on the duration of the initiation and propagation phases can be used to predict the overall service life and life cycle cost of RC structures exposed to chloride environments. The service life prediction and life cycle cost analysis methods are presented in next section.
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4
SERVICE LIFE AND LIFE CYCLE COST OF RC STRUCTURES EXPOSED TO CHLORIDE ENVIRONMENTS In the previous section a review of various mechanisms and critical parameters
related to the chloride-induced corrosion in concrete was presented. This section will discuss various phases of the overall service life of RC structures exposed to chloride environments. Also presented is a brief review of mathematical models used to predict the service life and life cycle cost of such structures. Presently, new durable construction materials are becoming available. But, these materials typically have higher initial costs, often preventing their use. But these materials can be cost effective on a long-term basis considering their low repair cost when compared to that of conventional materials. A procedure for selecting construction materials on the basis of long-term cost effectiveness is also provided. 4.1 4.1.1
SERVICE LIFE OF RC STRUCTURES Definitions and influencing factors
The service life of RC structures can be defined as the time during which a structure is able to meet the user requirements with an acceptable level of safety. Steel reinforcement corrosion in RC structures is a time-dependent deterioration process. This deterioration is mainly due to the corrosive environmental conditions that changes the material characteristics and in turn affects the serviceability and load carrying capacity of the structure. As the serviceability or load carrying capacity of a structure is reduced, at
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some point in time the structure is no longer serviceable and safe and has to be either repaired or decommissioned. As discussed in Section 3, there are several internal and external factors that influence the service life RC structures. Figure 4.1 shows a simple schematic of some of the factors that can influence the service life of RC structures exposed to chloride environments.
Cracking and spalling threshold thickness
Corrosion rate
Other design parameters (Water-cement ratio, cover depth, etc.)
Critical chloride threshold value
Service life
Chloride diffusion coefficient
Other characteristics of steel, concrete and environment
Figure 4.1 Some important factors influencing the service life or RC structures exposed to chloride environments.
Quantitative and reliable information on these dynamic factors can be helpful in predicting the service life and performing the life cycle cost analyses for the RC structures.
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4.1.2
Various time phases and prediction of service life
The overall service life of a RC structures exposed to chloride environments can be classified with three phases as follows: • • •
The chloride-induced corrosion initiation phase, The chloride-induced corrosion propagation phase, and The repair and rehabilitation phase.
These phases, measured in time, comprise of the overall service life as follows: Overall Service Life = ( Initiation Phase + Propagation Phase + Repair Phase)
(4.1)
Figure 4.2 shows a simple schematic of the overall service life in terms of different phases of reinforcement corrosion.
Damage Level
Maximum Allowable Damage Level
Last Repair Intermediate Repairs First Repair Time Initiation Phase
Propagation Phase
Repair Phase
Overall Service Life
Figure 4.2 Different phases of overall service life in terms of reinforcement corrosion in RC structures (adapted after Trejo 2002).
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The crack width, the crack width to cover depth ratio, the area of the cracked, spalled, or delaminated concrete, the amount of steel reinforcement corrosion, and the reduction in the cross sectional area of the steel reinforcement are some parameters that can assist engineers and technicians in identifying damage levels for corroding RC structures. Also in Figure 4.2, the maximum allowable damage level indicates the critical limits for these parameters. In Figure 4.2, the break in the horizontal line over the time axis indicates that the initiation phase could be much longer that the propagation and repair phases. More detailed explanations on these phases are provided in the following subsections. 4.1.3
The chloride-induced corrosion initiation phase
When RC structures are exposed to chloride environments the chloride ions migrate slowly towards the steel reinforcement and eventually initiate corrosion of the steel reinforcement. The corrosion initiation phase is composed of the time when the structure is placed into service until the time when corrosion of the reinforcement begins. The initiation phase is typically much longer than the propagation and repair phases. The key parameters to determine the time of the corrosion initiation phase, ti , are the critical chloride threshold level of the embedded steel reinforcement, Cthreshold , the chloride diffusion coefficient of the concrete cover, D , the clear cover depth, d , and the chloride build-up rate at the concrete surface, Φ (t ) . As addressed in Section 3, chloride concentrations near the steel reinforcement gradually increases as chloride ions from the
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surrounding environment (i.e., sea water, de-icing/anti-icing salt, etc.) penetrate into the concrete. Assuming a constant surface chloride concentration, Φ (t ) = C0 , for a RC structure, substituting x , and C ( x, t ) in Equation 3.7 with cover depth, d , and Cthreshold of the embedded steel reinforcement, respectively, gives: ⎛ Cthreshold = Ci + (C0 − Ci ) ⎜1 − erf ⎜ ⎝
⎛ d ⎜ ⎜ 4 Dt i ⎝
⎞⎞ ⎟⎟ ⎟⎟ ⎠⎠
(4.2)
Equation 4.2 can be directly used to determine ti provided the terms Cthreshold , Ci , C0 , D , and d , as defined earlier, are known (Collepardi et al. 1972, Clifton 1993, Maage et al. 1996, Prezzi et al. 1996, and Amey et al. 1998). Equation 4.2 can also be used for determining the ti of bridge structures where the de-icing/anti-icing salts are applied only when average constant value for chloride concentrations over of time at the concrete surface can be assumed. In general, Equation 4.2 can be used for determining the corrosion initiation time for any RC structure, provided the chloride concentration in the surrounding environment is assumed to be constant over the specified time period. For the case where the surface concentration increases in a linear fashion, substituting x and C ( x, t ) in Equation 3.10 with cover depth, d , and Cthreshold of the embedded steel reinforcement gives:
Cthreshold
⎧⎪⎛ d2 = kti ⎨⎜1 + ⎪⎩⎝ 2 Da ti
⎞⎛ ⎟ ⎜⎜ 1 − erf ⎠⎝
⎛ d ⎜ ⎜ 4D t a i ⎝
⎞⎞ ⎛ d ⎟⎟ − ⎜ ⎟⎟ ⎜ π D t a i ⎠⎠ ⎝
⎞ ⎛ −d 2 ⎞ ⎫⎪ ⎟ exp ⎜ ⎟⎬ ⎟ ⎝ 4 Da ti ⎠ ⎪⎭ ⎠
(4.3)
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Equation 4.3 can be used to determine ti by iterative or trial and error methods, provided the Cthreshold , k , Da , and d are known (Amey et al. 1998). The value of Da can be taken as the mean of chloride diffusion coefficients from several representative samples collected from the actual RC structure. The apparent diffusion coefficient is a more "practically obtainable" parameter than the exact diffusion coefficient. In general, Equation 4.3 can be used for any RC structure where the chloride concentration in the surrounding environment is assumed to be linearly increasing over the specified exposure period. For the cases where the surface chloride concentration varies as a function of square root of time (i.e., Φ (t ) = k t ), Equation 3.11 is used. Substituting x and C ( x, t ) in Equation 3.11 with cover depth, d , and Cthreshold of the embedded steel reinforcement, respectively, we get: ⎛ −d 2 ⎞ ⎛ d π ⎪⎧ Cthreshold = k ti ⎨exp ⎜ ⎟ − ⎜⎜ D t 4 a i ⎝ ⎠ ⎝ 4 Da ti ⎪⎩
⎞ ⎛ ⎟ ⋅ ⎜1 − erf ⎟ ⎜ ⎠ ⎝
⎛ d ⎜ ⎜ 4D t a i ⎝
⎞ ⎞ ⎪⎫ ⎟ ⎟⎬ ⎟⎟ ⎠ ⎠ ⎪⎭
(4.4)
Equation 4.4 can be used to determine the time to corrosion initiation by iterative or trial and error methods, provided the Cthreshold , k , Da , and d are known (Amey et al. 1998). For de-icing/anti-icing salt applications, a linear or square root buildup rate for chloride concentration at the concrete surface is more reasonable than assuming a constant value. Therefore, Equations 4.3 and 4.4 are most suited for evaluating air-born and direct de-icing/anti-icing salt applications on bridge decks (Amey et al. 1998, and
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ACI 365.1R-00 2000) until the surface chloride concentration that is equal to kt or k t reaches saturated chloride concentration in concrete, Csat .
If the surface chloride
concentration is equal to or more than Csat , Equation 4.2 could be used by substituting C0 with Csat and Ci with C ( x, t ) at the time of saturation, tsat . Therefore a combination of Equation 4.3 (or 4.4) and Equation 4.2 could be used for determining time of corrosion initiation of RC structures (e.g., bridge decks) exposed to de-icing/anti-icing salts. 4.1.4
The chloride-induced corrosion propagation phase
Once corrosion of the steel reinforcement initiates, the corrosion process continues, provided corrosive conditions prevail. The corrosion products exert expansive stresses on the surrounding concrete. At some point in time, the expansive stresses cause cracking and spalling of concrete cover.
The time from corrosion initiation of the
reinforcement to the time of first repair or the time at which the damage level reaches some allowable level is defined as the corrosion propagation phase. The duration of this phase is typically shorter than the initiation phase. Cady and Weyers (1984a) reported that the propagation time ranges between 2 to 5 years. Browne (1982) reported a larger range, 6 months to 5 years, for the time of propagation phase. The duration of the corrosion propagation phase, t p , can be assumed to be equal to the time to cracking or spalling, which is assumed to be equal to time to cracking, tcrack . The corrosion rate, the bar spacing, the bar diameter, the cover depth, the density of steel, the density of corrosion products, tensile strength of the concrete, and the interconnectivity of pores in the concrete are important factors influencing tcrack (Bazant
80
1979a, Bazant 1979b, Cady and Weyers 1983, and Liu 1996). The water-cement ratio influences strength parameters and the interconnectivity of the pores in the concrete (Neville 1998). Many researchers including Liu (1996), and Morinaga (1989) have developed mathematical models to determine tcrack for concrete cover. The mathematical model for determining tcrack developed by Liu (1996) is presented below. The rate of production of corrosion products, dWcp dt
=
dWcp dt
, can be written as follows:
kp Wcp
(4.5)
where Wcp is the amount of corrosion products produced, t is time, and k p is a function related to the rate of metal loss. The term k p can be expressed in terms of corrosion rate, icorr (in mA / ft 2 ) , as follows: k p = 2.59 × 10−6 (1/ α ) π D icorr
(4.6)
where D is steel diameter (inches) and α is ratio of molecular weight of steel to that of the corrosion products. If SI system of units is used, the Equation 4.7 can be rewritten as follows: k p = 7.08 × 10−10 (1/ α ) π D icorr
(4.7)
The integration of Equation (4.6) over time, t , gives: t
(Wcp ) = 2∫ k p dt 2
0
(4.8)
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As Wcp reaches Wcrit , which can be obtained as discussed in Section 3, t reaches tcrack . [Wcp ] lim = Wcrit
(4.9)
t →tcrack
Hence, the solution of Equation 4.9 gives the tcrack as follows: tcrack
(Wcrit ) 2 = 2k p
(4.10)
Morinaga (1989) provided another model for the determination of the time to cracking (or propagation time). The author reported that tcrack could be determined using the critical mass of corrosion products to cause cracking, Qcrack , and the corrosion rate, icorr , as follows: tcrack (days ) =
Qcrack ( grams ) icorr ( grams / day )
(4.11)
where Qcrack can be empirically determined as follows:
⎛ 2d ⎞ Qcrack = 0.602 D ⎜ 1 + ⎟ D⎠ ⎝
0.85
⎛ g ⎞ × 10−4 ⎜ 2 ⎟ ⎝ cm ⎠
(4.12)
where D is the diameter of the reinforcement ( mm ), and d is the cover depth ( mm ). Trejo and Pillai (2003a) reported that the time to cracking (or propagation time), tcrack , could be determined using the critical cracking or spalling threshold thickness, Tcrit , steel , defined in Section 3, and the corrosion rate as follows: tcrack ( yrs ) =
Tcrit , steel (mils ) corrosion rate (mils / yr )
(4.13)
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The time to first repair can then be determined by summing the initiation time and the propagation time as follows: Time to first repair = (tinitiation + t propagation ) = (ti + tcrack )
(4.14)
Cady and Weyers (1984b) reported that the service life to rehabilitation, which is the sum of initiation and propagation times, ranged from 16 to 39 years with a mean value of 29 years for bridge decks in the states of New York, Pennsylvania, and Virginia, reinforced with conventional carbon steel. 4.1.5
The repair and rehabilitation phase
Repairing a deteriorated RC structure can extend the overall service life. The repair phase can be defined as the time from first repair to the time of decommissioning. Multiple repairs may be necessary to maintain structural safety to an acceptable level. But, continued repairs become uneconomical and at some point it becomes more economical to decommission and replace the structure. The life and frequency of repairs depends on the repair methodology, the repair materials, and the structural component or material being repaired. Sprinkel et al. (1991) provided recommendations for repair strategies for RC structures and provided guidance on estimated life expectancy of repairs. In the majority of the cases, the time of the repair phase, tr , is typically shorter than the time of initiation phase, ti . Table 4.1 shows the expected life of various repair methods employed for concrete bridges that can be used for appropriate prediction of the time of the repair and rehabilitation phase.
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Table 4.1 Life expectancy for various repair and rehabilitation methods (Sprinkel et al. 1991 and Koch et al. 2001)
Life Expectancy (years)
Repair method Bituminous concrete patch Portland cement concrete patch Bituminous concrete with membrane Polymer overlay/sealer Portland cement concrete overlay (includes latex modified concrete) Impressed current (cathodic protection) Electrochemical removal of chlorides
4.1.6
Range 1-3 4-10 4.5-15 6-25
Average 1 7 10 10
14-23
18.5
15-35 10-20
35 15
Methodology for predicting service life of RC structures exposed to chloride environments
Information on the critical chloride threshold values of steel reinforcement, cover depths, chloride diffusion coefficients, and chloride exposure conditions can be used to determine the time to corrosion initiation.
The information on critical cracking or
spalling threshold thickness of the concrete cover and corrosion rate of the steel reinforcement can be used to determine the time of corrosion propagation.
The
information on life expectancy of repair and rehabilitation methods can be used to determine time of repair and rehabilitation phase. Before presenting the methods for the life cycle costing, a flow chart for predicting the overall service life of an RC structure is provided in Figure 4.3.
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Identify the RC element and construction materials Determine design and material parameters (cover depth, chloride diffusion coefficient, water-cement ratio, etc.) Determine critical chloride threshold value of steel reinforcement Identify and evaluate chloride environment (seawater, de-icing/anti-icing salts, etc.) Determine surface chloride concentration and build up rate
Determine time to corrosion initiation
Determine corrosion rate during the corrosion propagation phase
Determine critical cracking or spalling threshold thickness
Determine time of corrosion propagation phase Identify possible repair and rehabilitation methods and determine life expectancy Determine total time of repair and rehabilitation phase
Determine the overall service life Figure 4.3 A flow chart for predicting the overall service life of RC structures exposed to chloride environments.
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4.2
LIFE CYCLE COST OF RC STRUCTURES
This subsection presents a simple mathematical model to determine the life cycle cost (LCC) of RC structures exposed to chloride environments. It should be noted that this subsection addresses only those financial issues that are related to chloride-induced corrosion of steel reinforcement in RC structures. 4.2.1
Definition and factors contributing to the life-cycle cost
The LCC of a structure can be defined as the overall cost from concept to decommissioning.
These costs include the initial design costs, construction costs,
operating and maintenance costs, repair costs, user costs, and disposal costs associated with the structure. However, in this thesis, as a means of simplification, only initial construction costs and predicted future repair costs of RC structures will be considered for the LCC analysis and comparative studies on the cost effectiveness of using various construction materials. The initial design costs, operating and maintenance costs, user costs, disposal costs and other costs will not be included in the present study. Figure 4.4 shows a schematic of some important factors contributing to the LCC.
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Repair and rehabilitation costs
Initial design cost
Initial construction cost
Operation and maintenance costs
Life cycle cost
User costs
Other direct and indirect costs
Figure 4.4 Some important factors contributing to the life cycle cost of RC structures.
4.2.2
Life cycle cost analysis
The information on the time duration of various phases of service life and the overall service life (Figure 4.2) can be used to determine the LCC of RC structures exposed chloride environments. Apart from this estimated service life parameters or time data, the cost data on materials and activities related to construction and repairs, the discount rate, and other economic data can be used to determine the LCC of RC structures. The LCC of a RC structure can be determined as the present cost of the sum of all the costs associated with the structure during its overall service life, defined as the time period from the conceptual design to decommissioning. Since the cash flow of various costs related to the structure can be complex in nature, a simplified model is preferred for practical purposes. For a simplified model, the following assumptions can be made:
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The initial cost is equal to the sum of the costs of concrete and steel reinforcement. • • •
All the individual repair costs are the same. The duration of repair cycles is equal for the same type of construction material. The operation and maintenance costs are zero.
The construction costs for different construction methods and materials can be obtained from manufacturers, contractors, or other sources.
Table 4.2 shows the
construction costs of various construction materials (Darwin et al. 2002 and Trejo 2003).
Table 4.2 Construction costs of various construction materials (Darwin et al. 2002 and Trejo 2003)
Construction material
In-place cost
Cost $/m2 ($/yd2)
Concrete
$350/yd3
$98.33/m2 ($82.6/yd2)
ASTM A706 steel reinforcement
$0.58/lb
$34.09/m2 ($28.64/yd2)
ASTM A615 steel reinforcement
$0.57/lb
$33.35/m2 ($28.02/yd2)
MMFX microcomposite steel reinforcement
$0.86/lb
$50.02/m2 ($42.02/yd2)
SS304 stainless steel reinforcement
$1.00/lb
$87.77/m2 ($73.73/yd2)
SS316LN stainless steel reinforcement
$2.00/lb
$117.02/m2 ($98.30/yd2)
According to Darwin et al. (2002), for a typical bridge, the repair costs can be taken as the sum of costs for a deck overlay, reconstruction of bridge rails, and of approach guard rail, the mobilization costs for the equipment and materials, and the
88
traffic control and miscellaneous costs. To simplify the analysis procedure, the repair costs for the bridge rail and the approach guard rail are avoided in this present study. Table 4.3 shows the costs of various repair and rehabilitation methods and works. This data along with the information on the repair cycles can be used in calculating the cost of repair and rehabilitation.
Table 4.3 Costs related to various repair and rehabilitation methods or works (Sprinkel et al. 1991, Koch et al. 2001, and Darwin et al. 2002)
Repair methods and related works Bituminous concrete patch Portland cement concrete patch Bituminous concrete with membrane Polymer overlay/sealer Portland cement concrete overlay (includes latex modified concrete) Low slump dense concrete overlay Impressed current (cathodic protection) Electrochemical removal of chlorides Bridge rail modification Approach guard rail modification Equipment and materials mobilization Traffic control and miscellaneous
Cost $/m2 ($/yd2) Range Average 39-141 (33-118) 90 (76) 322-469 (270-394) 395 (332) 30-86 (25-72) 58 (49) 14-182 (12-153) 98 (82) 151-187 (127-157)
170 (143)
-92-137 (77-115) 53-129 (45-108) -----
67 (80) 114 (136) 91 (108) 19 (23) 23 (27) 26 (31) 13 (15)
To calculate the present cost of the sum of all the costs, the cash flow and the discount rate must be known. In a typical cash flow diagram, as shown in Figure 4.5, only the construction costs and repair costs are considered, the operating and maintenance costs are assumed to be zero, the repair costs increases as more and more
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repair work is performed, and the repair frequency increases as more and more repair works are performed.
2
Cost ($/m )
Repair times
Design and Construction time
Time
Time Initial costs Repair (Future) costs Present Cost, PC
Figure 4.5 A typical cash flow diagram.
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Using basic engineering economic principles the present cost, PC , of a future cost, F , can be determined as follows (Fleischer 1994): PC =
F (1 + i )t
(4.15)
where i is the discount rate and t is the time. In the case of multiple future costs, which can be incurred at different times, the present cost of all such future costs can be calculated using the following formula: N
Fn tn n =1 (1 + i )
PC = ∑
(4.16)
where N is the total number of individual future costs, n is the count, and t is the time when nth cost is incurred. While performing LCC analysis of RC structures, the present worth of all the predicted or expected future costs including repair costs, are calculated using Equation 4.16 or Equation 4.17. Figure 4.6 shows a flowchart for determining the life cycle cost of RC structures exposed to chloride environments.
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Identify the construction materials (concrete, steel reinforcement type, etc.)
Determine the service life parameters (corrosion initiation time, corrosion propagation time etc.)
Determine the construction costs
Identify the repair strategies and evaluate repair costs
Determine the related costs Any other related costs?
Yes
No Determine the cash flow and discount rate
Determine the present worth of sum of all the costs
Life-cycle cost = the present cost of sum of all the costs Figure 4.6 A flow chart for determining the life cycle cost (LCC) of RC structures exposed to chloride environments.
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A simple flow chart for selecting durable and cost effective construction materials for RC structures susceptible to chloride-induced corrosion is shown in Figure 4.7.
Select the RC element
Identify the construction materials available (concrete types, steel reinforcement types, etc.)
Select material combinations and alternatives
Determine the service life parameters (corrosion initiation time, corrosion propagation time etc.)
Determine life cycle cost
Any other material combinations or alternatives?
Yes
No Compare service life period and life cycle costs of the structure with different material combinations or alternatives
Choose the most durable and cost effective material combination or alternative
Figure 4.7 A simple flowchart for selecting the most durable and cost effective material combination or alternative for RC structures.
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The analysis of service life and LCC with various construction materials can be performed at the design stage of the RC structures. LCC analyses provide a more rational basis for the selection of bridge types (or materials for bridges) that will minimize LCC (Freyermuth 2001). Freyermuth (2001) also reported that a minimum service life of 100 years could be achieved with a marginal increase in the initial cost. The author suggested a minimum design life of 150 years for major bridges in urban environments. This section discussed mathematical formulations for service life prediction and life cycle cost analysis of RC structures susceptible to chloride-induced corrosion. A section on current test methodologies to evaluate various material characteristics, which can be used for predicting service life of RC structures susceptible to chloride-induced corrosion follows.
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5
CURRENT TEST METHODS TO PREDICT SERVICE-LIFE OF RC STRUCTURES EXPOSED TO CHLORIDE ENVIRONMENTS Electrical, electrochemical, chemical and other principles have been widely used
to study chloride-induced corrosion and related phenomena in concrete. This section presents current test methodologies used to evaluate critical chloride characteristics used for the prediction of service life of RC structures exposed to chloride environments. 5.1
ACCELERATED METHODS FOR CHLORIDE PENETRATION
Under field conditions, it can take years or decades for chloride ions to penetrate good quality concrete cover, reach the reinforcement, attain the critical chloride threshold concentration, and cause initiation and propagation of corrosion of the reinforcement. This long duration makes it uneconomical to perform non-accelerated corrosion tests in the laboratory. To reduce the duration of laboratory tests for chloride-induced corrosion in concrete samples, accelerated chloride transport techniques are necessary. Researchers have developed and implemented test procedures that accelerate the rate of chloride transport in concrete, thereby reducing the test period. A brief discussion on some accelerated chloride transport methods and the issues associated with each follows. 5.1.1
Chloride penetration by cyclic wet-dry exposure
Structures, especially marine structures in the splash zone, are exposed to cyclic wet-dry conditions. Such real life situations can be simulated in laboratory by cyclicponding with chloride solution and drying of concrete samples. Hong and Hooton (1999) reported that the chloride ingress rate is strongly influenced by the sequence and duration
95
of wetting and drying and that there exists a good relationship between the depth of chloride penetration and the square root of the number of cycles. The researchers also observed that a 3-day cycle resulted in a greater chloride concentration at concrete depths when compared to a corresponding 1-day cycle. It was concluded that longer drying times in a wet-dry cycle facilitates faster chloride ingress and chloride accumulation at different concrete depths. AASHTO T259 (1980), Standard Method of Test for Resistance of Concrete to Chloride Ion Penetration, is often referred to as the salt ponding test because it involves
cyclic ponding with salt solution and drying of concrete specimens. This test is a longterm test method that provides information on one-dimensional chloride ingress profiles in normal concrete over a testing duration of 90 days. Stanish et al. (1997) reported that there have been difficulties in developing a sufficient chloride profile in higher quality concretes over a duration of 90-days for this test. In general, reinforcement with very high critical chloride threshold levels embedded in high quality concrete could take many years or decades to initiate active corrosion. Hence, a faster chloride penetration technique is needed. 5.1.2
Electrically accelerated chloride penetration
The use of applying potential gradients (or voltages) to accelerate the chloride penetration in concrete was first investigated by Whiting (1981). Several researchers including Dhir et al. (1990), Andrade (1993), Jacobsen et al. (1996), and McGrath and Hooton (1996) have used potential gradients to accelerate the penetration of chlorides
96
into concrete under laboratory test settings. These test methods were developed mainly to determine the chloride diffusion coefficient of concrete. It has been established that the application of a potential gradient across the anode and the cathode (or across the concrete specimen) can accelerate the chloride ion penetration into the concrete specimen. The rate of chloride penetration and the test duration depends largely upon the voltage applied (McGrath and Hooton 1996), the distance between the anode and cathode, the material properties of the specimen, and other factors (Stanish et al. 1997). The major problems associated with the application of potential gradients while testing corrosion characteristics of embedded steel specimens will be discussed in the following paragraphs. Castellote et al. (2002) reported that the application of a potential gradient could induce electrical polarization (i.e., a forced shift in potential) of the steel specimen. Although it was found that the steel specimen recovers the initial value of potential once the electrical field is switched off, the induced polarization can change the surface characteristics of the steel specimen resulting in erroneous or unreliable corrosion test results. The lower the amount of this induced potential shift, the higher will be the accuracy of the test results. Another potential challenge associated with electrically induced migration of chlorides in concrete is the unwanted migration of the hydroxyl and other negatively charged ions in the same direction as that of the negatively charged chloride ions. The mobility of hydroxyl ions is "nearly twice" that of chloride ions (Prince et al. 1999). The authors also reported that this hydroxyl ion migration results in an unwanted increase in
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the pH at the anode region, which in turn results in a deviation from the actual chemical conditions in the field concrete exposed to chlorides and complicates the test. A third problem is the generation of heat due to the application of voltage or potential gradient. El-Belbol and Buenfeld (1989) conducted a study on heat generation at various voltages as shown in Table 5.1.
Table 5.1 Heat generation due to applied voltage (after El-Belbol and Buenfeld 1989)
Voltage applied for 4 days 60 40
Increase in temperature of test samples 18% Negligible
A voltage low enough to avoid the adverse problems (i.e., forced polarization of the test sample, forced pH variations near the test sample, forced heating of the test sample, and other problems), yet high enough to ensure a reasonably short test duration needs to be determined. 5.2
CORROSION RATE MEASUREMENT BY MASS LOSS TESTS
There are standard test methods available for the determination of corrosion rates by gravimetric mass loss measurements. Two of these methods are ASTM G1-99ε1 (1999), Standard Practice for Preparing, Cleaning, and Evaluating Corrosion Test Specimens, and NACE Standard TM019 (2000), Standard Test Method for Laboratory Corrosion Testing of Metals.
98
In the case of uniform or general corrosion, these mass loss test methods provide reliable information on the average corrosion rates over a period of time. In the case of localized corrosion (pitting corrosion, crevice corrosion, etc.) these mass loss methods are not as reliable in determining corrosion rates, especially if expressed in terms of average thickness of metal corroded per unit time. This is because for localized corrosion the actual thickness of metal corroded is highly uneven across various locations on the metal surface because of pits and crevices and the mass loss methods usually results in average corrosion rates and average depths of corrosion obtained from the total mass loss over the entire area of exposure of the test coupon or specimen. These mass loss methods do not give information on the actual corrosion rates or depths of corrosion at local or specific area (e.g., pits and crevices) attacked by localized corrosion. Moreover, the destructive nature of these tests makes them unsuitable for the on-site corrosion rate measurement on corroding RC structures. 5.3
ELECTROCHEMICAL METHODS FOR CORROSION MONITORING
Visual observations can be made to monitor the initiation and the propagation of corrosion of steel immersed in simulated concrete pore solutions. Visual observations cannot be made when the steel is embedded in concrete due to the concrete cover. Electrochemical principles can be used, remotely and non-destructively, to monitor the initiation and the propagation of corrosion of steel reinforcement embedded in concrete.
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5.3.1
Half-cell potential measurements
The half-cell potential forms the basis for predicting the tendency of a corrosion reaction to occur (Fontana and Greene 1978). The tendency of any reaction to occur increases as the half-cell potential of the reaction decreases. This property can be used in monitoring corrosion of steel in concrete. Therefore, the measurement of the half-cell potential of the steel reinforcement is one of the easiest and more typical procedures used for the routine corrosion inspections of RC structures. Half-cell potential measurement can be easily made if the following are available: • • •
a suitable reference electrode, a voltmeter (or potentiometer), and a conductive bridge or region between the reference electrode and the steel sample.
ASTM C876-91 (1999), Standard Test Method for Half-Cell Potentials of Uncoated Reinforcing Steel in Concrete, describes the methodology for using half-cell
potential measurements to detect the probable areas of corrosion on reinforcement. ASTM G109-99ε1 (1999), Standard Test Method for Determining the Effects of Chemical Admixtures on the Corrosion of Embedded Steel Reinforcement in Concrete Exposed to Chloride Environments uses the half-cell potential measurements to detect the
probable areas of corrosion on reinforcement. Table 5.2 shows the corrosion detection method used by ASTM C876-91 (1999).
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Table 5.2 Half-cell potential Vs probability of corrosion occurrence (ASTM C876-91 1999)
Half-cell potential mV Vs CSEa mV Vs SHEb < -350 mV -200 mV > -130 mV
Probability of corrosion occurrence >90% "Uncertain"
