A Compilation of Elevated 

Temperature Concrete 

Material Property Data and 

Information for Use in 

Assessments of 

Nuclear Power Plant 

Reinforced Concrete 

Structures

Office of Nuclear Regulatory Research 

NUREG/CR-7031 

ORNL/TM-2009/175 
A Compilation of Elevated 

Temperature Concrete 

Material Property Data and 

Information for Use in 

Assessments of 

Nuclear Power Plant 

Reinforced Concrete 

Structures 

Manuscript Completed: October 2010 

Date Published: December 2010 

Prepared by 

D.J. Naus 

Oak Ridge National Laboratory 

Managed by UT-Battelle, LLC 

Oak Ridge, TN 37831-6283 

H.L. Graves, NRC Project Manager 

NRC Job Code N6511 

NUREG/CR-7031 

ORNL/TM-2009/175 

iii

ABSTRACT

The objective of this limited study was to provide a compilation of data and information on the effects of elevated 

temperature on the behavior of concrete materials for use in assessments of nuclear power plant reinforced concrete 

structures that are subjected temperatures in excess of the current American Society of Mechanical Engineers Code 

limitations. In meeting this objective the physicochemical processes in Portland cement concrete as a function of 

temperature are noted. The general behavior of Portland cement, aggregate, and concrete materials under elevated 

temperatures is summarized. Data and information on the effect of elevated temperature and testing conditions on 

the mechanical and physical properties of concrete are presented. Mechanical property-related items addressed 

include: stress and strain characteristics, Poisson’s ratio, modulus of elasticity, compressive strength, thermal 

cycling, tensile strength, shrinkage and creep, concrete-steel reinforcement bond strength, fracture energy and 

fracture toughness, long-term exposure, radiation shielding effectiveness, and multiaxial conditions. Physical 

properties and thermal effects addressed include: porosity and density, coefficient of thermal expansion, thermal 

conductivity, thermal diffusivity, specific heat, heat of ablation and erosion rates, moisture diffusion and pore 

pressure, and simulated hot spots. A general description of heavyweight concrete materials utilized for radiation 

shielding is provided and the effect of elevated temperature on properties of several shielding concretes is identified. 

Design codes and standards that address concrete under elevated temperature conditions are described. Examples of 

methods that can be utilized for assessment of concrete exposed to high temperatures are identified. Temperaturedependent properties of mild steel and prestressing materials for use with Portland cement concretes are provided. 
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FOREWORD

Under normal conditions most nuclear plant concrete structures are subjected to a range of temperature no more 

severe than that imposed by ambient environmental conditions. However, there are cases where these structures 

may be exposed to much higher temperatures (e.g., building fires and chemical and metallurgical applications in 

which the concrete is in close proximity to furnaces). Also designs of some new generation reactor concepts 

indicate that concrete may be exposed to long-term steady-state temperatures in excess of the present American 

Society of Mechanical Engineers Boiler and Pressure Vessel Code (ASME Code) limit of 650C. Under such an 

application the effect of elevated temperature on certain mechanical and physical properties may determine whether 

the concrete will maintain its structural integrity. 

The purpose of this research was to provide an overview of the effects of elevated temperature on the behavior of 

concrete materials. The effects of elevated temperatures on the properties of ordinary Portland cement concretes and 

constituent materials are summarized. The effects of elevated temperature on high-strength concrete materials are 

noted and the performance compared to normal strength concretes. A limited discussion of elevated temperature on 

radiation shielding concrete is also provided. Nuclear power plant and general civil engineering design codes are 

described. Finally, design considerations and analytical techniques for evaluating the response of reinforced 

concrete structures to elevated-temperature conditions are presented. 

The major findings contained in this NUREG/CR are: 1) deterioration of concrete’s mechanical properties can be 

attributed to physiochemical changes in cement paste, physiochemical changes in aggregate, and thermal 

incompatibility of cement paste and aggregate; 2) although a large amount of data and information on effects of

elevated temperature on concrete properties are available, many of the elevated temperature tests on concrete did not 

use representative materials nor representative nuclear power plant environmental conditions, and quantitative 

comparison of results can be difficult because of different test procedures, constituents and proportions, and testing 

conditions; and 3) several research projects have been conducted to investigate the behavior of reinforced concrete 

structures at elevated temperature, however, the overall level of effort has not been sufficient for establishment of 

widely accepted elevated-temperature concrete design procedures. 

On the basis of these findings, if a nuclear plant concrete structure in one of the proposed advanced reactors is 

required to maintain its functional and performance requirements at temperatures in excess of ASME Code limits, or 

at moderately elevated temperatures (e.g., 200-300 ̊C) for extended periods of time, techniques for optimizing the 

design of structural elements to resist these exposures should be investigated (i.e., the material selection and design).
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CONTENTS

Page
viii

Properties............................................................................................................................. B-1
ix

LIST OF FIGURES

Figure Page

1.1 Physiochemical processes in Portland cement concrete during heating.....................................................2

2.1 Weight change and residual compressive strength of siliceous gravel and 

limestone concretes as a function of temperature ......................................................................................3

2.2 Examples of dependence of concrete’s thermo-hydral-mechanical properties 

on heat-load-time sequence of the test regime ...........................................................................................5

2.3 Thermal strains of aggregate prisms during heat cycle to 600 ̊C at 2 ̊C/min.............................................6

2.4 Examples of range of thermal stabilities and processes that take place 

in aggregates during heating ......................................................................................................................6

2.5 X-ray diffraction results for mortar samples of concrete monitoring 

Ca(OH)2 and Ca(CO)3 fired at different temperatures...............................................................................8

2.6 OPC paste specimens immediately after and two days after 800 ̊C heat treatment ...................................9

2.7 Thermal strain during first heat cycle to 600 ̊C of initially moist cementitious materials.........................9

2.8 Ultimate compressive strength and modulus of elasticity of

hydrated Portland cement paste (w/c = 0.3) at elevated temperature .......................................................10

2.9 Residual and hot strength of Portland cement paste as a function of temperature ...................................10

2.10 Effect of temperature on residual compressive strength of cement paste materials.................................11

2.11 Residual compressive strengths of cement paste and cement paste with polypropylene fibers...............12

2.12 Residual flexural strengths of cement paste and cement paste with polypropylene fibers.......................12

2.13 Schematic of concrete deterioration under moderate heating and cooling...............................................13

2.14 Change in morphology of hardened concrete after exposure to elevated temperature.............................14

2.15 Differential thermal analysis of selected normal weight concretes..........................................................14

2.16 Schematic of temperature and loading histories ......................................................................................16

2.17 Stress-strain diagrams for sealed and unsealed limestone concrete .........................................................17

2.18 Influence of temperature on the stress-strain relation of unsealed quartz aggregate 

and sandstone aggregate concrete ............................................................................................................18

2.19 Effect of elevated temperature on stress-strain behavior of a quartz concrete .........................................18

2.20 Stress-strain relationship for basalt concrete derived in strain-rate controlled tests ................................19

2.21 Stress-strain curves for expanded shale lightweight masonry concrete tested at temperature .................19

2.22 Comparison of high-strength and normal-strength concrete load-deformation curves............................20

2.23 Stress-strain relationships for several concrete types before and after exposure to 200 ̊C ......................20

2.24 Stress-strain curves for high-strength siliceous aggregate concrete without 

and with steel fiber reinforcement ...........................................................................................................21

2.25 Stress-strain curves for high-strength carbonate aggregate concrete without 

and with steel fiber reinforcement ...........................................................................................................21

2.26 Comparison of compressive stress-strain curves for different strain rates at 20 ̊ and 65 ̊C.....................22

2.27 Stress-strain relationship of normal concrete with specimens loaded during the heating period.............22

2.28 The wffects of temperature and preloading on unsealed specimens 

tested at temperature: 70 MPa concrete ..................................................................................................23

2.29 The wffects of temperature and preloading on unsealed specimens 

tested at temperature: ultra high-performance concrete ..........................................................................23

2.30 Temperature-dependent stress-strain curves in unstressed condition for 

specimens tested at temperature...............................................................................................................24

2.31 Temperature-dependent stress-strain curves in stressed condition for specimens 

tested at temperature ................................................................................................................................24

2.32 Test specimen for evaluation of cyclic and sustained thermal loads on concrete stress-strain curve ......25

2.33 Effect of sustained temperature and thermal cycling on concrete stress-strain curve ..............................26

2.34 Effect of moisture condition and temperature on concrete stress-strain curve.........................................26

2.35 Effect of cooling regime on stress-strain curves following thermal exposure .........................................27

2.36 Direct uniaxial tensile stress-strain results for an unsealed 

high performance concrete tested at several temperatures.......................................................................28

2.37 Biaxial test system ...................................................................................................................................28

2.38 Biaxial test results: 1: 2 = 1: 0 (uniaxial)..............................................................................................29
x

2.39 Biaxial test results: 1: 2 = 1: 0.4 ...........................................................................................................29

2.40 Biaxial test results: 1: 2 = 1: 1 ..............................................................................................................30

2.41 Effect of temperature on biaxial stress-strain curves for tension-compression loading...........................31

2.42 Instantaneous stress-strain curve for concrete at ambient temperature ....................................................33

2.43 Instantaneous stress-strain curves at elevated temperature: comparison of 

predicted with experimental data ............................................................................................................34

2.44 Poisson’s ratio for uniaxially loaded concrete at high temperature .........................................................38

2.45 Poisson’s ratio as a function of temperature for a quartzite concrete.......................................................38

2.46 Effect of aggregate type and concrete strength on Poisson’s ratio...........................................................39

2.47 Poisson's ratio results for sealed and unsealed concrete specimens.........................................................39

2.48 Variation of Poisson's ratio of basalt aggregate concrete with number of 

thermal cycles (ambient to 176.7 ̊C to ambient) ......................................................................................40

2.49 Variation of Poisson's ratio of basalt aggregate concrete with length of exposure to 176.7 ̊C ................40

2.50 Temperature dependence of the concrete modulus of elasticity (normalized).........................................41

2.51 Modulus of elasticity of different concretes at elevated temperature.......................................................42

2.52 Influence of water-cement ratio on modulus of elasticity of concrete at elevated temperature ...............42

2.53 Normalized modulus of elasticity vs temperature relationships: NSC and HSC ....................................43

2.54 Effect of temperature on the modulus of elasticity of concrete: hot and cold test results........................43

2.55 Compilation of data on modulus of elasticity vs temperature ..................................................................44

2.56 Compilation of data on relative modulus of elasticity vs temperature – ordinary Portland cement 

concretes and concretes containing supplementary cementitious materials.............................................45

2.57 Compilation of data on relative modulus of elasticity vs temperature –

ordinary Portland cement concretes.........................................................................................................45

2.58 Compilation of data on relative modulus of elasticity vs temperature - concretes containing 

supplementary cementitious materials.....................................................................................................46

2.59 Compilation of data on residual modulus of elasticity vs temperature – ordinary Portland cement 

concretes and concretes containing supplementary cementitious materials.............................................46

2.60 Compilation of data on residual modulus of elasticity vs temperature –

ordinary Portland cement concretes.........................................................................................................47

2.61 Compilation of data on residual modulus of elasticity vs temperature –

concretes containing supplementary cementitious materials ...................................................................47

2.62 Compilation of data on relative modulus of elasticity vs temperature for 

preloaded specimens – ordinary Portland cement concretes and concretes 

containing supplementary cementitious materials ...................................................................................48

2.63 Effect of cooling regime on initial tangent modulus following thermal exposure ...................................48

2.64 Effect of cooling regime on relative residual tangent modulus following thermal exposure ...................49

2.65 Normalized residual modulus of elasticity results for specimens subjected to 

elevated temperature and quenched in water prior to testing ...................................................................50

2.66 Compilation of data on relative modulus of elasticity vs temperature for 

lightweight/insulating concrete specimens – ordinary Portland cement concretes 

and concretes containing supplementary cementitious materials.............................................................51

2.67 Compilation of data on relative modulus of elasticity versus temperature for fibrous concretes.............51

2.68 Influence of type of cement on strength loss of mortars ..........................................................................54

2.69 Influence of partial replacement of (a) OPC and (b) slag on residual compressive strength ...................54

2.70 Effect of water-cement ratio on residual compressive strength of ordinary Portland cement..................55

2.71 Residual normalized strength vs temperature ..........................................................................................55

2.72 Residual compressive strength vs temperature ........................................................................................56

2.73 Effect of curing conditions prior to elevated-temperature exposure on

relative compressive strength...................................................................................................................56

2.74 Effect of temperature on compressive strength ratio of concrete.............................................................57

2.75 Effect of curing age and exposure condition on residual compressive strength 

(S = sealed, U = unsealed) .......................................................................................................................57

2.76 Effect of rate of heating on residual strength of slowly cooled concrete .................................................58

2.77 Effect of exposure time on residual strength of coral sand and basalt sand concretes.............................59

2.78 Relative strength development of concrete exposed to elevated temperature .................................................59

2.79 Effect of exposure time on residual strength of carbonate and siliceous aggregate concretes.................60
xi

2.80 Compressive strength of concretes with limestone and other aggregate types ........................................60

2.81 Effect of temperature exposure on compressive strength of concrete: tested cold...................................61

2.82 Effect of temperature exposure on compressive strength of concrete: tested hot ....................................62

2.83 Effect of temperature on uniaxial compressive strength of PCPV unsealed concretes 

(H = hot, C = cold)...................................................................................................................................62

2.84 Compilation of data on concrete compressive strength vs temperature ...................................................63

2.85 Compilation of data on relative compressive strength vs temperature – ordinary Portland cement 

concretes and concretes containing supplementary cementitious materials.............................................64

2.86 Compilation of data on relative compressive strength vs temperature – ordinary Portland cement 

concretes ..................................................................................................................................................64

2.87 Compilation of data on relative compressive strength vs temperature – ordinary Portland cement 

concretes, fc’ < 60 MPa............................................................................................................................65

2.88 Compilation of data on relative compressive strength vs temperature – ordinary Portland cement 

concretes, fc’ > 60 MPa............................................................................................................................65

2.89 Compilation of data on relative compressive strength vs temperature – concretes containing 

supplementary cementitious materials.....................................................................................................66

2.90 Compilation of data on residual compressive strength vs temperature – ordinary Portland cement 

concretes and concretes containing supplementary cementitious materials.............................................67

2.91 Compilation of data on residual compressive strength vs temperature – ordinary Portland cement 

concretes ..................................................................................................................................................67

2.92 Compilation of data on residual compressive strength vs temperature – ordinary Portland cement 

concretes, fc’ < 60 MPa............................................................................................................................68

2.93 Compilation of data on residual compressive strength vs temperature – ordinary Portland cement 

concretes, fc’ > 60 MPa............................................................................................................................68

2.94 Compilation of data on residual compressive strength vs temperature – ordinary Portland cement 

concretes with carbonate-type aggregates, fc’ < 60 MPa .........................................................................69

2.95 Compilation of data on residual compressive strength vs temperature – ordinary Portland cement 

concretes with siliceous- or gravel-type aggregates, fc’ < 60 MPa ..........................................................69

2.96 Compilation of data on residual compressive strength vs temperature – concretes containing 

supplementary cementitious materials.....................................................................................................70

2.97 Compilation of data on residual compressive strength vs temperature – concretes containing 

supplementary cementitious materials, fc’ < 60 MPa ..............................................................................70

2.98 Compilation of data on residual compressive strength vs temperature – concretes containing 

supplementary cementitious materials, fc’ > 60 MPa ..............................................................................71

2.99 Compilation of data on residual compressive strength vs temperature – concretes containing 

silica fume supplementary cementitious materials...................................................................................71

2.100 Compilation of data on residual compressive strength vs temperature – concretes containing fly 

ash supplementary cementitious materials...............................................................................................72

2.101 Compilation of data on residual compressive strength vs temperature – concretes containing 

metakaolin supplementary cementitious materials ..................................................................................72

2.102 Compilation of data on residual compressive strength vs temperature – concretes containing slag 

supplementary cementitious materials.....................................................................................................73

2.103 Compilation of data on compressive strength vs temperature for

Portland cement concretes and concretes containing supplementary cementitious materials..................74

2.104 Compilation of data on relative compressive strength vs temperature for 

preloaded specimens – ordinary Portland cement concretes....................................................................74

2.105 Compilation of data on relative compressive strength vs temperature for 

pre-loaded specimens – concretes containing supplementary cementitious materials.............................75

2.106 Effect of cooling regime on relative residual compressive strength following thermal exposure............75

2.107

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2.108

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2.109

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2.110

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2.111

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2.112

concrete specimens – ordinary Portland cement concretes......................................................................79

2.113 Compilation of data on residual compressive strength vs temperature for lightweight/insulating 

concrete specimens – concretes containing supplementary cementitious materials ................................79

2.114 Compilation of data on relative compressive strength vs temperature – fibrous concrete .......................80

2.115 Compilation of data on residual compressive strength vs temperature – fibrous concrete ......................80

2.116 Effect of different cooling regimes on residual compressive and tensile strengths of 0.1 % 

polypropylene fiber reinforced concrete – 88.6 MPa compressive strength ............................................81

2.117 Effect of different cooling regimes on residual compressive and tensile strengths of 0.3% 

polypropylene fiber reinforced concrete – 83.4 MPa compressive strength ............................................81

2.118 Effect of different cooling regimes on residual compressive and tensile strengths of 0.07% 

polypropylene plus 1% steel fiber reinforced concrete – 87.2 MPa compressive strength ......................82

2.119 Effect of temperature cycles on limestone concretes...............................................................................84

2.120 Influence of thermal cycling on - response of sealed concrete tested at 149°C...................................85

2.121 Comparison of effect of elevated temperature on compression and tensile strengths of concretes 

fabricated using different types of conventional aggregate materials......................................................86

2.122 Effect of elevated temperature curing on residual compressive and tensile strengths and modulus........87

2.123 Effect of length of room temperature and elevated temperature curing on relative 

residual tensile strength results for sealed and unsealed specimens.........................................................87

2.124 Relationship between residual ratios and heating temperatures, 

cement types, and water-cement ratios ....................................................................................................88

2.125 Residual splitting-tensile strength of concrete after one month’s exposure

at various temperatures ............................................................................................................................89

2.126 Residual tensile strengths of HSC and NSC ............................................................................................89

2.127 Effect of curing age and exposure condition on residual splitting-tensile strength of

a flint-gravel aggregate concrete..............................................................................................................90

2.128 Relative variations in splitting-tensile strength........................................................................................91

2.129 Comparison of the effect of elevated-temperature exposure on residual compressive (fcu), tensile 

(splitting-tension) (ft), and bend strengths (notched beams) (fr) of siliceous aggregate concrete ............91

2.130 Residual splitting-tensile and bending strengths after thermal exposure .................................................92

2.131 Effect of test conditions on flexural strength of gravel and limestone concretes, unsealed .....................93

2.132 Flexural strength of autoclaved gravel and limestone concretes..............................................................93

2.133 Variation of flexural strength with temperature and exposure period for 

concrete mixes with and without fly ash..................................................................................................94

2.134 Normalized residual property results for specimens subjected to elevated temperature and 

quenched in water prior to testing............................................................................................................95

2.135 Effect of cooling rate on residual flexural strength..................................................................................96

2.136 Effect of cooling rate on residual splitting-tensile strength......................................................................96

2.137 Effect of temperature on splitting-tensile strength of plain concrete .......................................................97

2.138 Effect of temperature and fiber reinforcement type and quantity on

splitting-tensile strength of specimens tested at temperature ...................................................................97

2.139 Effect of temperature and silica fume content on residual compressive and 

splitting-tensile strengths of a lightweight concrete................................................................................98

2.140 Effect of several factors on autogenous shrinkage of concrete ..............................................................100

2.141 Variation of long-term shrinkage of concrete with relative humidity of storage ...................................101

2.142 Shrinkage measurements obtained from sealed cylinders cast from Wylfa concrete mix .....................101

2.143 Shrinkage design limits for Wylfa PCPV ..............................................................................................102

2.144 Interaction of loading, restraint, and humidity conditions .....................................................................103

2.145 Illustration of concrete creep under simultaneous drying and loading...................................................104

2.146 Effect of aggregate type on drying shrinkage and creep of concrete .....................................................104

2.147 Creep of concrete stored at different relative humidities .......................................................................105
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2.148 Typical creep-time curves under multiaxial compression: (a) biaxial, and (b) triaxial..........................105

2.149 Reversibility of drying shrinkage and creep: (a) drying and rewetting, and 

(b) loading and unloading under uniaxial compression .........................................................................106

2.150 Relation between creep and logarithm of time under load for concretes stored at different 

temperatures and loaded to different stress levels..................................................................................107

2.151 ..................108

2.152 ...................................................108

2.153 .....109

2.154 .....110

2.155 ...........................111

2.156 ..............................................................................111

2.157 ........................................................112

2.158 ........................................................112

2.159 .........................................................113

2.160 .........................114

2.161 .................................114

2.162 .....................................................115

2.163 .........................................................................................116

2.164 ..................................................................................................117

2.165

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2.166

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2.167 ..................................................119

2.168 .......................................................120

2.169 ....................................................121

2.170 .......................................................121

2.171 ......................................................122

2.172 .............................................................................122

2.173 .......................................................................123

2.174

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2.175 .............................................................................124

2.176 .............................................................125

2.177

.................................................................................................................................................125

2.178 Effect of bar diameter on bond strength after elevated-temperature exposure 

for ribbed and plain round bars..............................................................................................................126

2.179 Bond-slip relationship at elevated temperature for cold deformed steel and 

prestressing steel ....................................................................................................................................127

2.180 Relative bond strength as a function of temperature ..............................................................................128

2.181 The effect of elevated-temperature exposure time on the relative residual bond strength of 

No. 3 bars embedded in a concrete cube................................................................................................129

2.182 Bond between concrete and deformed bars exposed to high temperature .............................................129

2.183 Relative variation in bond strength at start of pull-out for a hard sandstone aggregate concrete 

after various heating periods at 175 ̊C...................................................................................................130

2.184 Effect of elevated temperature on residual strength ratios.....................................................................130

2.185 Residual bond stress versus slip relationships at various temperatures: 16-mm diameter bars with 

55 mm cover and stressed at 3.7 MPa while heated...............................................................................131

2.186 Stressed and unstressed hot and residual bond stress results at various temperatures ...........................132

2.187 Variation of maximum bond stress with temperature for different cover depths...................................132

2.188 Effect of load cycling on the residual bond strength for 16-mm-diameter deformed rebar...................133

2.189

2.190

2.191

2.192 Effect of temperature on fracture energy (Gf)........................................................................................137
xiv

2.193 Effect of exposure period (th) on fracture energy (Gf) of normal strength concrete...............................138

2.194 Effect of curing period on fracture energy (Gf)......................................................................................138

2.195 Effect of exposure temperature on residual fracture energy of ordinary 

and high-strength concretes ...................................................................................................................139

2.196 Effect of temperature on residual fracture toughness of a concrete .......................................................139

2.197 Effect of thermal cycling on residual fracture toughness of a concrete .................................................140

2.198 Comparison of laboratory and actual sample long-term compressive strength data ..............................142

2.199 Relationship of relative strength and temperature of mass concrete (sealed).............................................142

2.200 Relationship of elasticity ratio and temperature of mass concrete (sealed)...................................................143

2.201 Relationship of strength and elasticity and temperature of unsealed concrete.............................................143

2.202 Relative residual compressive strength of limestone concrete after 4 month exposure to 

various temperatures (up to 450°C) .......................................................................................................144

2.203 Long-term (3.5-year) heating effect on compressive strength and modulus..........................................145

2.204 Effect of thermal cycling on compressive strength and modulus...........................................................146

2.205 Effect of exposure temperature on residual compressive strength, tensile strength, 

and modulus of elasticity after 90-d exposure .........................................................................................148

2.206 Weight loss of heated concrete ..............................................................................................................149

2.207 Compressive strength and modulus of elasticity of heated concretes ....................................................149

2.208 Shear strength of heated concrete ..........................................................................................................150

2.209 Residual compressive strength vs exposure time: unsealed specimens ................................................151

2.210 Residual compressive strength vs exposure time: sealed specimens .....................................................152

2.211 Residual modulus of elasticity vs exposure time: unsealed specimens.................................................152

2.212 Residual modulus of elasticity vs exposure time: sealed specimens......................................................153

2.213 Change in weight vs exposure time: unsealed specimens......................................................................153

2.214 Change in weight vs exposure time: sealed specimens.........................................................................154

2.215 Change in splitting-tensile strength ratio versus exposure time: sealed specimens ..............................154

2.216 Thermal neutron distribution in ordinary concrete as a function of temperature ...................................155

2.217 Multiaxial test methods..........................................................................................................................157

2.218 Multiaxial test results: (a) strength of concrete under biaxial loading and (b) triaxial failure 

envelopes within octahedral plane 0 = 34.5 MPa................................................................................158

2.219 Complete stress-strain curves for biaxial compression for different strain ratios ( )............................159

2.220 Biaxial strength envelopes in stress space and strain space ...................................................................160

2.221 Stress ratios considered(f2/f1).................................................................................................................160

2.222 Biaxial strength envelopes.....................................................................................................................161

2.223 Stress-strain relationships for wall and dome concrete under: (a) biaxial compression, 

(b) combined compression and tension, and (c) biaxial tension ............................................................162

2.224 Failure modes under biaxial loading for different stress ratios..............................................................163

2.225 Influence of aggregate content and water-cement ratio on uniaxial and

biaxial results for tests at 300°C and 600°C...........................................................................................164

2.226 Influence of aggregate content and water-cement ratio on modulus of elasticity 

results for tests at 300°C and 600°C ......................................................................................................165

2.227 Biaxial compressive strength at different temperatures...............................................................................166

2.228 Effect of loading platen type on concrete biaxial strength .....................................................................166

2.229 Effect of temperature on concrete biaxial strength ................................................................................167

2.230 Change in tensile strength with temperature ..........................................................................................168

2.231 Effect of temperature on the relationship between tensile strength and stress ratios.............................168 

2.232 Effect of temperature on the concrete biaxial tension-compression failure envelopes ..........................169

2.233 mac2T apparatus for multiaxial compression of concrete at elevated temperature .................................170

2.234 Average temperature-strain relationships from all LHS and HS tests: (a) load-induced thermal 

strain and shrinkage, (b) load-induced thermal strain ............................................................................171

2.235 Variation of Portland cement paste porosity with temperature ..............................................................173

2.236 True density, bulk density, and porosity of cement paste vs temperature..............................................173

2.237 Density of concretes having different coarse aggregate types ..............................................................174

2.238 Density of a nuclear power plant concrete vs temperature ....................................................................175

2.239 Density of normal and lightweight concretes vs temperature ................................................................175

2.240 Length change of Portland cement paste specimens at elevated temperature ........................................176
xv

2.241 Linear thermal expansion of various rocks with temperature ................................................................178

2.242 Coefficients of thermal expansion of neat cements, mortars, and concretes..........................................178

2.243 Thermal strain of Portland cement, mortar, and concrete on heating ....................................................179

2.244 Thermal expansion of limestone aggregate concretes............................................................................180

2.245 Thermal expansion of siliceous aggregate concretes .............................................................................180

2.246 Variation of thermal strain with temperature and load for Portland cement concretes: 

0 = no load, 0.1 = load corresponding to 10% fc’ ..................................................................................181

2.247 Effect of temperature on thermal conductivity of initially saturated concrete .......................................183

2.248 Thermal conductivity of normal and lightweight concretes...................................................................184

2.249 Thermal conductivity of ordinary concretes with different aggregates.......................................................184

2.250 Thermal conductivity as a function of density and moisture content.....................................................185

2.251 Thermal conductivity as a function of temperature ...............................................................................185

2.252 Thermal conductivity of a nuclear power plant concrete .......................................................................186

2.253 Comparison of experimental results and thermal conductivity predicted by 

different models for normal-strength concretes .....................................................................................187

2.254 Thermal diffusivity of limestone aggregate concrete.............................................................................189

2.255 Thermal diffusivity of siliceous aggregate concrete ..............................................................................189

2.256 Thermal diffusivity of a nuclear power plant concrete ..........................................................................190

2.257 Thermal diffusivity variation of a Korean nuclear plant concrete as a function of temperature ............190

2.258 Effect of temperature on thermal diffusivity of normal weight concrete 

excluding latent heat effects...................................................................................................................191

2.259 Thermal diffusivity of normal and lightweight concretes......................................................................191

2.260 Specific heats of various concretes........................................................................................................192

2.261 Specific heat capacity of limestone aggregate concrete .............................................................................193

2.262 Specific heat capacity of siliceous aggregate concrete ..........................................................................193

2.263 Specific heat of a nuclear power plant concrete.....................................................................................194

2.264 Specific heat of a Korean nuclear power plant concrete ........................................................................194

2.265 Specific heat of normal and lightweight concretes ................................................................................195

2.266 Comparison of experimental results and specific heat predicted by 

different models for normal-strength concretes .....................................................................................197

2.267 Penetration depths vs time for various melt temperature conditions .....................................................198

2.268 Cross section of test specimen after MEK-T1A test ..............................................................................199

2.269 Illustration of process of pressure buildup in a concrete element under 

the effect of a pressure gradient .............................................................................................................201

2.270 Pore pressure build up in sealed concrete specimen ..............................................................................203

2.271 Temperature distribution and pore pressure data for 30.5-cm-thick heated concrete wall.....................203

2.272 Pressure distribution for vented and sealed pore pressure test at given times........................................204

2.273 Phase diagrams for water in concrete at 887 days for concretes of various lengths 

and hot face temperature of 125 ̊C.........................................................................................................204

2.274 Changes in moisture content in specimen after heating and moisture content profile 

at end of test period................................................................................................................................206

2.275 Simulated section of mass concrete wall ...............................................................................................206

2.276 Details of simulated section of mass concrete wall and measurement positions ...................................207

2.277 Temperature distribution at various times in simulated mass concrete wall with 

and without a venting system.................................................................................................................208

2.278 Moisture distribution at various times in simulated mass concrete wall section 

with and without a venting system..........................................................................................................208

2.279 Water discharge from vent pipe of simulated mass concrete wall section .............................................209

2.280 Change in strain distribution with time in simulated mass concrete wall section 

with and without venting .......................................................................................................................209

2.281 Compressive strength test results at selected locations in simulated mass concrete wall section 

with and without venting .......................................................................................................................210

2.282 Modulus of elasticity test results at selected locations in 

simulated mass concrete wall section with and without venting............................................................210

2.283 Measured relative humidity distributions through Barsebäck containment wall at two locations .........211

2.284 Isometric of ORNL thermal cylinder test structure................................................................................213
xvi

2.285 Cross section showing discoloration in the vicinity of the hot-spot heating ..........................................214

2.286 Full-scale Oldbury hot-spot model.........................................................................................................215

3.1

..........................231

3.2

...............................................................................234

3.3 Weight loss and residual compressive strength after thermal treatment for 

baryte and quartz aggregate concretes ...................................................................................................236

3.4

..........................................................................................................238

3.5 ...........239

3.6 ..................................................................................240

3.7 ........241

4.1 BS 8110 design curves for strength variation with the temperature of dense concrete and 

lightweight concrete...............................................................................................................................246

4.2 ........247

4.3 ..........................................................248

4.4 ................248

4.5

..............................................................................................................................249

4.6

..........................................................................................................249

4.7

........................................................................................250

4.8 ..............................................250

4.9 ..........251

4.10 .....251

5.1 Example of evaluation process for structure subjected to thermal excursion 

such as resulting from a fire ...................................................................................................................253

5.2 Discoloration and cracking in concrete resulting from elevated temperature exposure .........................256

5.3 Appearance of flint-aggregate concrete after heating ............................................................................257

5.4 .......................................259

5.5 Illustration of application of drilling resistance method to fire-damaged concrete column...................260

5.6

............................................................................................................261

5.7 ....................................262

5.8

...................................................263

B.1 Stress-strain curves for structural steel (ASTM A36)............................................................................B-1

B.2 Stress-strain curves for structural steel (ASTM A36) at high temperatures...........................................B-2

B.3 Stress-strain curves for steel reinforcement (S220 and S420) ...............................................................B-2

B.4 Stress-strain relationships of reinforcing bars at elevated temperature ..................................................B-3

B.5 Effect of temperature on stress-strain curves of a prestressing steel (ASTM A421). ............................B-3

B.6 Stress-strain curve for 1770 MPa 5-mm-diameter low-relaxation prestressing steel wire 

tested at temperature and at room temperature ......................................................................................B-4

B.7 Modulus of elasticity at elevated temperatures for structural steel and steel reinforcing bars...............B-4

B.8 Influence of temperature on Young's modulus and elongation of reinforcing bars ...............................B-5

B.9 Normalized modulus of elasticity results for 1770 MPa 5-mm-diameter low-relaxation 

prestressing steel wire tested at temperature ..........................................................................................B-5

B.10 Effect of elevated temperature on yield and ultimate strength of a

typical structural steel (ASTM A36)......................................................................................................B-6

B.11 Yield strength of steel reinforcement vs temperature ............................................................................B-6

B.12 Strength of selected steels at high temperature ......................................................................................B-7

B.13 Effect of elevated temperature on residual yield and ultimate strength, 

and elongation of ASTM A615 steel rebar ............................................................................................B-7
xvii

B.14 Effect of exposure temperature on residual yield strength of steel reinforcement (S220 and S420) .....B-8

B.15 Effect of exposure temperature on residual tensile strength of 

steel reinforcement (S220 and S420)....................................................................................................B-8

B.16 Yield strength and ultimate tensile strength of reinforcing bars at elevated temperature ......................B-9

B.17 Residual yield and ultimate strength of high-strength weldable rebars..................................................B-9

B.18 Effect of elevated temperature on yield and ultimate strength of 

a prestressing steel (ASTM A421).......................................................................................................B-10

B.19 Normalized yield and tensile strength results for 1770 MPa 5-mm-diameter 

low-relaxation prestressing steel wire tested at temperature ................................................................B-11

B.20 Normalized residual yield and tensile strength results for 1770 MPa 

5-mm-diameter low-relaxation prestressing steel wire tested after heating .........................................B-12

B.21 Relationship between temperature and normalized tensile strength for 

hot-rolled reinforcing steel bar and cold-drawn prestressing steel wire...............................................B-12

B.22 Tensile strength results of different prestressing steels in the heated state ..........................................B-13

B.23 Effect of exposure temperature on toughness of steel reinforcement (S220 and S420).......................B-13

B.24 Residual normalized absorbed energy of high-strength weldable rebars.............................................B-14

B.25 Density of different steels....................................................................................................................B-14

B.26 Mean specific heat of different steels...................................................................................................B-15

B.27 Thermal conductivity of different steels ..............................................................................................B-15

B.28 Thermal diffusivity of different steels..................................................................................................B-16

B.29 Coefficient of expansion of different steels .........................................................................................B-16

B.30 Thermal expansion of ferritic steel at high temperature.......................................................................B-17
xviii

List of Tables

Table Page

2.1 Influence of environmental factors on heated concrete..............................................................................4

2.2 Values for parameters to describe ascending branch of stress-strain relationship for concrete at 

elevated temperature ................................................................................................................................36

2.3 Properties of concrete mixes at the age of 28 days ..................................................................................50

2.4 Concrete mix properties at the age of 28 days .........................................................................................94

2.5 Residual splitting-tensile strengths of plain and fibrous concretes after temperature exposure...............98

2.6 Mix proportions for concrete mixtures ..................................................................................................147

2.7 Concrete mixture proportions ................................................................................................................150

2.8 Preferred concrete physical properties...................................................................................................172

2.9 Coefficients of thermal expansion of different rocks and concrete at normal temperature....................177

2.10 Effect of elevated temperatures on the coefficient of thermal expansion on selected rocks..................177

2.11 Coefficients of thermal expansion at elevated temperature for 

concretes made with different aggregates..............................................................................................179

2.12 Thermal conductivities at ambient temperature of concrete constituents ..............................................182

2.13 Typical values of thermal conductivity..................................................................................................183

2.14 Thermal diffusivity values for concrete with different coarse aggregate ...............................................188

2.15 Heat of ablation values of different concretes........................................................................................198

2.16 Estimated heats of transformation for quartzitic and limestone aggregate concretes 

exposed to elevated temperatures...........................................................................................................199

2.17 Erosion of different concretes by liquid steel.........................................................................................200

2.18 Summary of conditions for simulated mass concrete wall section test ..................................................207

3.1 Examples of aggregate materials used to produce heavyweight concretes............................................229

3.2 Examples of mechanical and physical properties reported in literature 

for heavyweight aggregate concretes.....................................................................................................230

3.3 Effect of temperature on properties of a serpentine-iron concrete .........................................................232

3.4 Effect of temperature on properties of a magnetite-serpentine concrete................................................232

3.5 Effect of elevated temperature on the physical and mechanical properties of 

hydrous-iron aggregate concrete ............................................................................................................233

3.6 Effect of elevated temperature on the physical and mechanical properties of 

magnetite aggregate concrete.................................................................................................................234

3.7 Effect of elevated temperature on the physical and mechanical properties of 

a ferrophosphorus aggregate concrete....................................................................................................235

3.8 Effect of elevated temperature on mechanical properties of a baryte concrete......................................237

4.1 Condition categories and temperature limits for concrete and prestressing systems for PCRVs...........245

5.1 Listing of possible nondestructive approaches for use in assessment of fire-damaged concrete ...........254

5.2 Simplified visual concrete fire damage classification............................................................................254

5.3 Initial assessment of damage and probable treatment required..............................................................255

5.4 Indication of concrete color change on heating .....................................................................................258

5.5 Appearance of concrete after being subjecting to very high temperatures.............................................258

5.6 Visual indications of elevated temperature effects on concrete .............................................................261

5.7 Criteria used to classify the extent of microcracking in concrete ..........................................................263
xix

Acknowledgment

The author would like to acknowledge the continuing support and guidance throughout the program provided by the

U.S. Nuclear Regulatory Commission Technical Monitor, Mr. Herman L. Graves, III.
xx
xxi
xxii
1

1 INTRODUCTION

Under normal conditions, most concrete structures are subjected to a range of temperature no more severe than that 

imposed by ambient environmental conditions. However, there are important cases where these structures may 

experience much higher temperatures (e.g., jet aircraft engine blasts, building fires, chemical and metallurgical 

industrial applications in which the concrete is in close proximity to furnaces, and some nuclear power-related 

postulated accident conditions). Under elevated temperature exposure reinforced concrete structures can fail in a 

number of different ways [1.1,1.2]. For load-bearing slabs, if the strength of the steel reinforcement is lost due to 

heating then there may be bending or tensile strength failure. Reinforced members may also fail when the bond 

between the concrete and reinforcement is lost, with associated concrete tensile failure. Shear or torsion failures are 

also influenced by concrete tensile strength, but are poorly defined experimentally. Finally, compressive failures are 

usually associated with temperature-related loss of concrete compressive strength in the compression zone. In 

practice, failure is related to structural performance in situ (e.g., restraint effects). 

Of primary interest in the present study is the behavior of reinforced concrete elements in designs of new-generation 

reactor concepts in which the concrete may be exposed to long-term steady-state temperatures in excess of the 

present American Society of Mechanical Engineers Pressure Vessel and Piping Code (ASME Code) limit of 65 C 

[1.3]. Secondary interests include performance of concrete associated with radioactive waste storage and disposal 

facilities and postulated design-basis accident conditions involving unscheduled thermal excursions. Under such 

applications the effect of elevated temperature on certain mechanical and physical properties of concrete may 

determine its ability to maintain structural integrity as well as its ability to continue to provide adequate structural 

margins. 

The performance of Portland cement-based materials under elevated temperature exposure is very complicated and 

difficult to characterize. Concrete’s thermal properties are more complex than for most materials because not only 

is the concrete a composite material whose constituents have different properties, but also its properties depend on 

moisture and porosity. Exposure of concrete to elevated temperature affects its mechanical and physical properties. 

The changes in properties result from three processes that take place at elevated temperature: (1) phase 

transformations (e.g., loss of free water at about 100 ̊C, decomposition of calcium hydroxide at about 450 ̊C, and 

crystal transformation of quartz at 573 ̊C from the - to the -form), (2) pore structure evolution (e.g., volume and 

surfaces of pores increase up to a temperature of about 500 ̊C and then decrease with further temperature increase, 

and (3) coupled thermo-hygro-chemo-mechancial processes (e.g., temperature gradients leading to thermal stresses, 

multiphase transport of water, and chemical changes that affect pore pressure and structure) [1.4]. Figure 1.1 

provides a summary of the physiochemical processes in Portland cement concrete during heating [1.1]. Under 

thermal loading elements could distort and displace, and, under certain conditions, the concrete surfaces could spall 

due to the build up of steam pressure. Because thermally-induced dimensional changes, loss of structural integrity, 

and release of moisture and gases resulting from the migration of free water could adversely affect plant operations 

and safety, a complete understanding of the behavior of concrete under long-term elevated-temperature exposure as 

well as both during and after a thermal excursion resulting from a postulated design-basis accident condition is 

essential for reliable design evaluations and assessments. Because the properties of concrete change with respect to 

time and the environment to which it is exposed, an assessment of the effects of concrete aging is also important in 

performing safety evaluations.

Bonded reinforcement (i.e., deformed bars) is provided to control the extent and width of cracks at operating 

temperatures, resist tensile stresses and computed compressive stresses for elastic design, and provide structural 

reinforcement where required by limit condition design procedures. Bonded reinforcement in nuclear power plant 

structures is often used in conjunction with prestressed steel. The prestressed steel provides the structural rigidity 

and the major part of the strength while the bonded reinforcement distributes cracks, increases ultimate strength and 

reinforces those areas not adequately strengthened by the prestressed steel, and provides additional safety for 

unexpected conditions of loading. Steel reinforcement is normally protected by the concrete against significant 

elevated temperature exposure because of concrete’s low thermal diffusivity that results in slow propagation of 

thermal transients. However, under certain conditions such as long-duration thermal exposure, thin-section 

members, or occurrence of concrete spalling, exposure the reinforcement to elevated temperature can occur. If the 

temperatures experienced by the steel are high enough, phase transformations can occur that produce changes in its 

physical and mechanical properties.
2

Figure 1.1 Physiochemical processes in Portland cement concrete during heating.

Source: Adaptation of Figure 2 in G.A. Khoury, ―Effect of Fire on Concrete and Concrete Structures,‖ Progress in 

Structural Engineering Materials 2, pp. 429-447, 2000.

This report contains physical and mechanical property data and information on effects of thermal loadings on 

reinforced concrete materials that either have been used to construct existing plants or that could be used for 

fabrication of future plants (e.g., high performance concretes). This report supplements previous information 

provided in NUREG/CR-6900 [1.5] by providing additional material properties data, relationships between 

properties and temperature for use in analytical assessments, information and mateial properties data on shielding 

concretes, identification of codes and standards that address concrete under elevated temperature conditions, and 

identifies several potential methods for assessment of concrete exposed to high temperatures. The primary

application of this information is to assessments of nuclear power plants that have concrete structures experiencing 

concrete temperatures above ASME Code limits or structures that have experienced temperatures above the Code 

limits. Although not a primary objective of this activity, limited information is provided to assist in the assessment 

of the response of concrete structures under thermal loadings representative of accident conditions. 

References

1.4 K. Willam and I. Rhee, ―Thermal Degradation Effects in Concrete Material Systems,‖ NSF-FHWA 

Workshop on Imaging and Simulation of Concrete Structures, Northwestern University, Evanston, Illinois, 

July 30-31, 2003.

1.5 D.J. Naus, ―The Effects of Elevated Temperature on Concrete Materials and Structures – A Literature 

Review,‖ NUREG/CR-6900, U.S. Nuclear Regulatory Commission, Washington, DC, March 2006.
3

2 EXPERIMENTAL RESULTS ON TEMPERATURE-DEPENDENT PROPERTIES AND 

PERFORMANCE OF PORTLAND CEMENT CONCRETE MATERIALS

Portland cements are manufactured by mixing finely divided calcareous materials (i.e., lime containing) and 

argillaceous materials (i.e., clay). The four compounds that make up more than 90% of the dry weight of the cement 

are tricalcium silicate (3CaO•SiO2), dicalcium silicate (2CaO•SiO2), tricalcium aluminate (3CaO•A12O3), and 

tetracalcium aluminoferrite (4CaO•A1203•Fe2O3). When water is added to Portland cement, an exothermic 

reaction occurs, and new compounds are formed (i.e., hydrated cement paste): tobermorite gel 

[(Ca5Si6O16(OH)2•4H2O)], calcium hydroxide, calcium aluminoferrite hydrate, tetracalcium aluminate hydrate, 

and calcium monosulfoaliminate. Mature cement paste is normally composed of 70–80% layered calcium-silicatehydrate (C-S-H) gel, 20% Ca(OH)2, and other chemical compounds [2.1]. The C-S-H gel structure is made up of 

three types of groups that contribute to bonds across surfaces or in the interlayer of partly crystallized tobermorite 

material: calcium ions, siloxanes, and water molecules. Bonding of the water within the layers (gel water) with 

other groups via hydrogen bonds determines the strength, stiffness, and creep properties of the cement paste. 

Tobermorite gel is the primary contributor to the cement paste structural properties. 

Concrete is a heterogeneous multiphase material with relatively inert aggregates that is held together by the hydrated 

Portland cement paste. When concretes are exposed to high temperatures, changes in mechanical properties and 

durability occur. It has been noted that there are two schools of thought on causes of thermal damage to Portlandcement-based materials when exposed to elevated temperature [2.2]. The first attributes the irreversible loss of 

strength and stiffness to microcracking induced by thermal mismatch of some mesoscopic properties and interfaces 

[2.3,2.4]. The second attributes thermal damage to the dehydration of the primary hydration products in the 

hardened cement paste matrix [2.5,2.6]. 

Under elevated-temperature exposure, Portland cement paste experiences physical and chemical changes that 

contribute to development of shrinkage, transient creep, and changes in strength. Nonlinearities in material 

properties, variation of mechanical and physical properties with temperature, tensile cracking, and creep effects 

affect the buildup of thermal forces, the load-carrying capacity, and the deformation capability (i.e., ductility) of the 

structural members. Property variations are due largely to changes in the moisture condition of the concrete 

constituents and the progressive deterioration of the cement paste-aggregate bond, which is especially critical where 

thermal expansion values for the cement paste and aggregate differ significantly. Concrete at room temperature may 

contain 15-20% water (by vol.) of which 10-12% can evaporate with the remainder chemically bound. The 

correlation between moisture change and properties is demonstrated in Figure 2.1 which provides a comparison of 

the weight change and residual compressive strength results for siliceous gravel and limestone coarse aggregate 

concretes as a function of temperature [2.7]. Weight change and residual compressive strength for siliceous gravel 

and limestone concretes both changed significantly at temperatures above about 800 ̊C. The bond region is affected

Figure 2.1 Weight change and residual compressive strength of siliceous gravel and 

limestone concretes as a function of temperature.

Source: O. Arioz, ―Effects of Elevated temperatures on Properties of Concrete,‖ Fire Safety Journal 42, pp. 516-

522, 2007.
4

by the surface roughness of the aggregate and its chemical/physical interactions [2.8]. Chemical interaction relates 

to the chemical reactions between the aggregate and cement paste that can be either beneficial or detrimental. 

Physical interaction relates to dimensional compatibility between aggregate materials and cement paste. Behavior of 

concrete at high temperature depends on exposure conditions (i.e., temperature-moisture-load-time regime) [2.8]. 

Curing influences the degree of hydration, while the temperature and load history prior to exposure to elevated 

temperature could have a significant effect on the behavior of the Portland cement paste and therefore the concrete. 

Concrete at elevated temperature is sensitive to the temperature level, heating rate, thermal cycling, and temperature 

duration (as long as chemical and physical transformations occur). Deterioration of mechanical properties of 

concrete due to thermal loading involves three material factors: (1) physicochemical changes in the cement paste, 

(2) physicochemical changes in the aggregate, and (3) thermal compatibility between the aggregate and the cement 

paste (e.g., pure cement paste exhibits two to three times greater thermal expansion than limestone) ; and are 

influenced by environmental factors such as temperature level, heating rate, applied loading, and external sealing 

influencing moisture loss from the surface [2.9]. Information presented in Table 2.1 provides more detailed 

Table 2.1 Influence of environmental factors on heated concrete

*** First order influence, ** Second order influence, * Third order influence.

Source: G. A. Khoury, Performance of Heated Concrete—Mechanical Properties, Contract NUC/56/3604A with 

Nuclear Installations Inspectorate, Imperial College, London, United Kingdom, August 1996.
5

information on environmental factors that affect heated concrete and an indication of their relative influence [2.8]. 

Although information in this table indicates that there are several negative effects of elevated temperature exposure, 

there are two potential positive aspects of elevated temperature exposure of concrete [2.9]. First, transient creep, or 

load-induced thermal strain, is much larger than elastic strain and contributes to a significant relaxation that helps to 

minimize stress gradients originating from thermal incompatibilities and temperature gradients in heated concrete 

[2.10]. Second, concrete in service generally is under load, which has a beneficial effect during heating in that it 

confines the concrete and tends to inhibit crack development [2.11]. 

Research has been conducted on the thermal behavior of concrete in connection with the development of prestressed 

concrete pressure vessels for nuclear power plants (i.e., 20 C to 200 C) and to study the behavior of reinforced 

concrete members under fire conditions (i.e., 20 C to 1000 C) [2.12]. Interpretation of these results can be difficult 

however because (1) test materials and curing conditions were different, (2) descriptions of materials are incomplete, 

(3) different test procedures were utilized (e.g., heating rates and exposure times), (4) test conditions are not 

comparable (e.g., tested at temperature or permitted to cool to room temperature prior to testing, and loaded or 

unloaded while heating), and (5) shape and size of the test articles were different (e.g., cube and cylindrical). 

Factors such as these can combine during first heating to influence the properties of concrete (e.g., measurements of 

concrete compressive strength at 150 ̊C have yielded results that can range from as low as 30% to as high as 150% 

the concrete initial cold strength) [2.13]. Figure 2.2 presents a global diagram of dependence of concrete’s thermohydral-mechanical properties dependence on the load-heat-time sequence [Note: influence of moisture boundary 

conditions (i.e., totally sealed, semi-sealed, or unsealed) is not shown; LITS = load-induced thermal strain)] [2.14].

Figure 2.2 Examples of dependence of concrete’s thermo-hydral-mechanical properties 

on heat-load-time sequence of the test regime (moisture boundary conditions not shown). 

Source: G.A. Khoury, Effect of Heat on Concrete, Department of Civil and Environmental Engineering Report, 

Imperial College, London, United Kingdom, 1995.

Concrete structural members are to be designed to satisfy the requirements of serviceability and safety limit states 

for various environmental conditions. Elevated temperature exposure is one of the environmental conditions that 

may need to be addressed for some of the new advanced reactor designs. Contained in the following sections of this 

chapter is a summary of literature that has been identified addressing the general behavior and pertinent mechanical 

and physical properties of concrete materials under elevated-temperature conditions.

2.1 General Behavior

If concrete made with Portland cement or blast furnace slag cement is subjected to heat, a number of transformations 

and reactions occur, even if there is only a moderate increase in temperature [2.15,2.16]. As aggregate materials 
6

normally occupy 60 to 80% of the concrete volume, the behavior of concrete at elevated temperature is strongly 

influenced by the aggregate type. The thermal stability of the concrete depends largely on the thermal stability of 

the aggregate (i.e., thermal strain depends on aggregate used). Commonly used fine and coarse aggregate materials 

are thermally stable up to 300 – 350 C as noted in Figure 2.3 where thermal expansion data for 20 – 30 mm prisms 

of gravel, limestone, and basalt aggregate materials is presented showing residual expansion after cooling. 

Figure 2.3 Thermal strains of aggregate prisms during heat cycle to 600 ̊C at 2 ̊C/min.

Source:

Thermal instability and breakup temperatures reported for a gravel, limestone, and basalt are 300 ̊C+, 400 ̊C+, and 

500 ̊C+, and 350 ̊C+, 600-700 ̊C+, and >600 ̊C, respectively [2.8]. Figure 2.4 presents information on the thermal 

stability and processes that take place on heating in several aggregate types. Aggregate characteristics of importance 

to behavior of concrete at elevated temperature include physical properties (e.g., thermal conductivity and thermal 

expansion), chemical properties (e.g., chemical stability at temperature), and thermal stability/integrity.

Figure 2.4 Examples of range of thermal stabilities and processes that take place in aggregates during heating.

Source: G.A. Khoury, Effect of Heat on Concrete, Department of Civil and Environmental Engineering Report, 

Imperial College, London, United Kingdom, 1995.

Aggregate materials may undergo crystal transformations leading to significant increases in volume [e.g., crystalline 

transformation of -quartz (trigonal) to -quartz (hexagonal) between 500 and 650 C with an accompanying 
7

increase in volume of ~5.7%]. Some siliceous or calcareous aggregates with some water of constitution exhibit 

moderate dehydration with increasing temperature that is accompanied by shrinkage (i.e., opal at 373 C exhibits 

shrinkage of ~13% by volume) [2.17]. Most nonsiliceous aggregates are stable up to about 600 C. At higher 

temperatures, calcareous aggregates (calcite – CaCO3), magnesite (MgCO3), and dolomite (MgCO3/CaCO3) 

dissociate into an oxide and CO2 (CaO + CO2). Calcium carbonate dissociates completely at 1atm pressure at 

898 C with partial dissociation occurring at temperatures as low as 700 C [2.18]. Aggregates produced from 

crystalline rocks may progressively disintegrate at high temperatures because of thermal incompatibility of 

aggregate itself and possible dehydration of any chemically combined moisture that is present. Above 1200 C and 

up to 1300 C some aggregates, such as igneous rocks (e.g., basalt), show degassing and expansion. Lightweight and 

refractory aggregates can be utilized to produce significant improvements in the heat resistance of Portland cement 

concretes. However, aggregate porosity of some lightweight aggregates varies from 0 to 50% or more and any 

water present in the aggregate is evaporable and can be expelled on heating to contribute to pore pressure build up in 

concrete [2.19]. It has been noted that the thermal stability of aggregates increases in order of gravel, limestone, 

basalt, and lightweight [2.20]. Hydrous aggregates (e.g., serpentine, limonite, and goethite) containing chemically 

combined water of crystallization (e.g., 4 -12%, by weight) used in concrete for radiation shielding have dehydration 

temperatures ranging from 175 ̊ to 500 ̊C depending on the aggregate type [2.21]. The descending order of fire 

endurance of aggregates is expanded slags, shales, slates, and clays; air-cooled slag; basic, finely grained igneous 

rocks such as granite and basalt; calcareous; siliceous aggregates; and lastly flint [2.19] 

Apart from the crystalline transformations occurring mainly in the aggregate materials during heating, a number of 

degradation reactions occur, primarily in the cement paste, that result in a progressive breakdown in the structure of 

the concrete. Loss of concrete structural quality (e.g., strength and fracture) under elevated temperature exposure is 

dependent in large measure on alteration of the physico-chemical composition of the hardened cement paste which 

affects its phase composition and pore structure (i.e., specific surface of hydration products reduces and pore 

structure becomes coarser) [2.22,2.23]. An increase in temperature produces significant changes in the chemical 

composition and microstructure of the hardened Portland cement paste. At low temperatures these reactions mainly 

take the form of dehydration and water expulsion reactions. Changes in the chemical composition and 

microstructure of unsealed hardened Portland cement paste occur gradually and continuously over a temperature 

range from room temperature to 1000 C. At room temperature, between 30 and 60% of the volume of saturated 

cement paste and between 2 and 10% of the volume of saturated structural concrete are occupied by evaporable 

water. At temperatures up to 80 ̊C the hydration products of ordinary Portland cement essentially remain chemically 

unaltered and changes in the properties can be attributed to physical effects (e.g., changes in van der Waals cohesive 

forces, porosity, surface, energy, and cracking) [2.14]. As the temperature to which the cement paste is subjected 

increases, evaporable water is driven off until at a temperature of about 105 C all evaporable water will be lost, 

given a sufficient exposure period. At temperatures above 105 C, the strongly absorbed and chemically combined 

water (i.e., water of hydration) are gradually lost from the cement paste hydrates. This represents the dominant 

process affecting performance as temperatures increase above about 100 ̊C [2.9] (Fig. 2.8 presents the influence of 

temperature on compressive strength and modulus of a cement paste). Dehydration of the calcium hydroxide is 

essentially zero up to about 400 C, increases most rapidly around 535 C, and becomes complete at about 600 C 

[2.24], and thus impacts concrete performance at higher temperatures. This is illustrated in Figure 2.5 where 

multiple x-ray diffraction patterns are presented for samples removed from the surface and interior of mortar 

specimens after exposure to temperatures up to 1000 ̊C (i.e., Ca(OH)2 reduces from 6.01% at ambient to 2.52% at 

500 ̊C to nil at 800 ̊C)* [2.25]. The calcium hydroxide and calcium carbonate results correspond to the diffraction 

lines of d = 4.90 Å and d = 3.03 Å, respectively. Results for Ca(OH)2 show a gradual reduction with increase in 

temperature indicating a minor effect on concrete performance at lower temperatures where loss of strongly 

absorbed and chemically combined water dominates, but has increasing importance at temperatures where 

decarbonation occurs. Results for Ca(CO)3 indicate negligible carbonation of the sample has occurred.

 

Relatively immature sealed cement paste tends to accelerate hydration and improve strength on heating, but at 

temperatures above 80 ̊ to 100 ̊C any hydration of cement paste could be beneficial or detrimental to strengthening 

depending on the CaO/SiO2 ratio of the gel [2.19]. When heated above 80 ̊ to 100 ̊C in sealed condition 

hydrothermal reactions occur that result in significant changes in its chemical/physical microstructure.

**

It has been noted that Ca(OH)2 content change with depth may be used as an indicator of exposure temperature at 

the concrete surface as well as depth of damage [2.26].
8

Figure 2.5 X-ray diffraction results for mortar samples of concrete monitoring 

Ca(OH)2 and Ca(CO)3 fired at different temperatures.

Source: S.K. Handoo, S. Agarwal, and S.K. Agarwal, ―Physiochemical, Mineralogical, and Morphological 

Characteristics of Concrete Exposed to Elevated temperatures,‖ Cement and Concrete Research 32(7), 

pp. 1009-1018, 2002.

The behavior of unsealed cement paste during, and after, cooling is dominated by absorption of moisture from the 

surrounding medium [2.19]. Changes that take place during heating are somewhat irreversible, and rewetting of the 

cement paste after cooling can lead to some restoration of bonds partly due to rehydration. The CaO produced 

during previous heating from the dehydration of Ca(OH)2 (above 400 ̊C) and decarbonation of the CaCO3 (above 

700 ̊C) absorbs water on cooling and rehydrates into Ca(OH)2 with a corresponding 44% increase in volume that can 

lead to development of cracks and a weakening of the paste. It has been noted that 400 ̊C is a critical temperature 

for breakdown of the ordinary Portland cement (OPC) pastes due to dehydration of the calcium hydroxide followed 

by expansive rehydration of lime after cooling that occurs with time that can lead to cracking [2.13, 2.27]. The 

potential damaging effect of rehydration is illustrated by a study conducted using an ordinary Portland cement paste 

and an ordinary Portland cement/blast furnace slag paste to investigate the short- and long-term effects of 

rehydration of lime [2.28]. Three combinations of OPC/ground granulated blast furnace slag materials were studied 

[i.e., 30, 50, and 65% (by weight) replacement of OPC with slag]. Cylindrical specimens 50-mm diameter by 

100-mm long were demolded 24 hours after casting and cured for 14 days in lime-saturated water at 23 ̊C followed 

by oven drying for 2 hours at 60 ̊C. The specimens were subjected to 800 ̊C for 1 hour, permitted to cool to room 

temperature, and placed into sealed plastic bags until testing. Two different ages of paste specimens were 

investigated: one set of OPC and OPC blast furnace slag paste specimens was examined 1 week after thermal 

exposure and a second set 1 year after thermal exposure. Figure 2.6 presents the evolution of rehydration in the 

OPC paste specimens for times representing immediately after heating to 800 ̊C and 2 days after heating. The OPC 

paste specimens exhibited a large quantity of visible cracks on removal from the furnace and essentially completely 

disintegrated 2 days after heating. The OPC/blast furnace slag paste specimens exhibited few to no visible cracks 

upon removal from the furnace and 2 days after removal indicating that replacement of OPC with slag can be 

beneficial with respect to reducing the harmful effects of rehydration of CaO when exposed to temperatures above 

400 ̊C. After 1 year the OPC pastes had disintegrated and thus exhibited no residual compressive strength while the 

OPC/slag specimens exhibited similar behavior to the 2 day results with few visible cracks and no visible distinction 

between 2 day and 1 year results while maintaining similar residual compressive strength. Differential 

thermogravimetric analysis results for unheated OPC pastes (reference specimens) showed initial weight loss 

between 100 ̊ and 200 ̊C (loss of capillary water), secondary weight loss between 400 ̊ and 500 ̊C [dehydration of 

Ca(OH)2], and tertiary weight loss at temperatures above 750 ̊C [decarbonation of Ca(CO)3]. Samples of OPC that 

had been heated did not exhibit initial weight loss, secondary weight loss occurred between 400 ̊ and 500 ̊C 

indicating that CaO rehydration took place with Ca(OH)2 amounts being greater than observed for the reference 

specimens indicating CaO rehydration results from dissociation of Ca(OH)2 and from decarbonation of Ca(CO)3, 

and tertiary weight loss above 750 ̊C did not occur. Differential thermogravimetric results for the 50% slag 
9

specimens exhibited similar behavior to the reference OPC paste specimens. For the specimens tested after 1 week, 

the peak representing dehydration of the Ca(OH)2 was significantly greater for the OPC pastes than the OPC/slag 

pastes indicating rehydration of CaO is greater in OPC pastes. For specimens tested after 1 year, results indicated 

that rehydration of the CaO throughout 1 year had a more significant effect in OPC pastes than OPC/slag pastes. 

The study results concluded that OPC pastes are more affected by the long-term effect of CaO rehydration than 

OPC/slag pastes.

Figure 2.6 OPC paste specimens immediately after and two days after 800 ̊C heat treatment.

Source: A. Mendes, J. Sanjayan, and F. Collins,, ―Long-Term Progressive Deterioration Following Fire Exposure 

of OPC Versus Slag Blended Cement Pastes,‖ Materials and Structures 42, pp. 95-101, 2009.

Key material features of Portland cement paste that influence properties of concrete at high temperature are its 

moisture state (i.e., sealed or unsealed), chemical structure (i.e., loss of chemically-combined water from C-S-H in 

unsealed condition, CaO/SiO2 ratio of hydrate in sealed condition, and amount Ca(OH)2 crystals in sealed and 

unsealed condition), and physical structure (i.e., total pore volume, average pore size, and amorphous/crystalline 

structure of solid) [2.8]. Figure 2.7 provides thermal strain data for Portland cement pastes fabricated from ordinary 

Portland cement, ordinary Portland cement and fly ash, and sulfate-resisting Portland cement demonstrating the

Figure 2.7 Thermal strain during first heat cycle to 600 ̊C of initially moist cementitious materials.

Source: G. A. Khoury, Performance of Heated Concrete—Mechanical Properties, Contract NUC/56/3604A with 

Nuclear Installations Inspectorate, Imperial College, London, August 1996.
10

influence of type of Portland cement as well as partial cement replacement. Figure 2.8 indicates the influence of 

temperature on the ultimate compressive strength and modulus of elasticity of a Portland cement paste (Type I 

Portland cement; water-cement = 0.33) [2.29]. Figure 2.9 presents a comparison of the relative strength of Portland 

cement paste (w/c = 0.3) tested either at temperature or after cooling to room temperature prior to testing [2.30]. 

Initial drop in hot strength was attributed to thermally energized swelling of physically bound water layers causing 

disjoining pressures, with regain of strength resulting from relief of these pressures due to drying giving rise to 

greater van der Waal’s forces as cement gel layers move closer together [2.30]. Scanning electron microscopy 

Figure 2.8 Ultimate compressive strength and modulus of elasticity of

hydrated Portland cement paste (w/c = 0.3) at elevated temperature.

Source: T.Z. Harmathy and J.E. Berndt, ―Hydrated Portland Cement and Lightweight Concrete at Elevated 

Temperature,‖ J. American Concrete Institute 63, pp. 93-112, January 1966.

Figure 2.9 Residual and hot strength of Portland cement paste as a function of temperature.

Source: W.P.S. Dias, G.A. Khoury, and P.J.E. Sullivan, ―Mechanical Properties of Hardened Cement Paste 

Exposed to Temperatures Up To 700 ̊C,‖ ACI Materials Journal 87(2), March-April 1990.
11

results indicate that microcracking increases significantly beyond 300 ̊C, first around Ca(OH)2 crystals and then

around grains of unhydrated cement [2.31]. Figure 2.10 compares the effect of temperature on residual strength of 

ordinary Portland cement paste (100% OPC) and blends of cement (90% OPC + 10% Silica fume, 70% OPC + 30% 

pulverized fly ash, and 35% OPC + 65% slag) [2.32]. The 100% OPC exhibited a decline in residual compressive 

strength at temperatures above 200 ̊C with the decline significant at temperatures above 300 ̊C. Results in which 

silica fume replaced 10% of the OPC were similar to those obtained from the 100% OPC specimens. However, the 

specimens in which slag and pulverized fly ash were used as partial replacements for OPC retained their strength 

much better than either the specimens with 100% OPC or silica fume as a partial replacement for OPC. The reason 

for this was attributed to lower Ca(OH)2 contents of these specimens [2.33]. It was further noted that cement paste 

and concrete containing 100% OPC tend to lose significant residual compressive strength above 300 ̊C, but the loss 

is not quite as significant if the specimens are tested at temperature. Figures 2.11 and 2.12 provide a comparison of

Figure 2.10 Effect of temperature on residual compressive strength of cement paste materials.

Source:

residual compressive and flexural strength results, respectively, for a plain Portland cement paste and a Portland 

cement paste containing polypropylene fibers (~0.08%, by volume) after elevated temperature exposures up to 

1000 ̊C (700 ̊C for fiber-reinforced paste) [2.34]. Water-cement ratio was 0.32 for both materials. As noted in the 

figure, the temperature range from 30 ̊ to 120 ̊C was characterized by strength losses, but after temperature 

exposures at 150 ̊ and 440 ̊C the residual compressive strengths equaled the reference strength at 20 ̊C. After higher 

exposure temperatures from 520 ̊ to 1000 ̊C the residual compressive strength of the plain paste decreased. 

Incorporation of the polypropylene fibers decreased both the reference and residual compressive strengths. The 

reference flexural strengths at 20 ̊C were identical in the test specimen with and without the polypropylene fibers. 

At a moderate temperature exposure of 50 ̊C, the residual flexural strength of the plain cement paste decreased 

significantly to about 47% its reference strength and at 120 ̊C it further reduced to about 26% its reference strength. 

Then in the temperature range from 150 ̊ to 440 ̊C the residual strengths increased before decreasing as the exposure 

temperature increased. Addition of polypropylene fibers improved residual flexural strength at low temperatures.

The aggregate-cement paste bond region has been shown to be the weakest link because it is normally weaker that 

the cement paste which is normally weaker than the aggregate [2.35]. The reactivity of the aggregate can be 

beneficial to bond strength or it can be harmful depending on the nature of the chemical reaction that takes place. 

Studies of bond strength at ambient temperatures have shown that, everything else being equal, variations in bond 

strength of up to 300% or more can occur depending on preconditioning of the aggregate (e.g., lubricating or 

polymer coating) [ 2.36,2.37]. If the aggregate-cement paste bond fails on heating, chemically or as a result of 

thermal incompatibility between the aggregate and cement paste, the concrete will exhibit a significant reduction in 

strength, even if both the aggregate and surrounding mortar matrix remain intact [2.19]. Interaction between the 

aggregate and cement paste can be physical or chemical. Physical interaction results from the differential thermal 

expansion between the aggregate and cement paste leading to weakening and disruption of the concrete at elevated 
12

temperature (i.e., on heating to 80 ̊ to 100 ̊C the cement paste expands slightly as water escapes and the aggregate 

expands, while at higher temperatures the cement paste continues to shrink and the aggregate expands leading to a 

large differential strain). Chemical interaction, as influenced by temperature, can occur in the form of the reaction 

between Ca(OH)2 crystals released by hydration of ordinary Portland cement and magnesium carbonate component 

of some limestone aggregates that is expansive leading to weakening and disruption of the concrete [2.19].

In general, during first heating concrete is a highly unstable and complex material. A good summary of the 

degradation reactions that occur in Portland cement concrete is available [2.12] and Figure 2.13 presents an 

illustration of the deterioration process during heating and cooling phases [2.38]. Changes in the microstructure of 

concrete occur as temperature rises. The thermal-hydral-mechanical properties also change with time and 

temperature depending upon factors such as the heating rate, initial moisture condition, boundary conditions, 

geometry and size of the heated member, loading condition, type of constituents, and chemico-physico interactions 

[2.26]. Upon first heating, substantial water evaporation occurs from the larger pores close to the concrete surface. 

Then, from 100 ̊C onward, the evaporation proceeds at a faster rate with water being expelled from the concrete near 

the surface as a result of above-atmospheric vapor pressure (i.e, steam flow). At 120 ̊C the expulsion of water 

physically bound in the smaller pores, or chemically combined, initiates and continues up to about 500 ̊C where the

Figure 2.11 Residual compressive strengths of cement paste and cement paste with polypropylene fibers.

Source: J. Komonen and V. Penttala, ―Effects of High Temperature on the Pore Structure and Strength of Plain and 

Polypropylene Fiber Reinforced Cement Pastes,‖ Fire Technology 39(1), pp. 23-34, 2003. 

Figure 2.12 Residual flexural strengths of cement paste and cement paste with polypropylene fibers.

Source: J. Komonen and V. Penttala, ―Effects of High Temperature on the Pore Structure and Strength of Plain and 

Polypropylene Fiber Reinforced Cement Pastes,‖ Fire Technology 39(1), pp. 23-34, 2003.
13

process is essentially complete. From 20 ̊ to 300 ̊C, in conjunction with evaporation, dehydration of the hardened 

cement paste occurs (first stage) with the maximum rate of dehydration occurring at about 180 ̊C {Tobermorite gel 

is stable up to a temperature of 150 ̊C [2.39]}. In the temperature range from 450 ̊ to 550 ̊C there is decomposition 

of the Portlandite {i.e., Ca(OH)2 CaO + H2O [2.36]}. At 570 ̊C the inversion of quartz takes place with 

the transformation being endothermic and reversible. A further process of decomposition of the hardened cement 

paste takes place between 600 C and 700 C with the decomposition of the calcium-silicate-hydrate phases and 

formation of -C2S. Between 600 C and 900 C the limestone begins to undergo decarbonation (i.e., CaCO3

CaO + CO2). The rate of decomposition and the temperature at which it occurs are not only dependent on 

temperature and pressure, but also on the content of SiO2 present in the limestone. Above 1200 C and up to 

1300 C, some components of the concrete begin to melt. Above 1300 C to 1400 C concrete exists in the form of a 

melt.

Figure 2.13 Schematic of concrete deterioration under moderate heating and cooling.

Source: Adaptation of figure in K.D. Hertz, ―Concrete Strength for Fire Safety Design,‖ Magazine of Concrete 

Research 57(8), pp. 445-453, 2005.

Apparently liquefaction of the concrete commences with melting of the hardened cement paste followed by melting 

of the aggregates [2.40-2.42]. The melting points of aggregates vary greatly. At 1060 C basalt is at the lower limit 

of all types of rock, with quartzite not melting below 1700 C [2.15]. Figure 2.14 presents scanning electron 

micrographs showing the changes in concrete morphology as temperature increases from ambient to 800 ̊C (i.e, well 

developed hydrated phases of Ca(OH)2 intermixed with calcium-silicate-hydrate and calcium aluminate at ambient, 

and presence of microcracks, voids increasing the porosity of deformed Ca(OH)2 crystals, and disrupted calciumsilicate-hydrate phase boundaries at 800 ̊C). Figure 2.15 presents differential thermal analysis results for normal 

weight concrete indicating effects of elevated temperature exposure.
14

Figure 2.14 Change in morphology of hardened concrete after exposure to elevated temperature.

Source: S.K. Handoo, S. Agarwal, and S.K. Agarwal, ―Physiochemical, Mineralogical, and Porphological 

Characteristics of Concrete Exposed to Elevated temperatures,‖ Cement and Concrete Research 32(7), pp. 

1009-1018, 2002.

Figure 2.15 Differential thermal analysis of selected normal weight concretes.

Source:
15

2.2 Mechanical and Physical Properties 

Material properties are closely related to the specific test method employed. The properties of concrete at elevated 

temperature can be defined from a number of viewpoints (e.g., ranging from transient such as representing fire 

conditions to steady-state such as a structure operating at elevated temperature). To interpret results, knowledge of 

the test condition(s) employed is required. Three main test parameters are involved in the development of data: 

heating, application of load, and control of strain [2.43]. These parameters can be fixed at constant values or be 

varied during testing to provide transient conditions. Six regimes or ―idealized‖ testing methods have been 

identified for determining properties of concrete at elevated temperature [2.12,2.43,2.44]:

Each of these methods or regimes determines a specific feature of material behavior. Regimes 1 through 4 are 

related to steady-state tests and regimes 5 and 6 transient tests. 

Figure 2.16 provides a schematic representation of testing procedures commonly used to evaluate concrete response 

to elevated temperature: stressed tests, unstressed tests, and unstressed residual. Stressed tests are a modified 

version of stress- or strain-controlled experiments performed under isothermal temperature conditions. A preload, 

generally in the range of 20 to 40 percent of the ultimate compression strength at room temperature, is applied to the 

concrete specimen prior to heating, and the load is sustained during the heating period. After the specimen reaches a 

steady-state temperature condition, the stress or the strain is increased at a prescribed loading rate until the specimen 

fails. Unstressed tests are carried out identically to the stress- or strain-controlled experiments of the steady-state 

type. Unstressed residual strength tests are experiments where the specimen is first cooled to room temperature after 

one or several cycles of heating without preloading. The load is then applied at room temperature under stress or 

strain control until the specimen fails. 

2.2.1 Mechanical Properties

It has been established that the mechanical properties of concrete can be adversely affected by elevated-temperature 

exposure [2.44-2.48]. Deterioration of concrete’s mechanical properties can be attributed to three material factors: 

(1) physicochemical changes in the cement paste, (2) physicochemical changes in the aggregate, and (3) thermal 

incompatibility between the aggregate and the cement paste; and the properties are influenced by environmental 

factors such as temperature level, heating rate, applied loading, and external sealing influencing moisture loss [2.9]. 

Several studies have been conducted to evaluate these factors, however, quantitative interpretation of available data is 

difficult because (1) samples were either tested hot or cold, (2) moisture migration was either free or restricted,
16

Figure 2.16 Schematic of temperature and loading histories.

Source: K. Willam, Y. Xi, K. Lee, J. Lee, and H. Basche, ―Confirmatory Analytical Research on Phenomena 

Associated with Structural Response to Aircraft Attack – Letter Report,‖ Appendix A – A Literature Review 

of Concrete Structures Subjected to High Temperatures, University of Colorado, Boulder, January 6, 2005.

(3) concrete was either loaded or unloaded while heated, (4) mix constituents and proportions varied, (5) test 

specimen size and shape were not consistent, (6) specimens were tested at different degrees of hydration, and 

(7) heat-soak duration varied from test to test. To provide a consistent basis for evaluation of data, it is 

recommended that several factors be taken into account [2.49]: (1) concrete strength class; (2) test specimen size; 

(3) thermal compatibility of aggregate and cement paste matrix; (4) cement and concrete composition; (5) level of 

temperature; (6) degree of hydration; (7) moisture content; (8) moisture gradients, rate of drying or wetting; 

(9) temperature gradient, rate of heating or cooling; (10) duration of temperature exposure; (11) loading during 

temperature exposure; (12) temperature-activated transformations in microstructure and chemical composition of 

cement; (13) state of specimens tested—hot or cold; (14) strength testing procedure; and (15) reference strength 

selected—wet, moist, or dry. Therefore great attention should be paid by the researcher who produces the data in 

unraveling and understanding the complexities in order to present the subject to the practicing engineer in whatever 

form (e.g., texts, diagrams, tables, and numerical modeling input) that are readily usable, as close to reality as 

possible, and as reliable as possible [2.26].

A review of methods used by various investigators for elevated-temperature testing of concrete indicates that, 

generally, the tests can be categorized according to cold or hot testing. In cold testing, specimens are gradually 

heated to a specified temperature, permitted to thermally stabilize at that temperature for a prescribed period of time, 

permitted to slowly cool to ambient, and then tested to determine residual mechanical properties. In hot testing, 

specimens are gradually heated to a specified temperature, permitted to thermally stabilize at the temperature for a 

prescribed period of time, and then tested at temperature to determine relative mechanical properties. During testing, 

specimens are maintained in either an open environment where water vapor can escape (unsealed) or a closed 

environment where the moisture is contained (sealed). The closed environment represents conditions for mass 

concrete where moisture does not have ready access to the atmosphere, and the open environment represents 

conditions where the element is either vented or has free atmospheric communication. During heating and cooling, 

the specimens may be either loaded or unloaded. Mechanical properties in which the specimens have been 

permitted to return to room temperature prior to testing are referred to as residual properties, and when properties are 

obtained from specimens tested at temperature they are referred to as relative properties.
17

The performance of concrete can be measured by the change of its stiffness, strength, or some other property that 

would affect its main function in service. Because concrete has a relatively low tensile strength, it is normally relied 

upon to take compressive forces, with tensile forces taken by steel reinforcement. As a consequence, much of the 

research conducted on concrete at elevated temperature has concentrated on compressive strength as the 

fundamental property in examining its deterioration. However, it has been noted that the compressive strength may 

not be as good an indicator of deterioration at elevated temperature as tensile or flexural strength under short-term 

loading [2.50]. 

2.2.1.1 Stress and Strain Characteristics

Information and Data. Evaluation of structures for small strain conditions involves elastic analysis procedures for 

which knowledge of the concrete modulus of elasticity and strength is sufficient. When large strains are involved, 

such as could occur when a structure is subjected to elevated temperature, elastic-plastic analysis procedures are 

required that involve use of the load-deformation or stress-strain relations developed for concrete at the temperature 

level of interest. The stress-strain relationship is one of the basic mechanical properties required for prediction of 

the overall response of concrete structures under thermal loadings [2.51]. Elevated temperature is one of the main 

factors that can have a significant effect on the concrete stress-strain relationship.

Uniaxial Compression Testing

The majority of stress-strain data reported in the literature are for concrete heated to test conditions without load or 

loaded under stress-controlled conditions. Figure 2.17 presents stress-strain diagrams for sealed and unsealed 

limestone aggregate concretes tested at temperature [2.52]. These results indicate that the unsealed specimens are 

stiffer than the sealed specimens, but strains at ultimate load were reduced. Figure 2.18 presents the influence of test 

temperature on the stress-strain relationship of a quartz aggregate concrete in a stress-rate controlled test conducted 

at test temperature [2.53] and a sandstone aggregate concrete with testing performed after specimens cooled to room 

temperature [2.54]. These data show a significant increase in ultimate strain and a loss of stiffness with increasing 

temperature. Figure 2.19 shows that specimens made from quartz aggregate concrete that are tested at temperature 

are stiffer and stronger than identical companion specimens heated to the same temperatures and then permitted to 

cool to room temperature before testing (i.e., up to 450°C the stress-strain curves of specimens tested at temperature 

do not change appreciably) [2.55]. It was also concluded from this study that the type of cement and the duration of 

thermal treatment had a minor effect on the slope of the stress-strain curve. Figure 2.20 presents results for basalt 

aggregate concrete derived under strain-rate controlled test conditions [2.56]. Figure 2.21 provides stress-strain results for 

a lightweight masonry concrete tested at several temperatures. Results show that as temperature increases the 

ultimate stress decreases and the ultimate strain increases.

Figure 2.17 Stress-strain diagrams for sealed and unsealed limestone concrete.

Source: R. Kottas, J. Seeberger, and H.K. Hilsdorf, ―Strength Characteristics of Concrete in Temperature Range of 

20 ̊ to 200 ̊C,‖ Paper H01/4 in 5

th International Conference on Structural Mechanics in Reactor 

Technology, Elsevier Publishers, North-Holland, The Netherlands, August 1979.
18

Figure 2.18 Influence of temperature on the stress-strain relation

of unsealed quartz aggregate and sandstone aggregate concrete.

Sources: (a) Y. Anderberg and S. Thelanderson, ―Stress and Deformation Characteristics of Concrete at High 

Temperatures, 2-Experimental Investigation and Material Behaviour Model,‖ Bulletin 54, Lund 

Institute of Technology, Lund, Sweden, 1976.

(b) T. Harada, T. Takeda, S, Yamane, and F, Furumura, ―Strength, Elasticity and Thermal Properties of 

Concrete Subjected to Elevated Temperatures,‖ Paper SP-34-21 in Concrete for Nuclear Reactors, 

pp. 377-406, American Concrete Institute, Farmington Hills, Michigan, 1972.

Currently, utilization of high strength (e.g., fc

’

> 55 MPa) or high performance concretes (e.g., improved durability) 

has become an option to the use of traditional normal strength concrete. This is due to significant economic,

Figure 2.19 Effect of elevated temperature on stress-strain behavior of a quartz concrete.

Source: H. Weigler and R. Fischer, ―Influence of High Temperatures on Strength and Deformations of Concrete,‖ Paper 

SP 34-26 in Special Publication SP-34, Vol. I–III, American Concrete Institute, 1972.

 

Technical Committee 363, ―High-Strength Concrete,‖ Seventh International Symposium on Utilization of High Strength/High 

Performance Concrete, ACI Special Publication 228, pp. 79-80, American Concrete Institute, Farmington Hills, Michigan, June 

1, 2005. Does not include concretes produced using exotic materials or techniques.
19

Figure 2.20 Stress-strain relationship for basalt concrete derived in strain-rate controlled tests.

Source: U. Schneider, ―Concrete at High Temperatures – A General Review,‖ Fire Safety Journal 13, pp. 55-68, 

1988.

Figure 2.21 Stress-strain curves for expanded shale lightweight masonry concrete tested at temperature.

Source: M.S. Abrams, ―Behavior of Inorganic Materials in Fire,‖ Design of Buildings for Fire Safety, ASTM 

STP 685, American Society for Testing and Materials, West Conshohocken, Pennsylvania, 1979.

architectural, and structural advantages that high-strength concrete can provide relative to normal-strength concrete 

(e.g., high-strength concrete used for primary load-bearing members such as structural framing consisting of beams 

and columns) [2.57]. The improved strength and performance is obtained by reducing the amount of water in the 

mix and use of chemical or mineral admixtures to reduce permeability and impart improved workability. However, 

with the use of lower water-cement ratios and admixtures the concretes become more brittle and exhibit reduced 

permeability that can make the concretes more susceptible to spalling under rapid thermal loadings. Stress-strain 

results for a high-strength (91.8 MPa) and a normal-strength (32.9 MPa) concrete tested at temperature are presented in 

Figure 2.22 and indicate that the high-strength concrete has steeper and more linear stress-strain curves and tends to fail in a 

more brittle manner than normal-strength concrete. [2.58]. In order to address the potential limitations of high-strength 

concretes under thermal loadings, polypropylene or steel fibers have been incorporated into the mixes. Additions of
20

Figure 2.22 Comparison of high-strength and normal-strength concrete load-deformation curves.

Source: C. Castillo and A.J. Durrani, ―Effect of Transient High Temperature on High-Strength Concrete,‖ ACI 

Materials Journal 87(1), American Concrete Institute, pp. 47-53, January-February 1990.

relatively small amounts (e.g., 0.3%, by volume) of polypropylene fibers to the concrete mix improves the resistance 

to spalling as the polypropylene fibers melt at around 165 ̊C to relieve the vapor pressure [2.59]. Steel fiber 

additions (e.g., 1 to 2%, by volume) enhance the mechanical behavior of high-strength concrete at elevated 

temperature and significantly improve the concrete ductility [2.60]. Figure 2.23 presents a comparison of stressstrain curves for specimens containing polypropylene fibers with several concretes tested at room temperature 

Figure 2.23 Stress-strain relationships for several concrete types before and after exposure to 200 ̊C.

Source: A. Noumowé and C. Galle, ―Study of High Strength Concretes at Raised Temperature up to 200 ̊C: 

Thermal Gradient and Mechanical Behaviour,‖ Transactions of the 16th Structural Mechanics in Reactor 

Technology, Paper #1580, Washington, DC, August 2001.

(reference) and after exposure for 7 days at 200 ̊C (residual) [2.61]. Types of concrete tested included control high 

strength concrete, concrete incorporating polypropylene fibers (0.22%, by volume), lightweight aggregate concrete 

(expanded clay), and ordinary concrete. Reference compressive strengths for the concretes were 70, 72, 45, and 

48.5 MPa, respectively. Data indicate that the results for the mix with polypropylene fiber and the reference high 

strength concrete mix were similar, addition of lightweight aggregate reduced the slope of the stress-strain curve but 

improved the heat resistance, and ordinary concrete endured the effect of temperature better than high strength and 

lightweight concretes. Figures 2.24 and 2.25 present examples of stress-strain curves obtained at temperature for 

high-strength siliceous (granite) and carbonate (limestone) aggregate concretes, respectively, with and without steel 

fiber additions. Ninety-one-day compressive strengths of all mixes were in the range of 78 to 86 MPa. Highstrength fiber-reinforced mixes contained about 0.5% steel fibers by volume. Results indicate that aggregate type

has an effect on ultimate strain under high temperature conditions (i.e., carbonate aggregate concrete strain at peak 

strength was up to 40% greater than that for siliceous aggregate concrete) and addition of steel fibers increased 

ductility at elevated temperature of both concretes. 
21

Figure 2.24 Stress-strain curves for high-strength siliceous aggregate concrete

without and with steel fiber reinforcement.

Source: F-P. Cheng, V.K.R. Kodur, and T-C. Wang, ―Stress-Strain Curves for High Strength Concrete at Elevated 

Temperatures,‖ Report NRCC-46973, Institute for Research in Construction, National Research Council 

Canada, March 15, 2004 (http://irc.nrc-cnrc.gc.ca/ircpubs).

Figure 2.25 Stress-strain curves for high-strength carbonate aggregate concrete 

without and with steel fiber reinforcement.

Source: F-P. Cheng, V.K.R. Kodur, and T-C. Wang, ―Stress-Strain Curves for High Strength Concrete at Elevated 

Temperatures,‖ Report NRCC-46973, Institute for Research in Construction, National Research Council 

Canada, March 15, 2004 (http://irc.nrc-cnrc.gc.ca/ircpubs).

Limited data are available on the effect of strain rate on concrete stress-strain characteristics at temperatures of 20 ̊, 

65 ̊, and 90 ̊C [2.62]. Nominal concrete compressive strengths were either 30 or 35 MPa after 60 days. Specimens 

10-cm diam by 20-cm long were cured for five weeks in a water bath followed by five weeks at either 65 ̊ or 90 ̊C 

prior to testing. A comparison of compressive stress-strain curves at 20 ̊ and 65 ̊C for strain rates from 10-5/sec to 

101/sec is presented in Figure 2.26. Results indicate that initial elastic modulus is degraded as a result of elevated 

temperature exposure and increases with strain rate gradually approaching an upper shelf.
22

Figure 2.26 Comparison of compressive stress-strain curves for different strain rates at 20 ̊C and 65 ̊C.

Source: T. Yagashita, K. Shirai, C. Ito, and H. Shiojiri, ―Effect of Strain Rate on Concrete Strength Under High 

Temperature,‖ International Conference on Structures Under Shock & Impact VI, pp. 539-548, Cambridge, 

United Kingdom, WIT Press, July 2000.

Many nuclear power plant structures, such as prestressed concrete pressure vessels, will be under a compressive load 

prior to heating. Figure 2.27 demonstrates the beneficial effect of applied preload load (0, 10, or 30% the reference 

strength) during exposure to temperatures of either 250 C or 450 C on the normalized stress-strain curves [2.12]. 

Figure 2.27 Stress-strain relationship of normal concrete with specimens loaded during the heating period.

Source: U. Schneider, Behaviour of Concrete at High Temperature, HEFT 337, Deutscher Ausschuss fur 

Stahlbeton, Wilhelm Ernst & Sohn, Munich, Germany, 1982.

Figure 2.28 provides data on the effect of preloading during heating on the concrete compressive stress-strain 

behavior for a 70 MPa ordinary Portland cement concrete and Figure 2.29 presents results for an ultra highperformance concrete (> 100 MPa) [2.63]. The unsealed specimens were tested at temperatures from room 

temperature to 500 ̊C (700 ̊C for ultra high-performance concrete) while either unloaded or preloaded to 20% of the 

specimen’s compressive strength. Results indicate that the elastic modulus decreased as the test temperature 

increased for the specimens not preloaded and that the modulus of the specimens under sustained loading was larger 

than the modulus of the specimens without sustained loading at the same test temperature. The extent of linearity of 

the stress-strain curve for the ultra high-performance concrete was greater than that for the lower strength concrete 

(i.e., 80% of maximum stress versus 50%).
23

Figure 2.28 The effects of temperature and preloading on unsealed specimens 

tested at temperature: 70 MPa concrete.

Source: G.A. Khoury, C.E. Majorana, F. Pesavento, and B.A. Schrefler, ―Modelling of Heated Concrete,‖ 

Magazine of Concrete Research 54(02), pp. 77-101, April 2002.

Figure 2.29 The effects of temperature and preloading on unsealed specimens 

tested at temperature: ultra high-performance concrete.

Source: G.A. Khoury, C.E. Majorana, F. Pesavento, and B.A. Schrefler, ―Modelling of Heated Concrete,‖ 

Magazine of Concrete Research 54(02), pp. 77-101, April 2002.

Additional stress-strain results for specimens either unstressed or stressed (40% preload applied) during heating are 

presented in Figures 2.30 and 2.31, respectively [2.64]. Four concretes were investigated: normal-strength ordinary
24

Figure 2.30 Temperature-dependent stress-strain curves in

unstressed condition for specimens tested at temperature.

Source: Y.F. Fu, Y.L. Wong, C.S. Poon, and C.A. Tang, ―Stress-Strain Behaviour of High-Strength Concrete at

Elevated Temperature,‖ Magazine of Concrete Research 57(9), pp. 535-544, 2005.

Figure 2.31 Temperature-dependent stress-strain curves in 

stressed condition for specimens tested at temperature.

Source: Y.F. Fu, Y.L. Wong, C.S. Poon, and C.A. Tang, ―Stress-Strain Behaviour of High-Strength Concrete at 

Elevated Temperature,‖ Magazine of Concrete Research 57(9), pp. 535-544, 2005.
25

Portland cement concrete, high-strength ordinary Portland cement concrete, ordinary Portland cement concrete in 

which 30% of cement was replaced by fly ash, and ordinary Portland cement concrete in which 5% of cement was 

replace by metakaolin. Granite aggregate having a 20-mm-maximum size was used in all the mixes. Specimens 

were permitted to thermally stabilize 60 minutes prior to testing. For the unstressed condition (Figure 2.30) the 

ascending branch of the curves consists of two components: (1) an elastic recoverable strain component, that is 

temperature dependent and is affected principally by the thermal load level during initial heating to the test 

temperature; and (2) an irrecoverable plastic strain component. For the normal strength concrete the ascending 

branch remained linear till about 40% peak strength while it was linear till about 60% for the high-strength concretes 

indicating that irrecoverable components of strain were small. Above 400 ̊C the stress-strain curves became 

increasingly nonlinear as the temperature increased and the peak strains corresponding to peak strength increased by 

up to a factor of two. For the stressed condition (Figure 2.31) the ascending portions of the stress-strain curves were 

significantly different than for the unstressed tests with the curves tending to be closer to the room temperature 

curves for each of the concretes. Nonlinearity increased with temperature but there were no significant increases in 

peak strain at peak strength. Strain values at peak strength were all less than those obtained for the unstressed 

condition. Mineral additions were noted to have improved the elevated temperature performance of the concrete.

The effect of cyclic load at elevated temperature and duration of heating has been investigated [2.65]. Specimens 

utilized were 15.24-cm-diam by 45.7-cm-long cylinders that were sealed in copper jackets after casting (Figure 

2.32) and cured for 90-days prior to testing. Type II Portland cement and limestone fine and coarse aggregate 

(19.05 mm maximum size aggregate) were utilized in the concrete mix that had an average 28-d compressive 

strength of 44.7 MPa. Strains were determined by wire-resistance strain gages positioned as shown in Figure 2.32 

and described elsewhere [2.66]. The effect of several factors on the stress-strain behavior of the concrete was 

investigated: (1) sustained temperature of 149 ̊C, (2) number of thermal cycles from 21 ̊ to 149 ̊ to 21 ̊C, 

(3) influence of permitting moisture to freely escape from concrete during thermal treatment (holes punched in 

copper sealing jackets), and (4) influence of testing concrete at 149 ̊C compared to permitting concrete to cool to 

room temperature prior to testing. Figure 2.33a provides stress-strain results obtained at temperature for specimens 

that were subjected to 149 ̊C for periods up to 25 days and Figure 2.33b provides results indicating the effect of 

Figure 2.32 Test specimen for evaluation of cyclic and sustained thermal loads on concrete stress-strain curve. 

Source: V.V. Bertero and M. Polivka, ―Influence of Thermal Exposures on Mechanical Characteristics of 

Concrete,‖ Paper SP-34-28 in Concrete for Nuclear Reactors, pp. 505-531, American Concrete Institute, 

Farmington Hills, Michigan, 1972.

 

Metakaolin is a manufactured pozzolanic mineral admixture, derived from kaolin clay, that significantly enhances 

many performance characteristics of cement-based mortars, concretes, and related products.
26

Figure 2.33 Effect of sustained temperature and thermal cycling on concrete stress-strain curve.

Source: V.V. Bertero and M. Polivka, ―Influence of Thermal Exposures on Mechanical Characteristics of 

Concrete,‖ Paper SP-34-28 in Concrete for Nuclear Reactors, pp. 505-531, American Concrete Institute, 

Farmington Hills, Michigan, 1972.

Figure 2.34 Effect of moisture condition and temperature on concrete stress-strain curve.

Source: V.V. Bertero and M. Polivka, ―Influence of Thermal Exposures on Mechanical Characteristics of 

Concrete,‖ Paper SP-34-28 in Concrete for Nuclear Reactors, pp. 505-531, American Concrete Institute, 

Farmington Hills, Michigan, 1972.

number of thermal cycles (one to fourteen) from 21 ̊ to 149 ̊ to 21 ̊C. As the duration of thermal conditioning 

increased and the number of thermal cycles increased the slope of the stress-strain curve decreased and in general 

the strain at peak stress decreased. However, if moisture is permitted to freely escape during thermal cycling (five 

cycles) the stress-strain curve is not significantly affected as noted in Figure 2.34a. Overall, specimens from which 

moisture was permitted to escape exhibited little difference in characteristics whether tested at 21 ̊ or 149 ̊C. 

Figure 2.34b indicates that for sealed specimens tested at 149 ̊C the mechanical characteristics of the concrete are 

somewhat inferior to those obtained when tested at 21 ̊C. Main parameters that affected the stress-strain relationship 

at elevated temperature were the moisture condition, the number of thermal cycles, and the temperature at which the 

concrete is tested.

Limited information is available on the effect of cooling regime (e.g, cooling rate) on the stress-strain response of 

concrete. One study has been identified in which stress-strain curves were developed as part of an investigation of 

 

Cooling regimes are of importance for conditions where concrete has been subjected to thermal loading that is 

followed by rapid cooling (e.g., accident condition with water quenching). 
27

strength and stiffness characteristics of concrete subjected to various heating and cooling scenarios [2.67]. A granite 

aggregate ordinary Portland cement concrete having a 28-day reference compressive strength of 21.16 MPa was 

used to fabricate 101.6-mm-diameter by 203.2-mm-long cylindrical test specimens that were cured for eight weeks 

in a fog room (23 ̊C and 93% relative humidity). The specimens were heated in a furnace to 200 ̊, 400 ̊, 600 ̊, or 

800 ̊C with a holding time at temperature of four hours. Following heating, the specimens were cooled slowly 

(1 ̊C/minute), naturally (left in furnace and permitted to cool slowly), or rapidly (e.g., specimen removed from 

furnace and placed into a tank of water initially at 20 ̊C). Stress-strain curves for slow, natural, and water cooling 

are presented in Figure 2.35. The initial slope of the stress-strain curves decreased significantly as the temperature 

increased. The elastic region of the specimens subjected to slow or natural cooling and temperatures of 600 ̊C and 

beyond is nonlinear. For water cooling a similar trend was observed at temperatures of 400 ̊C and beyond. 

Although the peak stress following thermal exposure was less at each temperature investigated for the water-cooled 

specimens, the strain at peak stress at each temperature was greater than that for the specimens cooled either slowly 

or naturally.

Figure 2.35 Effect of cooling regime on stress-strain curves following thermal exposure.

Source: J. Lee, Y. Xi, and K. Willam, ―Concrete Under High Temperature Heating and Cooling,‖ Report SESM 

No.09/2006, Department of Civil, Environmental, and Architectural Engineering, University of Colorado at 

Boulder, 2006.

Uniaxial Tension Testing

Results presenting tensile stress-strain curves for concrete at elevated temperature are extremely limited. 

Figure 2.36 presents the effect of temperature on the tensile stress-strain behavior of a 102 MPa high performance 

concrete containing silica fume that was tested in direct tension. 

Multiaxial Testing

In large structures, such as prestressed concrete pressure vessels, the concrete is stressed either biaxially or triaxially. 

Unfortunately only limited data at elevated temperature exist for concrete loaded under either biaxial or triaxial 

conditions. 
28

The test frame presented in Figure 2.37 was used to conduct biaxial tests of 200 by 200 by 50 mm3gravel (quartz) 

concrete specimens [62.0 N/mm2(cube strength)] [2.68,2.69]. Load was applied to the specimen using four 

servohydraulic jacks and brush loading platens. The tests were conducted at temperature using unsealed specimens.

Figure 2.36 Direct uniaxial tensile stress-strain results for an 

unsealed high performance concrete tested at several temperatures.

Source: G.A. Khoury, C.E. Majorana, F. Pesavento, and B.A. Schrefler, ―Modelling of Heated Concrete,‖ 

Magazine of Concrete Research 54(02), pp. 77-101, April 2002

Figure 2.37 Biaxial test system.

Source: C. Ehm and U. Schneider, ―The High Temperature Behaviour of Concrete Under Biaxial Load,‖ Cement

and Concrete Research 25, pp. 27-34, 1985.
29

Figures 2.38, 2.39, and 2.40 presents reference uniaxial ( 1: 2 = 1:0), biaxial ( 1: 2 = 1:0.4), and biaxial ( 1: 2 = 

1:1), respectively, for test temperatures from 20 ̊ to 600 ̊C. The mechanical strains ( 1, 2, and 3) were found to be 

dependent on the stress level, the stress ratio, and the test temperature. The deformations increased with increasing 

load in all three orthogonal directions and at high stress levels they showed a nonlinear relation in the direction of 

greatest principal stress. The ultimate strains for all three axes shift to greater values with increasing temperature. 

At higher temperatures the stress-strain curves were highly nonlinear with the concrete behavior changing from 

brittle behavior to a soft and plastic behavior. Strains in the unstressed axis ( 3) reached high values at elevated 

temperatures and high stress ratios with a maximum value occurring when the stress ratio was 1.0.

Figure 2.38 Biaxial test results: 1: 2 = 1: 0 (uniaxial).

Source: C. Ehm and U. Schneider, ―The High Temperature Behaviour of Concrete Under Biaxial Load,‖ Cement 

and Concrete Research 25, pp. 27-34, 1985.

Figure 2.39 Biaxial test results: 1: 2 = 1: 0.4.

Source: C. Ehm and U. Schneider, ―The High Temperature Behaviour of Concrete Under Biaxial Load,‖ Cement

and Concrete Research 25, pp. 27-34, 1985.
30

Figure 2.40 Biaxial test results: 1: 2 = 1: 1.

Source: C. Ehm and U. Schneider, ―The High Temperature Behaviour of Concrete Under Biaxial Load,‖ Cement

and Concrete Research 25, pp. 27-34, 1985.

When concrete is under a biaxial tension-compression state of stress the capability of concrete to resist cracking is 

diminished, and both the tensile and compressive strengths are found to decrease rapidly [2.70]. Biaxial tensioncompression tests at elevated temperature (i.e., 20 ̊, 200 ̊, 300 ̊, 500 ̊, and 600 ̊C) have been conducted using five 

stress ratios (i.e., 0,0.1, 0.25, 0.50, and 0.75) [2.70]. Specimens 100 by 100 by 100 mm3 were fabricated using a 

20 mm maximum size crushed limestone aggregate concrete having reference compressive and tensile strengths of 

32.5 and 3.14 MPa, respectively. The specimens were held at temperature for six hours and then permitted to slowly 

cool to room temperature prior to testing. Stress ratios ( 2/ 3) investigated included 0 (uniaxial), -0.1, -0.25, -0.5, 

and -0.75. Figure 2.41 presents stress-strain curves for stress ratios of -0.1, -0.2, and -0.25. After high temperature 

exposure the concrete under tension-compression failed abruptly due to the tensile stress. For every stress ratio the 

initial elastic modulus and the peak stress values of the stress-strain curves in both principal compressive stress 

direction and principal tensile stress direction decreased with an increase in exposure temperature. With increasing 

exposure temperature the strains corresponding to peak stresses decreased, the stress peak inclined toward the strain 

axis, and the stress-strain curves tended to flatten. 

Examples of Stress-Strain Relations for Concrete at Elevated Temperature. Several conceptual models are 

available for the mechanical behavior of concrete at elevated temperature (e.g., Anderberg and Thelandersson 

[2.53]; Schneider [2.56]; Terro [2.71]; Khoury, Majorana, Pesavento, and Schrefler [2.63]; Li and Purkiss [2.72]; 

Gawin, Pesavento, and Schrefler [2.73]; and Khennane and Baker [2.74]). The models were developed in order to 

establish a general methodology for use in finite-element analysis of concrete structures. In most of these 

thermomechanical models the strain imposed is broken into four different types: (1) free thermal strain resulting 

from the change in temperature, (2) creep strain due to the dislocation of microstructures within the material 

(3) transient strain caused by changes in the chemical composition of the concrete, and (4) stress-related strain 

resulting from externally applied forces [2.72]. Frequently the creep strain and transient strain are combined as a 

load-induced thermal strain that at temperatures greater than 100 ̊C depends mainly on temperature rather than time 

[2.75]. As noted by results presented previously, the stress-related strain is a function of temperature, stress, and 

time. A critical review involving several of the stress-strain constitutive models identified above has been 

completed [2.72]. Formulations identified in the literature that have been used to calculate parameters that affect the 

stress-strain relationships include:

• concrete compressive strength [2.38]

• strain at peak stress [2.71]

• maximum compressive strain [2.71]

• concrete tensile strength [2.71]
31

• initial concrete modulus of elasticity [2.53]

• unrestrained thermal strain [2.76]

• yield strength of reinforcing bars [2.77] 

• bond strength [2.78]

• transient creep strain (results from changes in the chemical composition of concrete on first 

heating) [2.53,2.71,2.72].

Figure 2.41 Effect of temperature on biaxial stress-strain curves for tension-compression loading.

Source: Y. Song, A. Zhang, L. Qing, and C. Yu, ―Biaxial Tensile-cCompressive Experiment on Concrete at High

Temperatures,‖ Frontiers of Architecture and Civil Engineering in China 1(1), pp. 94-98, 2007.

Additional information to that referenced above and what is provided below on modeling the strain components of 

concrete during heating is available [2.8,2.79].

An instantaneous stress-strain relationship for concrete with parabolic ascending and descending branches has been 

proposed [2.77]

 (2.1)

 

A more detailed description of the nomenclature for the instantaneous stress-strain information utilized in this 

section is provided elsewhere [2.80].
32

 (2.2)

Relationships have been proposed that are parabolic for the ascending branch and linear for the descending branch 

[2.53]

 (2.3)

 (2.4)

(2.5)

(2.6)

A model has been proposed that accounts for the effect of concrete weight on the shape of the stress-strain curve by 

using a non-dimensional factor n. [2.81]. The recommended value of n is between 2.5 and 3.0 for lightweight and 

normal weight concrete,.

(2.7)

Another researcher [2.71] recommended use of n = 2 in the above model. 

To account for transient creep effects, each of the above considered the total strain to be composed of separate 

components [2.80]. The thermal strain is a function of the temperature so it can be separated from the total strain. 

Calculation of transient creep strain requires an assumption of the corresponding stress which leads to an iterative 

solution.

Additional compressive and tensile stress-strain relationships have been proposed for concrete [2.80] that are based 

on two prior models [2.82,2.83]. The models were modified by replacing and oc with the temperature-dependent 

terms and oTc and modeling transient creep by shifting the strain at maximum stress by the transient creep strain 

[2.84].

Modification of the first model [2.82] for elevated temperature response is done by changing the strain 50u for 

ambient temperature shown in Figure 2.42 to o.

(2.8)

(2.9)

(2.10)

(2.11)
33

(2.12)

 (2.13)

Figure 2.42 Instantaneous stress-strain curve for concrete at ambient temperature.

Source: D.C. Kent and R. Park, ―Flexural Members with Confined Concrete,‖ Journal Structural Engineering

97(ST7), pp. 1969-1990, American Society of Civil Engineers, 1971.

Modification of the second model [2.83] is done by estimating Eci by using 2*f’cT/ oT (r = 2).

(2.14)

(2.15)

f’ccT for circular sections is

(2.16)

The value of can be taken as equal to and for rectangular sections is available based on 

area of stirrups and their temperature-dependent yield strength [2.83]. 

A comparison between different instantaneous stress-strain models {[2.53], [2.77], modified model 1 [2.82], and 

modified model 2 [2.83]} and experimental results [2.12,2.85,2.86] is provided in Figure 2.43. As noted in the 
34

figure, all models did a good job of representing the ascending portion of the stress-strain curve. Reference [2.82] 

provided the most accurate representation of the descending branch of the curves at different temperatures.

Figure 2.43 Instantaneous stress-strain curves at elevated temperature: 

comparison of predicted with experimental data.

Source:

Additional stress-strain constitutive relationships for concrete at high-temperature have been developed. The 

American Society of Civil Engineers has proposed the following relationships for normal-strength concrete [2.87]:

(2.17)

(2.18)

, 20 ̊C < T < 450 ̊C (2.19)
35

, 450 ̊C < T < 874 ̊C (2.20)

, T > 874 ̊C (2.21)

, (2.22)

where:

c = stress in concrete, MPa,

= 28-day concrete compressive strength in MPa at room temperature,

= concrete compressive strength in MPa at temperature,

T = concrete temperature in  ̊C,

= concrete mechanical strain,

= concrete strain at maximum stress at temperature T, 

= concrete ultimate strain at temperature T, and 

= concrete strain at maximum stress at temperature T.

This relationship has been modified for high-strength concrete [2.88]: 

(2.23)

(2.24)

, T < 100 ̊C (2.25)

, 100 ̊C < T < 400 ̊C (2.26)

, 400 ̊C < T (2.27)

where,

, and .

Finally, Eurocode has developed a relationship for the ascending portion of the stress-stain curve for both normaland high-strength concretes [2.46]:

, where and (2.28)

= compressive stress at temperature 

= compressive strain at temperature 

= ultimate concrete strain at temperature 

= compressive strength of concrete at temperature and

= compressive strength of concrete at room temperature.
36

Values of and are obtained from Table 2.2.

Table 2.2 Values for parameters to describe ascending branch of 

stress-strain relationship for concrete at elevated temperature

Concrete

Temperature

( ̊C)

x 10-3

Normal Concrete Normal Concrete Lightweight

Siliceous Calcareous Concrete

20

100

200

300

400

500

600

700

800

900

1000

1100

1200

1

0.95

0.90

0.85

0.75

0.60

0.45

0.30

0.15

0.08

0.04

0.01

0

1

0.97

0.94

0.91

0.85

0.74

0.60

0.43

0.27

0.15

0.06

0.02

0

1

1

1

1

0.88

0.76

0.64

0.52

0.40

0.28

0.16

0.04

0

2.5

3.5

4.5

6.0

7.5

9.5

12.5

14.0

14.5

15.0

15.0

15.0

15.0

Source:

The Eurocode permits the use of the linear as well as the nonlinear descending branch in the numerical analysis.

A tensile stress-strain relationship has been proposed [2.80]. Until the cracking stress [ ] occurs, the ascending 

portion of the stress-strain curve is represented as a linear branch. Recommended values for [ ] are

 direct tension, and (2.29)

flexural tension. (2.30)

After cracking, the model in Ref. [2.84] can be modified to account for the reduction in tensile resistance and bond 

strength [2.80]:

. (2.31)

Information on a thermoplasticity model that has been developed for concrete under transient temperature and 

biaxial stress is available [2.79].

Summary. Relative to temperature effects on concrete’s stress-strain curve, several general observations can be 

made based on information provided above. The ascending branch of the stress-strain curve consists of three 

components: (1) an elastic recoverable strain that is temperature dependent and is strongly influenced by the load 

level during initial heating to the test temperature; (2) an irrecoverable plastic strain component; and (3) a timedependent creep component that is normally small at room temperature but can be significant at high temperatures, 

particularly above 550 ̊C [2.63]. The extent of linearity of the stress-strain curve increases as the concrete strength 
37

increases and the slope of the curve tends to decrease as the temperature increases, except if material is under load 

while heated in which case the curve tends to become more linear and remain closer to its original slope [2.63]. The 

reduced brittleness of concrete with increasing temperature has been attributed to the evaporation of moisture with 

evaporation of gel water having a significant influence on the brittleness of concrete [2.89]. 

Under steady-state conditions, the original concrete strength, water-cement ratio, heating rate, and type of cement 

have minor influence on the stress-strain behavior. Aggregate-cement ratio, aggregate type, and presence of a 

sustained load during heating affect the shape of the stress-strain curve [2.46,2.77]. Concretes made with hard 

aggregates (e.g., siliceous or basalt) generally have a steeper decrease of the initial slope at high temperature (e.g., 

>55 Concrete specimens tested at temperature tend to be stiffer 

and stronger than identical companion specimens heated to the same temperatures and then permitted to cool to 

room temperature before testing [2.55]. High-strength concrete has a steeper and more linear stress-strain curve and 

tends to fail in a more brittle manner than normal-strength concrete [2.58]. 

 

The mechanical strains in biaxial compression loading were found to be dependent on the stress level, the stress 

ratio, and the test temperature. The deformations increased with increasing load in all three orthogonal directions 

and at high stress levels they showed a nonlinear relation in the direction of greatest principal stress. The ultimate 

strains for all three axes shift to greater values with increasing temperature. At higher temperatures the stress-strain 

curves for biaxial loading were highly nonlinear with the concrete behavior changing from brittle behavior to a soft 

and plastic behavior. When concrete is under a biaxial tension-compression state of stress the capability of concrete 

to resist cracking is diminished, both the tensile and compressive strengths are found to decrease rapidly with 

increasing exposure temperature, and the concrete fails abruptly due to tensile stress. In addition, the initial elastic 

modulus and the peak stress values of the stress-strain curves in both principal compressive stress direction and 

principal tensile stress direction decreased with an increase in exposure temperature and the stress-strain curves 

tended to flatten. 

2.2.1.2 Poisson’s Ratio

Information and Data. Poisson’s ratio is needed for conducting structural analyses of flat slabs, arch dams, 

tunnels, tanks, and other statically indeterminate members. At normal ambient conditions Poisson’s ratio for 

concrete can vary from 0.11 to 0.32, but is generally in the range from 0.15 to 0.20. Under uniaxial compression 

loading, the beginning of matrix cracking occurs at the stress level where there is an apparent increase in the 

Poisson’s ratio value [2.90]. At ambient temperature there appears to be no consistent relationship between 

Poisson’s ratio and concrete characteristics such as water-cement ratio, curing age, and aggregate gradation, but it is 

generally lower in high-strength concrete and higher for saturated concrete and for dynamically-loaded concrete 

[2.91]. 

Data on the effect of elevated temperature on Poisson’s ratio are somewhat limited and tend to be inconsistent. 

Figure 2.44 provides Poisson’s ratio results for uniaxially loaded concrete for temperatures from 20 ̊ to 750 ̊C 

[2.92]. At 20 ̊C the Poisson’s ratio is constant until the load exceeds about 70% of the ultimate load, while as the 

temperature increases the deviation of Poisson’s ratio increases till in some cases it is >0.5. Figure 2.45 indicates 

that the Poisson’s ratio decreases with increasing temperature from about 0.2 at room temperature to about 0.1 at 

400 ̊C [2.93]. This drop was attributed to a change in state due to heating that resulted in water desorption (e.g., 

weakening of the microstructure caused by breakage of bonds due to heating and by microcracking). When the 

specimen was permitted to cool after a given change in state the variation of Poisson’s ratio with temperature was 

slight and nil when the evaporable water had been removed. In general, Poisson’s ratio values obtained after drying 

are less than before drying [2.16]. In measurements on sealed cylinders the change in Poisson’s ratio with 

increasing temperature was small (e.g., 0.2 to 0.18) [2.65]. When the concrete is under a confining pressure as it 

would be in many nuclear power plant concrete structures, it has been hypothesized that Poisson’s ratio at elevated 

temperature would be about the same as at room temperature [2.16]. 
38

Figure 2.44 Poisson’s ratio for uniaxially loaded concrete at high temperature.

Source: C. Ehm, ―Experimental Investigations of the Biaxial Strength and Deformation of Concrete at High 

Temperatures,‖ Dissertation, Technical University of Braunschweig, Germany, 1985. 

Figure 2.45 Poisson’s ratio as a function of temperature for a quartzite concrete. 

Source: J.C. Marechal, ―Variations in the Modulus of Elasticity and Poisson’s Ratio with Temperature,‖ Paper SP 

34-27 in Concrete for Nuclear Reactors, pp. 495-503, American Concrete Institute, Farmington Hills, 

Michigan, 1972.

The effect of elevated temperature on elastic properties of concretes containing either a carbonate, siliceous, or 

lightweight aggregate was evaluated using unsealed specimens [2.94]. A 8.53-mm maximum size aggregate was 

used in the mixes. All specimens were moist cured for 3 days at 23.9 ̊C and then stored in air at 50% relative 

humidity and 23.0 ̊C for 25 days. Two nominal concrete design compressive strengths were investigated: 27.6-

31.0 MPa and 41.4-44.8 MPa. Figure 2.46 presents Poisson’s ratio results as a function of temperature for the 

different aggregate materials and the two nominal concrete compressive strengths. The Poisson’s ratio values 

generally ranged from 0.11 to 0.25. Although results obtained at higher temperature were somewhat erratic and a 

general trend was not apparent, results did show lower Poisson’s ratio values at room temperature for the higher 

strength concretes. 
39

Figure 2.46 Effect of aggregate type and concrete strength on Poisson’s ratio.

Source: C.R. Cruz, ―Elastic Properties of Concrete at High Temperatures,‖ Department Bulletin 191, Journal of the 

Portland Cement Association Research and Development Laboratories 8(1), pp. 37-45, January 1966.

Elsewhere Poisson’s ratio for high strength concrete has been reported to range from 0.11 at 20°C to 0.25 at 400°C, 

while above 400°C it increased [2.95]. Additional data for high strength concrete indicated that when the stress did 

not exceed 50% of peak value, the Poisson’s ratio decreased with an increase in temperature [2.96]. Figure 2.47 

presents Poisson’s ratio results for a hard sandstone aggregate concrete after various heating periods (i.e., 1, 7, 28, 

and 91 d) at 175°C for specimens that were either sealed or unsealed during heating [2.97]. Poisson’s ratio ranged 

from 0.14 to 0.22 with the trend for it to be larger for the sealed specimens.

Figure 2.47 Poisson's ratio results for sealed and unsealed concrete specimens.

Source: K. Hirano, K. Ohmatsuzawa, T. Takeda, S. Nakane, T. Kawaguchi, and K. Nagao, ―Physical Properties of 

Concrete Subjected to High Temperature for MONJU,‖ Paper P2-25, Power Reactor and Nuclear Fuel 

Development Corporation, Tokyo, Japan.

The effect of thermal cycling (ambient to 180 ̊C to ambient) on Poisson’s ratio over a period of 1,024 days was 

evaluated for a concrete that was similar to that used for the Hanford nuclear underground storage tanks [2.98,2.99]. 

Two concrete mixes having nominal compressive strengths of 20.7 and 31.0 MPa were utilized in the investigation. 
40

The concrete was fabricated from Type II Portland cement and contained basalt coarse aggregate. Two thermal 

cycling programs were utilized, either a 14- or 28-day period for cycling from 21 ̊ to 176.7 ̊ to 21 ̊C. Poisson’s ratio 

was determined at room temperature. Figure 2.48 presents Poisson’s ratio results as a percentage of initial room 

temperature value versus number of temperature cycles. Measured values after one 28-day thermal cycle were 

approximately 35% lower than those obtained from the unheated specimens, however, with increasing number of

thermal cycles the Poisson’s ratio recovered somewhat. Results for the 14-day thermal cycle series were similar 

except the minimum in Poisson’s ratio occurred later (e.g., after 3 cycles) and the reduction was not as great. The 

magnitude of longitudinal and lateral strains measured was larger for the long-term heated concrete than for the 

unheated concrete, though their Poisson’s ratio was essentially the same. The variation in Poisson’s ratio with 

length of exposure to 176.7 ̊C is presented in Figure 2.49. As the length of exposure to the maximum temperature 

level increased the Poisson’s ratio tended to return to the unheated value with the lower strength concrete recovering 

more than the higher strength concrete. 

Figure 2.48 Variation of Poisson's ratio of basalt aggregate concrete with 

number of thermal cycles (ambient to 176.7 ̊C to ambient).

Source: C. Defigh-Price, ―Effects of Long-Term Exposure to Elevated Temperature on the Mechanical Properties 

of Hanford Concrete,‖ Report RHO-C-54, Construction Technology Laboratories, Portland Cement 

Association, Skokie, Illinois, October 1981.

Figure 2.49 Variation of Poisson's ratio of basalt aggregate concrete with length of exposure to 176.7 ̊C.

Source: C. Defigh-Price, ―Effects of Long-Term Exposure to Elevated Temperature on the Mechanical Properties 

of Hanford Concrete,‖ Report RHO-C-54, Construction Technology Laboratories, Portland Cement 

Association, Skokie, Illinois, October 1981.
41

Summary. Poisson’s ratio data at elevated temperature are very limited and results tend to be inconsistent and 

appear to not exhibit a strong dependence on temperature. Values obtained after drying are less than those obtained 

before drying and for moderate temperatures (e.g., T<180 ̊C) the values tend to increase with increasing 

temperature. Thermal cycling (e.g., T<180 ̊C) reduces Poisson’s ratio during the first few cycles, but as the number 

of cycles increases the Poisson’s ratio value tends to recover. The initial reduction in Poisson’s ratio with thermal 

cycling tends to increase as the concrete strength increases.

2.2.1.3 Modulus of Elasticity

Information and Data. Concrete’s modulus of elasticity—a measure of its stiffness or resistance to deformation—

is used extensively in the analysis of reinforced concrete structures to determine the stresses developed in simple 

elements and the stresses, moments, and deflections in more complicated structures. Because concrete’s stressstrain curve is nonlinear, the modulus of elasticity is determined either by the initial tangent modulus, secant 

modulus, or tangent modulus method. Most results presented in the literature correspond to static tests conducted 

under steady-state stress rate or steady-state strain rate control. Temperature can significantly affect the modulus 

values causing a reduction as a result of the breakage of bonds [2.16]. Drying which increases with temperature 

reduces the apparent modulus value to produce the bond rupture and it has been observed that fast drying results in a 

faster decrease in the modulus of elasticity than slow drying [2.100,2.101]. Several potential factors affecting the 

modulus of elasticity due to elevated temperature exposure have been investigated (e.g., mix proportions, aggregate 

type, cement type, concrete strength, sealed or unsealed, presence of sustained stress during heating, and duration of 

thermal exposure). Results addressing several of these factors are provided below.

Normal Weight Concretes

Normal weight concretes are typically concretes used for structural purposes that contain natural sand and gravel or 

crushed rock aggregates and weigh around 2400 kg/m3. Figure 2.50 presents results for normal-strength concrete 

(NSC) and high-strength concrete (HSC) from researchers in China [2.78]. Results show that the elastic modulus 

for the NSC decreased monotonically with increasing temperature. From the NSC and HSC elastic modulus results 

obtained at temperature (relative) or after thermal exposure (residual), Reference 2.78 notes that the elastic modulus 

after high-temperature exposure (residual) was lower than that obtained at temperature and was influenced by type 

aggregate, the elastic modulus decreased much more for concrete cured in water than for concrete cured in air, and 

the deterioration in elastic modulus was more related to the maximum temperature during heating than to the 

heating-cooling cycle. 

Figure 2.50 Temperature dependence of the concrete modulus of elasticity (normalized).

Source: J. Xiao and G. König, ―Study of Concrete at High Temperature in China—An Overview,‖ 

Fire Safety Journal 39, pp. 89–103, 2004.
42

Figure 2.51 presents the strong influence of aggregate type on modulus [2.15]. Conclusions from this study were 

that a sustained stress during heating affects the modulus significantly and the type of cement had little effect. 

Figure 2.52 illustrates the influence of water-cement ratio on static and dynamic modulus [2.12]. Results presented 

in Figure 2.53 for a 31-MPa and a 63-MPa limestone aggregate concrete tested at temperature indicate that, when 

normalized with respect to the room temperature modulus of elasticity, the strength of the concrete does not have a 

significant effect on the modulus-temperature response [2.58].

Figure 2.51 Modulus of elasticity of different concretes at elevated temperature.

Source: U. Schneider, C. Diererichs, and C. Ehm, ―Effect of Temperature on Steel and Concrete for PCRV’s,‖ 

Nuclear Engineering and Design 67, pp. 245–258, 1981.

Figure 2.52 Influence of water-cement ratio on modulus of elasticity of concrete at elevated temperature.

Source: U. Schneider, Behaviour of Concrete at High Temperature, HEFT 337, Deutscher Ausschuss für Stahlbeton, 

Wilhelm Ernst & Sohn, Munich, Germany, 1982 (W/Z = water-cement ratio).
43

Figure 2.53 Normalized modulus of elasticity vs temperature relationships: NSC and HSC.

Source: C. Castillo and A. J. Durani, ―Effect of Transient High Temperature on High-Strength Concrete,‖ J. American 

Concrete Institute, 87(10), pp. 47-53, January-February 1990.

Results from information presented above as well as those from other researchers have been used to develop a series 

of figures illustrating the effect of elevated temperature on the modulus of elasticity of concrete materials. 

References to data sources cited by others or used in figures are identified in Appendix A. Test results are presented 

in terms of testing procedures identified earlier in the report that are commonly used to evaluate concrete response to 

elevated temperature: unstressed tests (hot testing), unstressed residual (cold testing), and stressed tests.

Figure 2.54 summarizes results from several researchers on the temperature dependence of the concrete elastic 

modulus (normalized to reference room temperature value) [2.102]. Also shown in the figure are upper and lower

Figure 2.54 Effect of temperature on the modulus of elasticity of concrete: hot and cold test results.

Source: G. N. Freskakis, ―Behavior of Reinforced Concrete at Elevated Temperature,‖ Paper 3-4, Source: ASCE 

Conf. on Civ. Eng. and Nuclear Power 1, Paper 3-5, pp. 3-5-1 to 3-5-21, Knoxville, Tennessee, Sept. 15–

17, 1980.
44

bounds that were established from the data. The results were utilized as part of a study to investigate the behavior of 

reinforced concrete sections at elevated temperature.

Data on effects of elevated temperature on modulus of elasticity have also been assembled as part of a review of 

concrete properties for prestressed concrete pressure vessels [2.103]. Figure 2.55 presents a summary of data 

assembled under this activity for sealed and unsealed specimens that were either tested at temperature (hot) or after 

returning to room temperature (cold). It was concluded that generally the modulus decreased with increasing 

temperature for all types of testing, further decreases in modulus occurred with an increase in number of thermal 

cycles and exposure time, and sealed specimens were more sensitive to these factors with decrease in modulus 

becoming fairly significant at temperatures above 150 ̊C.

Figure 2.55 Compilation of data on modulus of elasticity vs temperature.

Source: R.K. Nanstad, A Review of Concrete Properties for Prestressed Concrete Pressure Vessels, ORNL/TM5497, Oak Ridge National Laboratory, Oak Ridge, Tennessee, October 1976.

Figure 2.56 presents a compilation developed under the present study on the effect of elevated temperature on the 

relative modulus of elasticity (i.e., unsealed, hot testing) for a number of different concrete mixtures (i.e., ordinary 

Portland cement concretes and ordinary Portland cement concretes that also included supplementary cementitious 

materials). Results obtained for only the Portland cement concretes are provided in Figure 2.57 and results for 

Portland cement concretes containing supplementary cementious materials are presented in Figure 2.58. The CEB 

design curve for unstressed normal weight concretes has been superimposed on the figures to facilitate 

comparisons. Results indicate similar trends toward a decrease in relative modulus of elasticity with increasing 

temperature for both the ordinary Portland cement concretes and the Portland cement concretes that contained 

supplementary cementitious materials. There may be a slight trend for the reductions in modulus to be slightly 

higher for the Portland cement concretes with supplementary cementitious materials than the ordinary Portland 

cement concretes, however, the concrete mixes with supplementary cementitious materials were all high-strength 

concretes (fc’ > 60 MPa) whereas the vast majority of ordinary Portland cement concretes were normal strength 

concretes.

 

Comites Euro-International Du Beton, Fire Design of Concrete Structures – in accordance with CEB/FIP Model 

Code 90, CEB Bulletin D’Information No. 208, Lausanne, Switzerland, July 1991.
45

Figure 2.56 Compilation of data on relative modulus of elasticity vs temperature – ordinary Portland cement 

concretes and concretes containing supplementary cementitious materials.

Figure 2.57 Compilation of data on relative modulus of elasticity vs

temperature – ordinary Portland cement concretes.
46

Figure 2.58 Compilation of data on relative modulus of elasticity vs 

temperature - concretes containing supplementary cementitious materials.

A data compilation developed under the present study on the effect of elevated temperature on the residual modulus 

of elasticity (i.e., unsealed, cold testing) for a number of different concrete mixtures (i.e., ordinary Portland cement 

concretes and Portland cement concretes that also included supplementary cementitious materials) is provided in 

Figure 2.59. Results obtained for only the Portland cement concretes are provided in Figure 2.60 and results for

Figure 2.59 Compilation of data on residual modulus of elasticity vs temperature – ordinary Portland cement 

concretes and concretes containing supplementary cementitious materials.
47

Portland cement concretes containing supplementary cementious materials are presented in Figure 2.61. The CEB 

design curve for unstressed normal weight concretes has been superimposed on the figures to facilitate comparisons. 

Results indicate similar trends toward a decrease in relative modulus of elasticity with increasing temperature for 

both the ordinary Portland cement concretes and the Portland cement concretes that contained supplementary 

cementitious materials. Comparing the results to those obtained from specimens tested at temperature, the retained 

modulus at temperature tends to be higher than that obtained from specimens permitted to cool to room temperature 

prior to testing. 

Figure 2.60 Compilation of data on residual modulus of elasticity vs

temperature – ordinary Portland cement concretes.

Figure 2.61 Compilation of data on residual modulus of elasticity vs 

temperature - concretes containing supplementary cementitious materials.

The effect of sustained stress during heating on the relative modulus of elasticity (unsealed, hot testing) is presented 

in Figure 2.62 for ordinary Portland cement concretes and Portland cement concretes that also included 

supplementary cementitious materials. Preload levels ranged from 20 to 40% the reference room temperature 

compressive strength of the concretes. Although results are limited, the presence of preload and its beneficial effects 

relative to retained modulus is evident.
48

Figure 2.62 Compilation of data on relative modulus of elasticity vs temperature for preloaded specimens –

ordinary Portland cement concretes and concretes containing supplementary cementitious materials.

The effect of cooling regime on the residual modulus of elasticity has been studied as part of an overall investigation 

to evaluate the post-thermal properties of concrete following rapid cooling (e.g., water quenching in conjunction 

with fire) [2.67]. A granite aggregate ordinary Portland cement concrete having a reference compressive strength of 

21.16 MPa and reference initial tangent modulus of 24.8 GPa was used to fabricate 101.6-mm-diameter by 203.2-

mm-long cylindrical test specimens that were cured for eight weeks in a fog room (23 ̊C and 93% relative humidity). 

The specimens were heated in a furnace to 200 ̊, 400 ̊, 600 ̊, or 800 ̊C with a holding time at temperature of four 

hours. Following heating, the specimens were cooled slowly (1 ̊C/minute), naturally (left in furnace and permitted 

to cool slowly), or rapidly (e.g., specimen removed from furnace and placed into a tank of water initially at 20 ̊C). 

Figure 2.63 presents the effect of cooling regime on the initial relative tangent modulus for concretes subjected to 

the maximum temperatures up to 800 ̊C.

 

Figure 2.63 Effect of cooling regime on initial tangent modulus following thermal exposure.

Source: J. Lee, Y. Xi, and K. Willam, ―Concrete Under High Temperature Heating and Cooling,‖ Report SESM 

No.09/2006, Department of Civil, Environmental, and Architectural Engineering, University of Colorado at 

Boulder, 2006.
49

Additional results on the effect of cooling regime on the residual relative initial tangent modulus are presented in 

Figure 2.64 for a limestone aggregate ordinary Portland cement concrete [2.104]. The 75-mm-diameter by 200-mmlong cylindrical test specimens were obtained by coring 700 x 500 x 200 mm3

concrete blocks that had been cast and 

cured in water at 20 ̊C for 28 days followed by conditioning in a laboratory environment (20 ̊C and 65% relative 

humidity) for several months to permit the moisture content to stabilize. Following curing and conditioning the 

specimens were placed into an oven and heated to target exposure temperatures of 217 ̊, 287 ̊, 320 ̊, 378 ̊, or 470 ̊C 

where they remained until temperatures at the outer surfaces and centers of the specimens attained uniformity. 

Following thermal treatment, half the specimens at each temperature were removed from the furnace and permitted 

to cool in a controlled environment at 20 ̊C and the other half were taken out of the furnace and sprayed with tap 

water for a duration of five minutes and then permitted to stabilize in a controlled environment at 20 ̊C. The cooling 

regime results show the sensitivity of the tangent modulus to water quenching, particularly in the lower temperature 

range. The effect of quenching on the tangent modulus decreases somewhat at the higher temperatures, however, 

the initial tangent modulus already had been severely reduced by exposure to these temperatures.

Figure 2.64 Effect of cooling regime on relative residual tangent modulus following thermal exposure.

Source: Y. Nassif, S. Rigden, and E. Burley, ―The Effects of Rapid Cooling by Water Quenching on the Stiffness 

Properties of Fire-Damaged Concrete,‖ Magazine of Concrete Research 51(4), pp. 255-261, August 1999.

The effect of elevated temperature followed by water quenching on the modulus of elasticity of high-strength 

concrete containing supplementary cementitious materials such as used in construction of high rise buildings was 

studied [2.105]. Four mixes were investigated: Mix 1 - 100% general purpose cement, Mix 2 - 62% slag plus 38% 

cement, Mix 3 - 20% fly ash plus 80% cement, and Mix 4 - 6.6% silica fume plus 93.4% cement. Crushed basalt 

aggregate having a maximum aggregate size of 10 mm and a water-cement ratio of 0.30 (by weight) was used for all 

the mixes. Table 2.3 presents reference 28-day properties of the mixes. Specimens used for dynamic modulus of 

elasticity determinations were 75 x 75 x 305 mm3prisms. Specimens were cured by storing in water at 20 ̊C for 

28 days and then placing into the laboratory environment of 20 ̊C and 65% relative humidity for 2 days. At this 

time the specimens were heated to temperatures of 200 ̊, 400 ̊, 600 ̊, 800 ̊, or 1000 ̊C with a total duration of

 

The effect of binder type and water quenching on the relative compressive, flexural, and tensile strengths of these 

concretes is presented later in this report.
50

Table 2.3 Properties of concrete mixes at the age of 28 days

Mix Binder Materials

Cylinder

strength,

MPa

Tensile

strength,

MPa

Flexural 

strength,

MPa

Dynamic 

modulus,

GPa

1 100% OPC 61.8 6.58 9.34 46.5

2 62% slag + 38% OPC 40.3 4.28 5.47 39.9

3 20% fly ash + 80% OPC 67.2 5.08 6.94 45.7

4 6.6% silica fume+ 93.4% OPC 76.4 6.93 8.51 55.2

Source: R.S. Ravindrarajah, R. Lopez, and H. Reslan, ‖Effect of Elevated Temperature on Properties of HighStrength Concrete Containing Cement Supplementary Materials,‖ 9

th International Conference on 

Durability of Building Materials and Components, 9 p., Brisbane, Australia, 17-20 March 2002.

7 hours (i.e., soak period from 6 to 2 hours depending on maximum temperature). Figure 2.65 presents the 

normalized residual relative modulus of elasticity results for specimens subjected to elevated temperature followed 

by quenching in water prior to testing. Elastic modulus decreased as the exposure temperature increased with the 

results influenced by the binder material type (e.g., silica fume concrete decreased the most followed by the fly ash 

concrete). At 200 ̊C the reduction in modulus ranged from 8 to 16%, at 800 ̊C the reduction ranged from 13 to 20%, 

and at 1000 ̊C the reduction ranged from 29 to 40%.

Figure 2.65 Normalized residual modulus of elasticity results for specimens subjected to

elevated temperature and quenched in water prior to testing.

Source: R.S. Ravindrarajah, R. Lopez, and H. Reslan, ‖Effect of Elevated Temperature on Properties of HighStrength Concrete Containing Cement Supplementary Materials,‖ 9

th International Conference on 

Durability of Building Materials and Components, 9 p., Brisbane, Australia, 17-20 March 2002.

Lightweight and Thermally Stable Aggregate Concretes

Lightweight concrete material systems for structural applications typically contain natural or pyro-processed 

aggregates with relatively low bulk densities (e.g., expanded shale, slate, or slag) and generally weigh less than 

1800 kg/m3. Relative modulus of elasticity results (unsealed, hot testing) for lightweight/thermally stable aggregate 

concretes compiled from the literature are presented in Figure 2.66. The CEB design curve for lightweight concrete 

is superimposed on the data. Results indicate that in general the lightweight/insulating concretes tend to retain their 

modulus of elasticity at temperature better than normal weight concretes. The improved performance can probably 

be attributed to the modulus of elasticity of the lightweight aggregates used in the mixes being closer to that of the 

Portland cement paste [2.19]. 
51

Figure 2.66 Compilation of data on relative modulus of elasticity vs temperature for lightweight/insulating concrete 

specimens – ordinary Portland cement concretes and concretes containing supplementary cementitious materials.

Fibrous Concretes

Fibrous concretes is made from hydraulic cements containing fine and coarse aggregate materials and discontinuous 

discrete fibers (e.g., steel, polypropylene, and aramid). Relative modulus of elasticity results (unsealed, hot testing) 

compiled from the literature for fibrous concretes and the CEB design curve noted earlier are presented in 

Figure 2.67. Insufficient data are available to draw any definite conclusions relative to incorporation of 

discontinuous fiber reinforcement in the concrete mixes. However, polypropylene fibers (and steel fibers) have been 

incorporated into concrete mixes (high strength concretes in particular) to improve fire resistance and enhance 

resistance of the concrete to thermal spalling. The polypropylene fibers melt at around 170 ̊C to provide ―channels‖ 

for the steam pressure that develops in the concrete to escape thus preventing internal pressurization that can 

produce spalling.

Figure 2.67 Compilation of data on relative modulus of elasticity versus temperature for fibrous concretes.
52

Examples of Relations for Concrete Modulus of Elasticity at Elevated Temperature. Assuming the secant 

modulus at a stress of is the elastic modulus, the following has been proposed for estimating the modulus at 

elevated temperature [2.106]:

for 20 ̊ < T < 300 ̊C, (2.32)

 for 300 ̊ < T < 700 ̊C, and (2.33)

 for T > 700 ̊C. (2.34)

A bi-linear equation has also been suggested [2.107]:

 for 20 ̊ < T < 60 ̊C, and (2.35)

 for 60 ̊ < T < 700 ̊C. (2.36)

A relationship has been developed for the normalized elastic modulus ( ) variation with temperature (T) 

[2.108]:

for 20 ̊ < T < 125 ̊C, and (2.37)

for 125 ̊ < T < 800 ̊C. (2.38)

Reductions in initial tangent modulus of elasticity E0 for concrete can be estimated from [2.72]:

E(T) = E0 for T > 60 ̊C, and (2.39)

for 60 ̊ < T < 800 ̊C. (2.40)

The effect of elevated temperature on the concrete secant modulus at a stress corresponding to 40% the room 

temperature compressive strength of concrete can be estimated from [2.78]:

for 20 ̊ < T < 300 ̊C, (2.41)

for 300 ̊ < T < 700 ̊C, and (2.42)

for T > 700 ̊C, (2.43)

where EC and are the concrete elastic moduli at room and elevated temperature. 

Summary. An increase in temperature generally leads to a continuous fall in the modulus of elasticity of both 

sealed and unsealed concrete, with the decrease being larger for sealed concrete. The decrease in elastic modulus 

with elevated temperature exposure is primarily due to breakage of bonds in the cement paste microstructure. 

Variations in results from different authors can be caused by a number of factors in addition to the type of aggregate 

used (e.g., water-cement ratio, method for determination of modulus, testing hot or cold, presence of pre-load, rate 

of heating and cooling, and rate of loading). Modulus of elasticity at temperature is higher than residual modulus. 

Variation of modulus values with temperature up to 80 ̊C is considerable, primarily as a result of use of different 

aggregate materials, and above 100 ̊C the decrease in modulus tends to be linear up to a critical temperature at 

which the concrete experiences severe deterioration [2.19]. Due to thermal incompatibility with the cement paste, 

concretes containing aggregates with low thermal expansion (e.g., limestone) experience a greater reduction in the 

modulus of elasticity than those with a higher thermal expansion such as gravels and sandstone [2.19]. The primary 
53

factors affecting the modulus of elasticity at high temperature are the type of aggregate and the presence of sustained 

stress during heating (sustained stress results in lower decreases in modulus with increasing temperature). Thermal 

cycling causes a decrease in modulus. Fast drying accompanying heating results in a greater decrease in modulus 

than slow drying due to increased breakage of bonds. Faster heating rates result in a lower modulus of elasticity. 

Lightweight aggregate/thermally stable concretes retain their modulus at elevated temperature better than normal 

weight aggregate concretes. Normal strength concretes (fc’ < 60 MPa) retain their modulus better at elevated 

temperature than high-strength concretes. Duration of temperature exposure, type of cement, water-cement ratio and 

original concrete strength appear to have little effect on modulus results. The age at test apparently also does not 

affect the residual modulus as noted for a flint/beach gravel concrete for which results were obtained to 150°C at 

concrete ages of 3 months and 1 year [2.109]. 

2.2.1.4 Compressive Strength

Information and Data. The concrete compressive strength is generally considered to be its most important 

properties and influences the load-carrying capacity of a structure. Thermal loadings in addition to affecting the 

concrete modulus of elasticity also affect the concrete compressive strength. Most results presented in the literature 

correspond to static tests conducted under steady-state stress rate or steady-state strain rate control, utilized unsealed 

specimens, and determined the residual compressive strength (unsealed, cold testing). However, a number of tests 

are also available in which the relative concrete compressive strength (unsealed, hot testing) was determined. 

Several potential factors affecting the compressive strength of concrete due to elevated temperature exposure were 

investigated (e.g., mix proportions, aggregate type, cement type, concrete strength, sealed or unsealed, presence of 

sustained stress during heating, and duration of thermal exposure). Results addressing several of these factors are 

provided below.

Cements, Mortars, and Normal Weight Concretes

Figure 2.68, which presents results for unsealed mortars fabricated from ordinary Portland cement, blast furnace slag 

cement, and truss cement (i.e., mixture of Portland cement, volcanic tuff, and gypsum) indicates little difference in 

strength-temperature characteristics, except the ordinary Portland cement mortar exhibited a sharper decrease in 

strength at temperatures greater than 500°C [2.12]. Another study utilizing a number of cementitious materials [e.g., 

ordinary Portland cement, fly ash, and blast furnace slag cement (cemsave)] also noted that up to 600°C there was 

little effect of the cement type [2.110]. Results presented in Figure 2.69 indicate that partial replacement of ordinary 

Portland cement with pulverized fly ash improves the residual strength and may even produce an increase in strength 

at higher temperatures [2.8,2.11].* Figure 2.70 indicates that for the three water-cement ratios investigated, the 

residual compressive strengths of unsealed cement paste specimens experienced similar trends [2.8]. These results 

indicate a peak strength at about 150°C where the residual strengths were 10 to 30% higher than the reference room 

temperature strength. Although the strength declined at temperatures higher than 300°C to 350°C, at temperatures 

of 300°C to 350°C it was still higher than the reference strength. Research presented elsewhere involving several 

aggregate types indicates that the effect of water-cement ratio had little influence on the residual compressive 

strength up to 600°C [2.111]. For normal strength Portland cement concretes, the concrete strength has a secondary 

effect on strength-temperature characteristics. Residual compressive strength results are presented in Figure 2.71 for 

concretes having compressive strengths ranging from 19.6 to 74 MPa [2.112]. This study concluded that a highstrength concrete (e.g, fc’ > 60 MPa) has a higher rate of reduction in residual compressive strength (and modulus) 

than a normal-strength concrete. Figure 2.72 presents results of unstressed tests for ordinary Portland cement 

concretes having reference compressive strengths either of 21, 42, or 60 MPa [2.113]. Results for each of the 

concretes were similar in that the compressive strength decreased at 100°C, recovered to the room temperature 

strength at 200°C, and then decreased monotonically with increasing temperature beyond 200°C. Figure 2.73 shows 

the effect of curing conditions (e.g., in water unsealed, sealed, in-air unsealed) prior to elevated-temperature 

exposure [2.97]. After > 91 days cure, the specimens, either in a sealed or unsealed condition, were subjected to 

175°C for up to 91days. Unstressed and residual compressive strengths were determined periodically over the 

exposure period. Although differences in compressive strengths occurred at smaller exposure ages, after 91 days 

exposure to 175°C, similar results were provided under all test conditions.

 

*A similar effect was achieved through partial replacement of the ordinary Portland cement with ground granulated 

blast furnace slag. Partial replacement with silica fume was not beneficial and in some cases produced detrimental 

residual strength results for unsealed specimens.
54

Figure 2.68 Influence of type of cement on strength loss of mortars.

Source: U. Schneider, Behaviour of Concrete at High Temperature, HEFT 337, Deutscher Ausschuss fiir 

Stahlbeton, Wilhelm Ernst & Sohn, Munich, Germany, 1982.

Figure 2.69 Influence of partial replacement of (a) OPC and (b) slag on residual compressive strength.

Source: G. A. Khoury, Performance of Heated Concrete—Mechanical Properties, Contract NUC/56/3604A with 

Nuclear Installations Inspectorate, Imperial College, London, United Kingdom, August 1996.
55

Figure 2.70 Effect of water-cement ratio on residual compressive strength of ordinary Portland cement.

Source: G. A. Khoury, Performance of Heated Concrete—Mechanical Properties, Contract NUC/56/3604A with 

Nuclear Installations Inspectorate, Imperial College, London, United Kingdom, August 1996.

Figure 2.71 Residual normalized strength vs temperature.

Source: T. Morita, H. Saito, and H. Kumajai, ―Residual Mechanical Properties of High Strength Concrete Members 

Exposed to High Temperature—Part 1. Test on Material Properties,‖ Summaries of Technical Papers of 

Annual Meeting, Architectural Institute of Japan, Naiigata, August 1992.
56

Figure 2.72 Residual compressive strength vs temperature.

Source: T. Furumura, T. Abe, and Y. Shinohara, ―Mechanical Properties of High Strength Concrete at High 

Temperatures,‖ Proceedings of 4th Weimar Workshop on High Performance Concrete: Material Properties 

and Design, held at Hochschule für Architektur und Bauwesen (HAB), Weimar, Germany pp. 237–254, 

4–5 October 1995.

Figure 2.73 Effect of curing conditions prior to elevated-temperature exposure on relative compressive strength.

Source: K. Hirano, K. Ohmatsuzawa, T. Takeda, S, Nakane, T. Kawaguchi, and K. Nagao, ―Physical Properties of 

Concrete Subjected to High Temperature for MONJU,‖ Paper P2-25, Power Reactor and Nuclear Fuel 

Development Corporation, Tokyo, Japan.

The age at which the concrete is subjected to elevated-temperature exposure has little effect on the strengthtemperature response as long as the concrete has adequate curing (e.g., 90 days). Figure 2.74 presents a comparison 

of residual unstressed compressive strength results for a nuclear power plant concrete (Temelin NPP) at ages of 28-

days and 90-days for temperatures to 280°C [2.114]. Figure 2.75 provides residual compressive strength results for 

sealed and unsealed specimens cast from a nuclear power plant siliceous aggregate concrete at curing ages of 

3 months and 1 year [2.109]. These results indicate that some improvement in residual compressive stress due to 
57

more extended curing occurs, but it was not significant, and that the strength of sealed specimens was lower than 

that of unsealed specimens. 

Figure 2.74 Effect of temperature on compressive strength ratio of concrete.

Source: F. Vodak, K. Trtík, O. Kapicková, S. Hisková, and P. Demo, ―The Effect of Temperature on Strength—

Porosity Relationship for Concrete,‖ Construction and Building Materials 18, pp. 529–534, 2004.

Figure 2.75 Effect of curing age and exposure condition on 

residual compressive strength (S = sealed, U = unsealed).

Source: J. Guo and P. Waldron, ―Deterioration of PCPV Concrete,‖ Nuclear Engineering and Design 198, pp. 211–

226, 2000.

As long as the rate of heating does not produce significant thermal gradients, the rate of heating has a secondary 

effect on the concrete strength-temperature response, particularly at high temperatures. Figure 2.76 indicates the 

effect of rate of heating on residual strength of a crushed basalt concrete [2.115]. These results indicate that the rate 

of heating had an insignificant effect on residual strength for temperatures of 600°C and 800°C. However, the 

heating rate had an effect on the residual compressive strength at lower temperatures. Exposure times at 

temperature beyond 1 hour had an effect on the residual compressive strength, but this effect diminished as the level 
58

of exposure temperature increased, with the majority of strength loss occurring in the first 2 hours of exposure 

(Fig. 2.77). Figure 2.78 shows the effect of duration of temperature exposure (exposure periods to 42 days) on the 

relative strength change of sealed and unsealed gravel and limestone aggregate concretes at exposure temperatures 

to 180°C. For all temperatures investigated, and for sealed and unsealed conditions, the river gravel concrete 

generally exhibited a slight strength increase. The compressive strength of the limestone aggregate concrete was 

relatively constant with increasing exposure period for the unsealed condition, but exhibited a decline in

compressive strength with exposure time at 180 ̊C for the sealed condition. The explanation for this behavior was 

that the strength loss of the limestone concrete was caused by changes in the microstructure of the hydrated cement 

paste when exposed to hydrothermal conditions. In the river gravel concrete the strength loss was counteracted by a 

reaction between the silicates of the fine aggregate particles and the Ca(OH)2 of the hydrated cement paste. 

The aggregate type is one of the main factors influencing the compressive strength of concrete at elevated 

temperature. Figure 2.79 presents a comparison of stressed, unstressed residual, and unstressed compressive 

strength results for carbonate and siliceous aggregate concretes at temperatures to 871°C [2.116]. Results of this 

study indicate the influence of the aggregate type, show the beneficial effect of stressing the specimens (within 

limits) while heated, and indicate that unstressed results obtained at temperature exceed those obtained under 

unstressed residual conditions (i.e., hot strength results were generally greater than residual strength results). The 

influence of aggregate type on results is further illustrated in Figure 2.80 where the compressive strength of 

limestone and other concretes is presented [2.12].* It was noted in this reference that for the temperature range 

shown in the figure, quartz and basalt aggregate are less sensitive to temperature effects than the limestone 

aggregate concrete.

Figure 2.76 Effect of rate of heating on residual strength of slowly cooled concrete.

Source: G. T. C. Mohamedbhai, ―Effect of Exposure Time and Rates of Heating and Cooling on Residual Strength 

of Heated Concrete,‖ Magazine of Concrete Research 38(136), pp. 151–158. September 1986.

 

*See also Figure 2.207.
59

Figure 2.77 Effect of exposure time on residual strength of coral sand and basalt sand concretes.

Source: G. T. C. Mohamedbhai, ―Effect of Exposure Time and Rates of Heating and Cooling on Residual Strength 

of Heated Concrete,‖ Magazine of Concrete Research 38(136), pp. 151–158, September 1986.

Figure 2.78 Relative strength development of concrete exposed to elevated temperature.

Source: R. Kottas, J. Seeberger, and H. K. Hilsdorf, ―Strength Characteristics of Concrete in the Temperature 

Range of 20° to 200°C,‖ Paper H01/4 in 5th International Conference on Structural Mechanics in Reactor 

Technology, 8 p., Elsevier Science Publishers, North-Holland, The Netherlands, August 1979.
60

Figure 2.79 Effect of exposure time on residual strength of carbonate and siliceous aggregate concretes.

Source: M. S. Abrams, ―Compressive Strength of Concrete at Temperatures to 1600°F,‖ SP-25, Temperature and 

Concrete, American Concrete Institute, pp. 33–58, 1971.

Figure 2.80 Compressive strength of concretes with limestone and other aggregate types.

Source: U. Schneider, Behaviour of Concrete at High Temperature, HEFT 337, Deutscher Ausschuss für Stahlbeton, 

Wilhelm Ernst & Sohn, Munich, Germany, 1982.
61

Figures 2.81 and 2.82 present normalized compressive strength results for unstressed specimens tested cold 

(residual) and at temperature (hot), respectively, that are based on results from several investigators. Also shown in 

the figure are upper and lower bounds that were established from the data. The results were utilized as part of a 

study to investigate the behavior of reinforced concrete sections at elevated temperature.

Additional information is presented in Figure 2.83 on the effect of elevated-temperature exposure on the 

compressive strength of unsealed nuclear power plant prestressed concrete pressure vessel-type concretes 

(limestone, basalt, or gravel aggregate materials) tested either hot (H) or cold (C). This figure plus the previous two 

figures indicate the influence of the concrete and the test condition on the residual compressive strength after 

thermal exposure. The general trend for a strength loss with increasing temperature indicates the influence of the 

cement paste and the increasing role of the aggregate materials at higher temperatures. Factors have been identified 

that may contribute to the general trend for loss of compressive strength with increasing temperature [2.8]:

aggregate damage; weakening of the cement paste-aggregate bond; and weakening of the cement paste due to an 

increase in porosity on dehydration, partial breakdown of the C-S-H, chemical transformation on hydrothermal 

reactions, and development of cracking. A number of material and environmental-related factors affect the response 

of concrete materials to elevated-temperature conditions (e.g., see Table 2.1, p. 4). As many of the aggregate 

materials are thermally stable up to temperatures of 300°C to 350°C, which includes the temperature range 

considered for most concrete applications, the compressive strength of concrete at elevated temperature is dependent 

in large measure on the interaction between the cement paste and aggregate.

Figure 2.81 Effect of temperature exposure on compressive strength of concrete: tested cold.

Source: G. N. Freskakis et al., ―Strength Properties of Concrete at Elevated Temperature,‖ Civil Engineering Nuclear 

Power, Vol. 1, ASCE National Convention, American Society of Civil Engineers, Boston, Massachusetts, April 

1979.
62

Figure 2.82 Effect of temperature exposure on compressive strength of concrete: tested hot.

Source: G. N. Freskakis et al., ―Strength Properties of Concrete at Elevated Temperature,‖ Civil Engineering 

Nuclear Power, Vol. 1, ASCE National Convention, American Society of Civil Engineers, Boston, 

Massachusetts, April 1979.

Figure 2.83 Effect of temperature on uniaxial compressive strength of PCPV unsealed concretes 

(H = hot, C = cold).

Source: G. A. Khoury, Performance of Heated Concrete—Mechanical Properties, Contract NUC/56/3604A with 

Nuclear Installations Inspectorate, Imperial College, London, United Kingdom, August 1996.
63

Data on effects of elevated temperature on concrete compressive strength have also been assembled as part of a 

review of concrete properties for prestressed concrete pressure vessels (PCPVs) [2.103]. Figure 2.84 presents a 

summary of data assembled under this activity for sealed and unsealed specimens that were either tested at 

temperature (hot) or after returning to room temperature (cold). Results indicate that the reference strength retention 

of unsealed specimens (open symbols) was better than that for sealed specimens (closed symbols). It was noted that 

this may be a concern for mass concrete having substantial amounts of retained moisture that experiences thermal 

loadings and that deleterious effects of elevated-temperature exposure on concrete compressive strength can be 

considered as significant only for sealed conditions in which significant amounts of free moisture are retained up to 

300 ̊C. Compressive strength was found to decrease with increases of temperature, time of exposure, free-moisture 

content, and thermal cycling. 

Figure 2.84 Compilation of data on concrete compressive strength vs temperature.

Source: R.K. Nanstad, A Review of Concrete Properties for Prestressed Concrete Pressure Vessels, ORNL/TM5497, Oak Ridge National Laboratory, Oak Ridge, Tennessee, October 1976.

Results from information presented above as well as those from other researchers have been used to develop a series 

of figures illustrating the effect of elevated temperature on the compressive strength of concrete materials. 

References to data sources cited by others or utilized in figures are identified in Appendix A. Test results are 

presented in terms of testing procedures identified earlier in the report that are commonly used to evaluate concrete 

response to elevated temperature: unstressed tests (hot testing), unstressed residual (cold testing), and stressed tests. 

All test results presented are for unsealed specimens.

Figure 2.85 presents a compilation of data on the effect of elevated temperature on the relative compressive strength 

(i.e., unsealed, hot testing) for a number of different concrete mixtures (i.e., ordinary Portland cement concretes and 

concretes that also included supplementary cementitious materials). The CEB design curve for unstressed normal 

weight concretes has been superimposed on the figures to facilitate comparisons. 

 

Comites Euro-International Du Beton, Fire Design of Concrete Structures – in Accordance with CEB/FIP Model 

Code 90, CEB Bulletin D’Information No. 208, Lausanne, Switzerland, July 1991.
64

Figure 2.85 Compilation of data on relative compressive strength vs temperature –

ordinary Portland cement concretes and concretes containing supplementary cementitious materials.

Results obtained for only the Portland cement concretes are provided in Figure 2.86 with the results subdivided into 

normal strength concretes (fc’ < 60 MPa) and high-strength concretes (fc’ > 60 MPa) and presented in Figures 2.87 

and 2.88, respectively. 

Figure 2.86 Compilation of data on relative compressive strength vs

temperature – ordinary Portland cement concretes.
65

Figure 2.87 Compilation of data on relative compressive strength vs

temperature – ordinary Portland cement concretes, fc’ < 60 MPa.

Figure 2.88 Compilation of data on relative compressive strength vs

temperature – ordinary Portland cement concretes, fc’ > 60 MPa.

Results obtained for Portland cement concretes containing supplementary cementious materials are presented in 

Figure 2.89.
66

Figure 2.89 Compilation of data on relative compressive strength vs temperature –

concretes containing supplementary cementitious materials.

Several general conclusions can be derived based on the relative compressive strength data (unsealed, hot testing). 

Considerable data scatter is exhibited by the relative compressive strength data particularly at the lower exposure 

temperatures. Scatter can probably be attributed to different materials and testing conditions utilized in the different 

studies (e.g., constituents and mix proportions, specimens size and shape, specimens tested at different degrees of 

hydration and initial moisture contents, and heating rates and thermal stabilization periods). As the exposure 

temperature increased the loss in relative compressive strength tended to become more linear. Although ordinary 

Portland cement high-strength concrete data reported are very limited, there may be a slight trend for the highstrength concretes to have increased rates of strength lose at temperatures lower than 400 ̊C than the normal-strength 

concretes, but at higher temperatures the rates of strength loss appear to be similar. Results for concretes containing 

supplementary cementitious materials, which with one exception were all high-strength concretes, also tended to 

show higher rates of compressive strength loss at temperatures lower than 400 ̊C than that experienced by the 

ordinary Portland cement concrete mixes.

Figure 2.90 presents a compilation of data on the effect of elevated temperature on the residual compressive strength 

(i.e., unsealed, cold testing) for a number of different concrete mixtures (i.e., ordinary Portland cement concretes and 

concretes that also included supplementary cementitious materials). Results obtained for only the Portland cement 

concretes are provided in Figure 2.91 with the results subdivided into normal strength concretes (fc’ < 60 MPa) and 

high-strength concretes (fc’> 60 MPa) and presented in Figures 2.92 and 2.93, respectively. Residual compressive 

strength results for ordinary Portland cement normal-strength concretes utilizing carbonate- or siliceous- and graveltype aggregate materials are presented in Figures 2.94 and 2.95, respectively. Residual compressive strength results 

for concretes containing supplementary cementious materials are presented in Figure 2.96, with the results 

subdivided into normal-strength concretes (fc’ < 60 MPa) and high-strength concretes (fc’> 60 MPa) and presented 

in Figures 2.97 and 2.98, respectively. Residual compressive strength results for concretes containing silica fume, 

fly ash, metakaolin, and slag as supplementary cementitious materials are presented in Figures 2.99 through 2.102, 

respectively. The CEB design curve for unstressed normal weight concretes has been superimposed on the figures 

to facilitate comparisons.
67

Figure 2.90 Compilation of data on residual compressive strength vs temperature –

ordinary Portland cement concretes and concretes containing supplementary cementitious materials.

Figure 2.91 Compilation of data on residual compressive strength vs temperature –

ordinary Portland cement concretes.

88.1-10%S- S-O-#40

Coarse aggregate type

Basalt (B)

Carbonate (C)

Exp. Shale (ESH)

Exp. Slate (ESL)

Firebrick (F) Granite (GT)

Gravel (G)

Limestone (L)

Lytag (LY)

Pumice (P) Quartz (Q)

Siliceous (S)

None (O)

Fly ash (F) Metakaolin (MK)

Silica fume (SF)

Slag (S)

Supplementary

Cementitious Material
68

Figure 2.92 Compilation of data on residual compressive strength vs

temperature – ordinary Portland cement concretes, fc’ < 60 MPa.

Figure 2.93 Compilation of data on residual compressive strength vs

temperature – ordinary Portland cement concretes, fc’ > 60 MPa.
69

Figure 2.94 Compilation of data on residual compressive strength vs temperature –

ordinary Portland cement concretes with carbonate-type aggregates, fc’ < 60 MPa.

Figure 2.95 Compilation of data on residual compressive strength vs temperature –

ordinary Portland cement concretes with siliceous- or gravel-type aggregates, fc’ < 60 MPa.
70

Figure 2.96 Compilation of data on residual compressive strength vs temperature –

concretes containing supplementary cementitious materials.

Figure 2.97 Compilation of data on residual compressive strength vs temperature –

concretes containing supplementary cementitious materials, fc’ < 60 MPa.
71

Figure 2.98 Compilation of data on residual compressive strength vs temperature –

concretes containing supplementary cementitious materials, fc’ > 60 MPa.

Figure 2.99 Compilation of data on residual compressive strength vs temperature –

concretes containing silica fume supplementary cementitious materials.
72

Figure 2.100 Compilation of data on residual compressive strength vs temperature –

concretes containing fly ash supplementary cementitious materials.

Figure 2.101 Compilation of data on residual compressive strength vs temperature –

concretes containing metakaolin supplementary cementitious materials.
73

Figure 2.102 Compilation of data on residual compressive strength vs temperature –

concretes containing slag supplementary cementitious materials.

Several general conclusions can be derived based on the residual compressive strength data (unsealed, cold testing). 

Considerable data scatter is exhibited by the residual compressive strength data, particularly at the lower exposure 

temperatures, which can be attributed to differences in material and environmental factors for the different tests. 

Comparing the residual (Figure 2.91) to the relative (Figure 2.86) compressive strength results indicates that in 

general the compressive strength retention of specimens tested at temperature is greater than that obtained from 

specimens permitted to cool to room temperature prior to testing. Results in Figure 2.91 seem to indicate that for the 

unsealed specimens the residual compressive strength generally tends to decrease somewhat at temperatures below 

about 100 ̊ to 120 ̊C, that can be followed by an increase in compressive strength as the temperature increases to 

200 ̊ to 250 ̊C, before again decreasing, with the main decrease occurring as the temperature exceeds about 300 ̊C. 

No definite trend could be derived on the effect of initial concrete strength when comparing normal- and highstrength ordinary Portland cement concrete results (Figures 2.92 and 2.93). Although data are limited, there appears 

to be a slight trend for the siliceous and gravel aggregate-type concretes (Figure 2.95) to exhibit higher residual 

compressive strengths than carbonate aggregate-type concretes (Figure 2.94) at temperatures below about 400 ̊C, but 

at higher temperatures the carbonate-aggregate type concretes may exhibit higher residual compressive strengths. 

Residual compressive strength results for concretes containing supplementary cementitious materials (Figure 2.96) 

exhibited a similar trend to that exhibited by the ordinary Portland cement concretes (Figure 2.91). Comparing 

Figure 2.97 with Figure 2.98 indicates that the residual compressive strength retention of the normal-strength 

concretes at temperatures above 200 ̊C may be slightly better than that for the high-strength concretes, but the data 

are limited. Although results are limited, Figures 2.99 through 2.102 indicates that the residual compressive strength 

retention is highest for the concretes containing slag and fly ash supplementary cementitious materials, followed by 

the concrete containing metakaoline, with silica fume concretes exhibiting the lowest residual compressive strength 

at the higher temperatures.

The effect of sustained stress during heating on the relative compressive strength (unsealed, hot testing) is presented 

in Figure 2.103 for ordinary Portland cement concretes and concretes that also included supplementary cementitious 

materials. Preload levels ranged from 20 to 40% the reference room temperature compressive strength of the 

concretes. Results have been separated into those for ordinary Portland cement concrete materials and those for 

concretes containing supplementary cementitious materials and are presented in Figures 2.104 and 2.105, 

respectively. Although results are limited, the presence of preload, within reasonable limits, and its beneficial 

effects relative to retained compressive strength is evident. Similar effects of pre-loading on the residual 

compressive strength were noted to that obtained for the ordinary Portland cement concretes, but data were too 

limited and scatter too great to draw conclusions comparing the relative benefit of preload on ordinary Portland 

cement concretes and concretes containing supplementary cementitious materials. It has been noted that the 

beneficial effects on retained compressive strength derived from preload during heating is a result of densification of 
74

the cement paste resulting in a large transient creep component and possibly to a reduction in porosity relative to the 

unloaded state, and pre-compression can reduce tensile stresses in the concrete, particularly during cooling [2.8].

Figure 2.103 Compilation of data on compressive strength vs temperature for preloaded specimens – ordinary 

Portland cement concretes and concretes containing supplementary cementitious materials.

Figure 2.104 Compilation of data on relative compressive strength vs temperature

for preloaded specimens – ordinary Portland cement concretes.
75

Figure 2.105 Compilation of data on relative compressive strength vs temperature

for pre-loaded specimens – concretes containing supplementary cementitious materials.

The effect of cooling regime on the residual compressive strength has been studied as part of an overall investigation 

to evaluate the post-thermal properties of concrete following rapid cooling (e.g., water quenching in conjunction 

with fire) [2.67]. A granite aggregate ordinary Portland cement concrete having an initial reference compressive 

strength of 21.16 MPa was used to fabricate 101.6-mm-diameter by 203.2-mm-long cylindrical test specimens that 

were cured for eight weeks in a fog room (23 ̊C and 93% relative humidity). The specimens were heated in a 

furnace to 200 ̊, 400 ̊, 600 ̊, or 800 ̊C with a holding time at temperature of four hours. Following heating, the 

specimens were cooled slowly (1 ̊C/minute), naturally (left in furnace and permitted to cool slowly), or rapidly (e.g., 

specimen removed from furnace and placed into a tank of water initially at 20 ̊C). Figure 2.106 presents the effect 

of cooling regime on the relative residual compressive strength for concretes subjected to temperatures up to 800 ̊C.

Figure 2.106 Effect of cooling regime on relative residual compressive strength following thermal exposure.

Source: J. Lee, Y. Xi, and K. Willam, ―Concrete Under High Temperature Heating and Cooling,‖ Report SESM 

No.09/2006, Department of Civil, Environmental, and Architectural Engineering, University of Colorado at 

Boulder, 2006.
76

The effect of different cooling regimes ranging from natural cooling to water spraying to water quenching on the 

relative residual concrete compressive strength has been investigated [2.117]. The limestone aggregate concrete 

utilized in the study contained silica fume, 10.7% (by weight) of the cementitious material, and had a reference 

compressive strength of 83.5 MPa. The 100 x 100 x 100 mm3specimens were cured in water at 20 ̊C until an age of 

58 days and then dried at 105 ̊C. After drying the specimens were exposed to temperatures either of 200 ̊, 400 ̊, 

600 ̊, or 800 ̊C with the temperature maintained for one hour. After elevated temperature the specimens were 

exposed to one of several cooling regimes: natural cooling, five-minute spraying with water, 30-minute spraying 

with water, 60-minute spraying with water, or quenching in water. When the specimens returned to room 

temperature they were tested. The effect of the cooling regimes on the relative residual compressive strength is 

presented in Figure 2.107. The cooling regime results show the sensitivity of the relative residual compressive 

strength to water spraying and quenching, particularly in the lower temperature range. The effect of quenching on 

the relative residual compressive strength decreases somewhat at the higher temperatures, however, the relative 

residual compressive strength already had been severely reduced by exposure to these temperatures.

Figure 2.107 Effect of different cooling regimes on 

relative residual compressive strength following thermal exposure.

Source: G-F. Peng, S-H. Bian, Z-Q. Guo, J. Zhao, X-L. Peng, and Y-C. Jiang, ―Effect of Thermal Shock Due to 

Rapid Cooling on Residual Mechanical Properties of Fiber Concrete Exposed to High Temperatures,‖ 

Construction and Building Materials 22, pp. 948-955, 2008.

Lightweight and Thermally Stable Aggregate Concretes

Relative compressive strength results (unsealed, hot testing) for lightweight/thermally stable aggregate concretes are 

presented in Figure 2.108. Relative compressive strength results for lightweight/thermally stable aggregate 

concretes fabricated using ordinary Portland cement as the binder and ordinary Portland cement plus supplementary 

cementitious materials as the binder are presented in Figures 2.109 and 2.110, respectively. The CEB design curve 

for lightweight concrete is superimposed on each of the figures. Residual compressive strength results (unsealed, 

cold testing) for lightweight/thermally stable aggregate concretes are presented in Figure 2.111. Residual 

compressive strength results for lightweight/thermally stable aggregate concretes fabricated using ordinary Portland 
77

cement as the binder and ordinary Portland cement plus supplementary cementitious materials as the binder are 

presented in Figures 2.112 and 2.113, respectively. Specimens tested at temperature (relative strength) in general 

appear to retain a higher percentage of their compressive strength than that retained by specimens permitted to cool 

to room temperature prior to testing (residual strength). No definite conclusions could be drawn when comparing 

results for specimens fabricated using ordinary Portland cement as the binder and specimens fabricated using 

ordinary Portland cement plus supplementary cementitious materials as the binder.

Figure 2.108 Compilation of data on relative compressive strength vs temperature 

for lightweight/insulating concrete specimens – ordinary Portland cement concretes 

and concretes containing supplementary cementitious materials.

Figure 2.109 Compilation of data on relative compressive strength vs temperature 

for lightweight/insulating concrete specimens – ordinary Portland cement concretes. 
78

Figure 2.110 Compilation of data on relative compressive strength vs temperature 

for lightweight/insulating concrete specimens – concretes containing supplementary cementitious materials.

Figure 2.111 Compilation of data on residual compressive strength vs temperature 

for lightweight/insulating concrete specimens – ordinary Portland cement concretes 

and concretes containing supplementary cementitious materials.
79

Figure 2.112 Compilation of data on residual compressive strength vs temperature 

for lightweight/insulating concrete specimens – ordinary Portland cement concretes.

Figure 2.113 Compilation of data on residual compressive strength vs temperature 

for lightweight/insulating concrete specimens – concretes containing supplementary cementitious materials.
80

Fibrous Concretes

Relative compressive strength results (unsealed, hot testing) for fiber-reinforced concretes (fibrous concrete) are 

presented in Figure 2.114. Residual compressive strength results for fibrous concretes are presented in Figure 2.115. 

At temperatures above about 200 ̊C the retained compressive strength for the specimens tested at temperature and 

specimens permitted to cool to room temperature prior to testing both tended to decrease linearly as the temperature 

increased. Insufficient data are available, especially for relative compressive strength testing, to draw any 

conclusions when comparing relative and residual compressive strength results.

Figure 2.114 Compilation of data on relative compressive strength vs temperature – fibrous concrete. 

Figure 2.115 Compilation of data on residual compressive strength vs temperature – fibrous concrete. 

Some information is available on the effect of different cooling regimes ranging from natural cooling to water 

spraying to water quenching on the relative residual concrete compressive strength of fibrous concretes [2.117]. 

The limestone aggregate concretes utilized in the study contained silica fume, 10.7% (by weight) of the cementitious 

material, and had a reference compressive strengths of 88.6, 83.4 MPa, and 87.2, respectively, for mixes containing 

 

Companion results for plain concrete subjected to the same cooling regimes were presented earlier in Figure 2.107.
81

0.1% polypropylene fibers, 0.3% polypropylene fibers, and a combination of 0.07% polypropylene and 1% steel 

fibers (by volume). The 100 x 100 x 100 mm3specimens were cured in water at 20 ̊C until an age of 58 days and 

then dried at 105 ̊C. After drying the specimens were exposed to temperatures of 200 ̊, 400 ̊, 600 ̊, or 800 ̊C with 

the temperature maintained for one hour. After elevated temperature exposure the specimens were subjected to one 

of several cooling regimes: natural cooling, five-minute spraying with water, 30-minute spraying with water, 

60-minute spraying with water, or quenching in water. When the specimens returned to room temperature they were 

tested. The effect of the cooling regimes on the relative residual compressive strengths is presented in Figures 2.116 

to 2.118 for the three fibrous concrete materials. Also shown in the figures are companion splitting-tensile strength 

results that were determined as part of the cooling regime study. Natural cooling retained the highest compressive 

and tensile strength values for each of the fibrous concretes with water quenching and 30 minute spraying producing 

the lowest strength retention. The hybrid concrete mix containing polypropylene and steel fibers exhibited 

improved retention of tensile strength relative to the polypropylene-reinforced concretes while still providing 

resistance to spalling. Notched beams tested as part of this study indicated that the hybrid fiber reinforced concrete 

also exhibited significant improvement in fracture energy at room temperature and after elevated temperature 

exposure compared to companion plain concrete beams.

Figure 2.116 Effect of different cooling regimes on residual compressive and tensile strengths 

of 0.1 % polypropylene fiber reinforced concrete – 88.6 MPa compressive strength.

Source: G-F. Peng, S-H. Bian, Z-Q. Guo, J. Zhao, X-L. Peng, and Y-C. Jiang, ―Effect of Thermal Shock Due to 

Rapid Cooling on Residual Mechanical Properties of Fiber Concrete Exposed to High Temperatures,‖ 

Construction and Building Materials 22, pp. 948-955, 2008.

Figure 2.117 Effect of different cooling regimes on residual compressive and tensile strengths 

of 0.3% polypropylene fiber reinforced concrete – 83.4 MPa compressive strength.

Source: G-F. Peng, S-H. Bian, Z-Q. Guo, J. Zhao, X-L. Peng, and Y-C. Jiang, ―Effect of Thermal Shock Due to 

Rapid Cooling on Residual Mechanical Properties of Fiber Concrete Exposed to High Temperatures,‖ 

Construction and Building Materials 22, pp. 948-955, 2008.
82

Figure 2.118 Effect of different cooling regimes on residual compressive and tensile strengths 

of 0.07% polypropylene plus 1% steel fiber reinforced concrete – 87.2 MPa compressive strength.

Source: G-F. Peng, S-H. Bian, Z-Q. Guo, J. Zhao, X-L. Peng, and Y-C. Jiang, ―Effect of Thermal Shock Due to 

Rapid Cooling on Residual Mechanical Properties of Fiber Concrete Exposed to High Temperatures,‖ 

Construction and Building Materials 22, pp. 948-955, 2008.

Examples of Relations for Concrete Compressive Strength at Elevated Temperature. The following has been 

proposed for concrete cube compressive strength as a function of temperature [2.107]:

(2.44)

where and are the concrete cube compressive strength at elevated temperature and room temperature, 

respectively. 

A relationship has been developed for estimating the peak concrete compressive strength at temperature ( ) based 

on temperature (T) and peak compressive strength at room temperature ( ) of 25 MPa [2.72]:

. (2.45)

A relationship has been developed for the normalized residual compressive strength of normal strength concretes 

( < 50 MPa) following elevated temperature exposure [2.108]:

for 20 ̊ < T < 200 ̊C, and (2.46)

for 200 ̊ < T < 800 ̊C (2.47)

where is the residual compressive strength, T is the temperature, and is the reference room temperature 

strength.

The following has been proposed by the American Society of Civil Engineers in their fire protection manual [2.87]:

for 20 ̊ < T < 450 ̊C, (2.48)
83

for 450 ̊ < T < 874 ̊C, and (2.49)

for T > 874 ̊C (2.50)

where is the residual compressive strength, T is the temperature, and is the reference room temperature 

strength.

The following has been proposed for high strength concrete ( > 55 MPa) residual compressive strength at 

elevated temperature [2.118]:

for T < 100 ̊C, (2.51)

for 100 ̊ < T < 400 ̊C, and (2.52)

for T > 400 ̊C (2.53)

where is the compressive strength at room temperature, T is temperature of interest, and is the strength at 

temperature T in MPa.

Summary. From about 22 ̊ to 120 ̊C the concrete compressive strength decrease is attributable to thermal swelling

of the physically-bound water that causes disjoint pressures. From 120 ̊C to about 300 ̊C there is a regain of 

compressive strength that is generally attributed to greater van der Waal’s forces as a result of the cement gel layers 

moving closer to each other during heating. At temperatures above 300 ̊C compressive strength losses can become 

significant probably due to differences in thermal expansion coefficients between the aggregate and cement paste 

and decomposition of calcium hydroxide. At temperatures above 450 ̊C concrete compressive strength drops 

significantly due to loss of bond between the aggregate and cement paste. Residual concrete compressive strength 

(cold testing) is generally lower than relative concrete compressive strength (hot testing). Moisture content at time 

of testing has a significant effect on the strength of concrete at elevated temperature with strength of unsealed 

specimens being higher than strength of sealed specimens. Some regain in compressive strength can occur if 

specimens are stored in water following thermal exposure. Results in the literature indicate that the original 

concrete strength of normal strength concrete, type of cement, aggregate size, heating rate, and water-cement ratio 

have little effect on the relative strength vs temperature characteristics; exposure times at temperatures beyond 1 h 

had an effect on residual compressive strength, but this effect diminished as the level of exposure temperature 

increased, with the majority of strength loss occurring in the first 2 h. Age of concrete is important in so far as 

concretes with relatively incomplete hydration of cement may indicate a strength increase for temperature up to 

400°C due to accelerated hydration; maximum test temperature influences the strength recovery in that after 

exposures to above about 600°C there is no strength recovery; aggregate-cement ratio has a significant effect on 

strength of concrete exposed to high temperature with the reduction being proportionally smaller for lean mixtures 

than for rich mixtures. Type of aggregate appears to be one of the main factors influencing concrete strength at high 

temperature with siliceous aggregate concrete having lower strength (by percentage) at high temperature than 

calcareous and lightweight concrete; and stressed specimens resulted in higher compressive strength retention at 

high temperature than unstressed specimens. 

2.2.1.5 Thermal Cycling

Data and Information. Thermal cycling, even at relatively low temperatures (65°C), can have some deleterious 

effects on concrete’s mechanical properties (i.e., cyclic heating generally gives lower strengths than a single heating) 

[2.119]. Figure 2.119 presents the affect of temperature cycling on a limestone aggregate concrete [2.120] and in 
84

Figure 2.120 for a sealed limestone concrete [2.65]. Results indicate that the compressive, tensile, and bond 

strengths, and the modulus of elasticity are reduced, and the Poisson’s ratio increased. As shown in Figure 2.120, 

the sealed limestone concrete exhibited a 50% loss of strength after 14 cycles at 150 ̊C. At higher temperatures 

(200 ̊ to 300 ̊C), the first thermal cycle causes the largest percentage of damage, with the extent of damage markedly 

dependent on aggregate type and is associated with loss of bond between aggregate and cement paste matrix [2.120].

Figure 2.119 Effect of temperature cycles on limestone concretes.

Source: D. Campbell-Allen and P. M. Desai, ―The Influence of Aggregate on the Behavior of Concrete at Elevated 

Temperature,‖ Nuclear Engineering and Design 6(1), pp. 65–77, August 1967.
85

Figure 2.120 Influence of thermal cycling on - response of sealed concrete tested at 149°C.

Source: V. V. Bertero and M. Polivka, ―Influence of Thermal Exposures on Mechanical Characteristics of 

Concrete,‖ Paper SP 34-28 in Concrete for Nuclear Reactors, pp. 505-531, American Concrete 

Institute, Farmington Hills, Michigan, 1972.

2.2.1.6 Tensile Strength

Information and Data. The tensile strength of concrete is important because it determines the ability of concrete to 

resist cracking. At room temperature, concrete’s tensile strength generally is from 7 to 11% its compressive 

strength. Direct measurement of concrete’s tensile strength is seldom made because of difficulties in gripping the 

specimen to apply loads. An indication of concrete’s tensile strength can be obtained by the splitting-tension and 

flexure tests. The splitting-tension test is an indirect test for tensile strength of concrete in that a horizontal concrete 

cylinder is loaded in compression through bearing strips placed along two axial lines that are diametrically opposite 

on the specimen [2.121]. Flexural strength of concrete is expressed in terms of modulus of rupture that is 

determined from beam specimens loaded in four-point bending until failure occurs. Because the modulus of rupture 

is calculated based on linear-elastic conditions, it is a fictitious value, but convenient for comparison purposes. For 

normal strength concretes tested at room temperature, the modulus of rupture is 60 to 100% higher than the direct 

tensile strength and 100 to 133% the splitting-tension strength [2.122]. Most tests to determine elevated-temperature 

effects on concrete tensile strength used splitting-tension tests with the residual tensile strength determined. As 

noted in Figure 2.121, the effect of elevated temperature on tensile strength shows a similar trend to its effect on 

compressive strength, but tensile testing of concrete is more sensitive to deterioration at elevated temperature [2.50].

The effect of elevated temperature on the splitting-tensile and direct tension strengths of 10-cm-diameter by 20-cmlong cylindrical test specimens fabricated from ordinary Portland and high alumina cements and different aggregate

materials (i.e., sandstone, andesite, limestone, and pumice) was investigated [2.54]. Figure 2.122 presents the effect
86

Figure 2.121 Comparison of effect of elevated temperature on compression and tensile strengths

of concretes fabricated using different types of conventional aggregate materials.

Source: R. Blundell, C. Diamond, and R. Browne, ‖The Properties of Concrete Subjected to Elevated 

Temperatures,‖ Report No. 9, Construction Industry Research and Information Association, Underwater 

Engineering Group, London, United Kingdom, June 1976.

of short term exposures (1 hour) to temperatures from 100 ̊ to 800 ̊C followed by gradual cooling to room 

temperature prior to testing on the reduction in strength and modulus of elasticity relative to room temperature 

reference values. All the concrete mixes using ordinary Portland cement exhibited similar compressive strength 

ratios (i.e., general linear reduction with temperature increase becoming approximately 60% at 400 ̊C) except for the 

pumice concrete that retained about 80% its reference strength at 400 ̊C. The tensile strength ratio of the silica 

aggregate concrete was close to that exhibited with respect to compressive strength ratio but the tensile strength ratio 

of the limestone aggregate concrete exhibited a fairly large reduction being about 30% at 400 ̊C. The modulus of 

elasticity ratio also decreased with increasing temperature with the rate of reduction increasing in order of andesite, 

silica, and limestone aggregate concretes. The pumice aggregate maintained a higher modulus ratio at elevated 

temperature than that for the other aggregate concretes, with the high alumina cement silica aggregate concrete 

performing the worst. Tensile strength results as a function of exposure period relative to 28-day values for 

specimens cured at room temperature and residual tensile strength results for specimens subjected to elevated

temperature for exposure times from 3 to 12 months at temperatures of 40 ̊ (direct tension), 60 ̊ (direct tension), or 

80 ̊C (splitting-tension) using specimens that were either sealed or unsealed during exposure are shown in 

Figure 2.123. The specimens were fabricated from ordinary Portland cement, river gravel, and had a water-cement 

ratio of 0.50. The tensile strength for unsealed specimens exhibited a steep drop after 3 months exposure, with the 

drop being more severe than that obtained for companion compression tests, however the change between 3 and 

12 months was not great. Sealed specimens cured at 60 ̊C exhibited about a 10% loss in strength in direct tension, 

while sealed specimens cured at 80 ̊C exhibited an increase in splitting tensile strength of about 20% at 3 months, 

but decreased with increasing curing time at temperature.

Residual splitting-tensile-strengths were determined at temperatures of 300°C and 600°C for unsealed specimens 

[2.110]. The specimens were fabricated either from normal Portland cement (N), fly ash cement (F), moderate heat 

Portland cement (M), blast furnace slag cement (B), or high alumina cement (A). Aggregates investigated included 

a hard sandstone, basalt, limestone, blast furnace slag, and fire-resistant bricks. Water-cement ratios utilized were 

either 0.45, 0.55, or 0.65. The 150-mm-diameter by 150-mm-long cylindrical test specimens were cured for 91 days 

in water prior to heating. Residual results for splitting-tensile, as well as compressive strength, and modulus of 

elasticity, are summarized in Figure 2.124. The ratio of results after heating to reference room temperature results 

indicates that the residual ratios decreased with the increase of heating temperature (e.g., results range from 

approximately 20 to 90% at 300 ̊C and between approximately 20 and 30% at 600 ̊C). The high alumina cement 

mix had a notably lower residual strength ratio (~23%) than that for the other cementitious materials which 

exhibited a similar effect (~66 to 87%), however there was little difference in tensile strength ratio for all 

cementitious materials at 600°C. No significant variation in residual tensile strength ratio resulted for the different 

water-cement ratios.
87

Figure 2.122 Effect of elevated temperature curing on residual compressive and tensile strengths and modulus.

Source: T. Harada, T. Takeda, S, Yamane, and F, Furumura, ―Strength, Elasticity and Thermal Properties of 

Concrete Subjected to Elevated Temperatures,‖ Paper SP-34-21in Concrete for Nuclear Reactors, pp. 377-

406, American Concrete Institute, Farmington Hills, Michigan, 1972.

Figure 2.123 Effect of length of room temperature and elevated temperature curing 

on relative residual tensile strength results for sealed and unsealed specimens.

Source: T. Harada, T. Takeda, S, Yamane, and F, Furumura, ―Strength, Elasticity and Thermal Properties of 

Concrete Subjected to Elevated Temperatures,‖ Paper SP-34-21 in Concrete for Nuclear Reactors, pp. 377-

406, American Concrete Institute, Farmington Hills, Michigan, 1972.
88

Figure 2.124 Relationship between residual ratios and heating temperatures, cement types, and water-cement ratios.

Source: K. Nagao and S. Nakane, ―Influences of Various Factors on Physical Properties of Concretes Heated to 

High Temperatures,‖ Paper H03/1 in 11th International Conference on Structural Mechanics in Reactor 

Technology, pp. 61–66, August 1991.

An investigation was undertaken to determine the changes in mechanical properties of concrete after long-term 

exposure to sustained temperatures ranging from 75 ̊ to 600 ̊C for periods up to eight months [2.123]. Three types 

of concrete made with normal Portland cement (Type I), normal Portland cement with 35% replacement with blast 

furnace slag (Type II), and normal Portland cement with 25% replacement with fly ash (Type III) were investigated. 

Dolomitic limestone was used as the coarse aggregate material. Water-cement ratios ranged from 0.45 to 0.60 (by 

weight). After moist curing for 28 days the specimens (except for reference specimens) were stored under 

laboratory conditions at 50% relative humidity for periods from 16 and 30 weeks. After thermal exposure, the 

specimens were permitted to cool to room temperature prior to testing. Figure 2.125 presents residual splittingtensile strength results relative to reference room temperature values after one-month exposure to elevated 

temperatures up to 600 ̊C. Series A and B, Series C and D, and Series E and F used cementitious materials Type I, 

Type II, and Type III, respectively. The first series in each group used a water-cement ratio of 0.45 (by weight) and 

the second series in each group used a water-cement ratio of 0.60 (by weight). After one-month’s exposure at 75 ̊C, 

the splitting-tensile strength of the 102-mm-diameter by 203-mm-long cylindrical test specimens fabricated with 

normal cement and slag was relatively unaffected, however, the normal Portland cement concrete and the normal 

Portland cement plus fly ash concrete residual splitting-tensile strengths were reduced by about 10%. Test results at 

this temperature after four and eight month’s exposure were similar. Concretes exposed to 150 ̊C for one month 

produced a decrease in residual splitting-tensile strength from 10% for normal Portland cement concrete with slag to 

a maximum of about 20% for the Portland cement concrete. After four month’s exposure at this temperature 

virtually all specimens had disintegrated. This behavior was attributed to instability of the limestone aggregate due 

to the presence of iron sulfide.

Tests were conducted to evaluate the effect of elevated temperature on the residual direct-tensile and splitting-tensile 

strengths of normal-strength (38.1-MPa) and high-strength (61.1-MPa) calcareous aggregate concretes [2.124]. 

Direct-tension tests utilized prisms 100 x 100 x 400 mm3and splitting-tensile tests were conducted on 160-mmdiameter by 320-mm-long cylindrical test specimens. The specimens were cured at 22 ̊C and 95% relative humidity 

until testing at an age of 2 months. The specimens were slowly heated to target test temperatures of 150 ̊, 300 ̊, 

450 ̊, or 600 ̊ C and maintained at the target temperature for one hour prior to permitting to slowly cool to room 

temperature. Experimental results, shown in Figure 2.126, indicate that the residual tensile strengths for both 

normal-strength and high-strength concretes decreased similarly and almost linearly with increasing temperature. 

Tensile strength of the high-strength concrete remained approximately 10-15% higher than that of the normalstrength concrete over the entire temperature range. Also, tensile strengths measured by the splitting-tension test 

were consistently higher than those obtained by the direct-tension test. 
89

Figure 2.125 Residual splitting-tensile strength of concrete after one month’s exposure at various temperatures.

Source: G. G. Carette, K.E. Painter, and V.M. Malhotra, ―Sustained High Temperature Effects on Concretes Made 

with Normal Portland Cement, Normal Portland Cement and Slag, or Normal Portland Cement and Fly 

Ash,‖ Concrete International 4(7), pp. 41–51, July 1982.

Figure 2.126 Residual tensile strengths of HSC and NSC.

Source: A. N. Noumowe, P. Clastres, G. Debicki, and J-L. Costaz, ―Thermal Stresses and Water Vapor Pressure of 

High Performance Concrete at High Temperature,‖ Proc. 4th International Symposium on Utilization of 

High-Strength/High-Performance Concrete, Paris, 1996.
90

An experimental study has been conducted to investigate the deterioration of concretes used in the construction of 

prestressed concrete pressure vessels in the United Kingdom [2.109]. The overall study focused on the dependence 

of the residual properties of thermally-aged concretes involving representative mix designs, temperature exposure 

levels from 20 ̊ to 450 ̊C, test ages from 3 months to 1 year, and exposure condition (sealed or unsealed). Three 

prestressed concrete pressure vessel concretes were investigated. Ordinary Portland cement or a combination of 

ordinary Portland cement and fly ash were used as the cementitious materials. Coarse aggregates included basalt, 

flint gravel, and dolomitic limestone having a maximum aggregate size up to 20 mm. Water-cement ratios utilized 

were either 0.43, 0.56, or 0.58. Sealed specimens subjected to temperatures below 100 ̊C were placed directly in 

water in a heat-controlled tank and sealed specimens tested at temperature above 100 ̊C were heated in an autoclave. 

Residual splitting-tensile strength results for sealed and unsealed specimens cast from a flint gravel aggregate 

concrete (i.e., w/c = 0.43 and ordinary Portland cement) after curing ages of 3 months and 1 year are presented in 

Figure 2.127. Results for specimens tested after 1 year curing generally exhibited higher residual splitting-tensile 

strength than those that were cured for 3 months either in the sealed or unsealed condition. Unlike the residual 

compressive strength results obtained in the investigation, the tensile strength of sealed concrete specimens that had 

experienced temperature exposure was higher than that for unsealed specimens (except at 150°C). Comparing 

compression and tensile results, the splitting-tensile strength appears to be less sensitive to the effect of moisture 

content (i.e., unsealed specimens did not exhibit a trend of decreasing residual strength with increasing temperature) 

and were more sensitive to the damage (i.e., microcracking) inside the concrete caused by temperature exposure. 

Relating the splitting-tensile strength values obtained after thermal treatment to the reference room-temperature 

values, all three mixes exhibited a trend for the value to decrease with increasing exposure temperature with the 

decrease being more marked than that exhibited by the compressive strength. In the sealed condition the decrease 

was greater for the mix using ordinary Portland cement with flint coarse aggregate than for the mixes using a 

combination of ordinary Portland cement and fly ash with either basalt or dolomitic limestone coarse aggregate. 

Figure 2.127 Effect of curing age and exposure condition on residual splitting-tensile strength

of a flint-gravel aggregate concrete (Solid line = sealed, dashed line = unsealed).

Source: J. Guo and P. Waldron, ―Deterioration of PCPV Concrete,‖ Nuclear Engineering and Design 198, pp. 211–

226, 2000.

Properties of concrete subjected to high temperature were determined for use in design of the reactor buildings at 

―Monju‖ [2.97]. Cylindrical test specimens 10-cm-diameter by 20-cm-long were cast from a hard sandstone 

aggregate concrete mix that used Type B (fly ash) cementitious material and had a water-cementitious materials 

ratio of 0.55 (by weight). Preheating conditions investigated included in water (unsealed), in water (sealed), in air 

(unsealed), and in air (sealed). The heating period started at a concrete age of 91 days and lasted from 1 to 91 days. 

Specimens were heat treated at 175 ̊C. Splitting-tensile strength results for the various conditions are summarized in 

Figure 2.128. Splitting-tensile results indicate that sealed specimens performed better than unsealed specimens. 

Unsealed specimens exhibited about a 20% reduction, on average, during heating.
91

Figure 2.128 Relative variations in splitting-tensile strength.

Source: K. Hirano, K. Ohmatsuzawa, T. Takeda, S. Nakane, T. Kawaguchi, and K. Nagao, ―Physical Properties of 

Concrete Subjected to High Temperature for MONJU,‖ Paper P2-25, Power Reactor and Nuclear Fuel 

Development Corporation, Tokyo, Japan.

The effect of moisture loss at high temperature on the brittleness of concrete was investigated by conducting threepoint bending tests of preheated notched beams 500 x 100 x 100 mm3

having an effective span of 400 mm and a 

notch depth of 50 mm [2.89]. The beams were fabricated from a Type III Portland cement, water-cement ratio of 

0.54, siliceous river sand, and 20-mm maximum size siliceous gravel aggregate. At an age of 14 days the beams 

were heated for different exposure periods up to 168 hours and then permitted to cool naturally in air for a period of 

12 hours prior to testing. After testing, splitting-tensile and compressive strength100-mm cubes were cut from the 

broken ends of the beam specimens and tested. A comparison of the effect of elevated-temperature exposure on 

residual compressive, tensile (splitting-tension), and bend strengths (notched beams) for the siliceous gravel 

concrete exposed to temperatures up to 600°C is presented in Figure 2.129. Results in the figure indicate that the 

residual tensile strength, either splitting-tensile or notched beam, is affected more significantly as the temperature 

increases than the compressive strength. 

Figure 2.129 Comparison of the effect of elevated-temperature exposure on residual compressive (fcu), 

tensile (splitting-tension) (ft), and bend strengths (notched beams) (fr) of siliceous aggregate concrete.

Source: B. Zhang, N. Bicanic, D.J. Pearce, and D.V. Phillips, ―Relationship Between Brittleness and Moisture Loss 

of Concrete Exposed to High Temperatures,‖ Cement and Concrete Research 32, pp. 363–371, 2002.
92

The effect of elevated temperature on the splitting-tensile and bending strength of normal- and high-strength 

concretes has been investigated [2.125]. Specimens were cast from a 20-mm maximum size crushed basalt 

aggregate concrete mix that incorporated Type II cement and fly ash. Water-cement ratio and proportions of cement 

and fly ash were varied to produce three concretes (C40, C60,and C70 having 28-day compressive strengths of 42.5, 

68.0, and 76.0 MPa, respectively). Splitting-tensile strength specimens were 100-mm cubes and bending strength 

specimens were 100 x 100 x 415 mm3. Specimens were cured for 28 days at 20 ̊C and 90% relative humidity and 

then heated to temperatures of 200 ̊, 400 ̊, 600 ̊, 800 ̊, or 1000 ̊C. After the temperature inside the furnace reached 

the target temperature on the specimen surface, the specimens were permitted to cool to room temperature prior to 

testing. Figure 2.130 presents residual splitting-tensile and bending strength results after temperature exposure. The 

splitting-tensile strength results are presented only for the C70 concrete while the bending results are for all three 

concrete mixes. Both the splitting-tensile and bending strengths decreased as the exposure temperature increased. 

The bending strength results for the two higher strength concretes dropped more sharply than that for the normalstrength concrete (C40).

Figure 2.130 Residual splitting-tensile and bending strengths after thermal exposure.

Source: M. Li, C.X. Qin, and W. Sun, ―Mechanical Properties of High-Strength Concrete After Fire,‖ Cement and 

Concrete Research 34, pp. 1001-1005. 2004.

Research related to prestressed concrete pressure vessels has been conducted to investigate the effect of moisture on 

the properties of Portland cement concrete [2.119]. Two concretes were used in the study, a gravel aggregate 

concrete and a limestone aggregate concrete. Type II Portland cement was used in the mixes with water-cement 

ratios of 0.40 and 0.42 for the limestone aggregate and gravel aggregate concretes, respectively. Nominal 28-day 

compressive strength of the two concretes was 41.4 MPa. Flexural strength specimens were 76.2 x 63.5 x 254 mm3. 

Prior to elevated temperature testing the specimens were cured in a fog room for 28 to 200 days. The specimens 

were subjected to temperatures up to 260 ̊C for periods ranging from 75 to 109 days both at atmospheric pressure 

and in an environment in which free moisture was contained in the specimen by a equilibrium saturated steam 

pressure. Flexural strength results for unsealed specimens as a percentage of reference room temperature value are 

presented in Figure 2.131 for both sealed and unsealed specimens. For the unsealed specimens both the gravel 

aggregate and limestone aggregate concretes heated at 79.4 ̊C and tested at temperature or after cooling to room 

temperature exhibited slight increases in flexural strength relative to unheated specimens. All specimens tested at 

121.1 ̊ and 260 ̊C exhibited a loss of flexural strength relative to the unheated specimens. The greatest loss of 

flexural strength occurred in specimens quenched from 260 ̊C prior to testing and in specimens cooled slowly from 

260 ̊C and then resaturated with water prior to testing. For the sealed concrete specimens flexural results are 

presented in Figure 2.132 after autoclaving and cooling to room temperature prior to testing. In general the gravel 

aggregate concrete exhibited a lesser degree of flexural strength loss that that exhibited by the limestone aggregate 

concrete, especially at the higher autoclave temperatures. 
93

Figure 2.131 Effect of test conditions on flexural strength of gravel and limestone concretes, unsealed.

Source: D. T. Lankard, D.L. Birkimer, F.F. Fondriest, and M.J. Snyder, ―Effects of Moisture Content on the 

Structural Properties of Portland Cement Concrete Exposed to Temperatures Up to 500°F,‖ SP-25 

Temperature and Concrete, pp. 59–102, American Concrete Institute, Farmington Hills, Michigan, 1971. 

Figure 2.132 Flexural strength of autoclaved gravel and limestone concretes.

Source: D. T. Lankard, D.L. Birkimer, F.F. Fondriest, and M.J. Snyder, ―Effects of Moisture Content on the 

Structural Properties of Portland Cement Concrete Exposed to Temperatures Up to 500°F,‖ SP-25 

Temperature and Concrete, pp. 59–102, American Concrete Institute, Farmington Hills, Michigan, 1971. 

The effect of elevated temperature on the flexural strength of ordinary Portland cement concretes and Portland 

cement concretes containing fly ash has been investigated [2.126]. Three reference primary mixes (Mix I, Mix II, 

and Mix III) were prepared in which the cement contents (330.9 to 404.4 kg/m3) and water-cement ratios (0.55 to 

0.45, by weight) were varied to produce concretes having 28-day compressive strengths from 28 to 35 MPa. In 

addition to each reference mix, three additional mixes were prepared for each of the reference mixes in which fly 

ash was utilized to replace 10, 20, or 30% (by weight) of the cement content, thus producing a total of 12 concrete 

mixes. The mixes utilized 20-mm maximum size gravel coarse aggregate. Flexural specimens 100 x 100 x 

500 mm3 were fabricated and subjected to elevated-temperature exposures of 100 ̊, 200 ̊, or 250°C for 1-, 2-, or 3-h 

durations. Figure 2.133 a-c presents the variation of maximum flexural strength with temperature for Mix I, Mix II, 

and Mix III control specimens and the three mixes with different partial placements of cement with fly ash exposed 

to elevated temperature for 60 min. Figure 2.133 d-f presents the effect on flexural strength of three different 

exposure periods for the Mix II test series. Conclusions of this study were that the fly ash consistently showed the 

same pattern of flexural behavior (i.e., trend) with temperature as that exhibited by concrete without fly ash for 

thermal exposures up to 250°C; fly ash concrete with fly ash content up to 20% exhibited improved performance 

relative to concrete without fly ash retaining more of its strength; the exposure time had an effect on the residual 

flexural strength of concrete, but the majority of strength loss occurred within the first hour of exposure.
94

Figure 2.133 Variation of flexural strength with temperature and exposure period 

for concrete mixes with and without fly ash. 

Source: M. P. Raju, M. Shobha, and K. Rambabu, ―Flexural Strength of Fly Ash Concrete Under Elevated 

Temperatures,‖ Magazine of Concrete Research 56(2), pp. 83–88, March 2004.

The effect of elevated temperature followed by water quenching on the properties of high-strength concrete 

containing supplementary cementitious materials such as used in construction of high-rise buildings was studied 

[2.105]. Four binder materials were investigated: Mix 1 - 100% general purpose cement, Mix 2 - 62% slag plus 

38% cement, Mix 3 - 20% fly ash plus 80% cement, and Mix 4 - 6.6% silica fume plus 93.4% cement. Crushed 

basalt aggregate having a maximum aggregate size of 10 mm and a water-cement ratio of 0.30 (by weight) was used 

for all the mixes. Table 2.4 presents reference 28-day properties of the mixes. Specimens used for compression and 

tensile tests were 100-mm-diameter by 200-mm-long cylinders. The dynamic modulus of elasticity specimens were 

75 x 75 x 305 mm3prisms. All specimens were cured by storing in water at 20 ̊C for 28 days and then placing into 

the laboratory environment of 20 ̊C and 65% relative humidity for 2 days. The specimens then were heated to 

temperatures of 200 ̊, 400 ̊, 600 ̊, 800 ̊, or 1000 ̊C with a total duration in the furnace of 7 hours (i.e., soak period

Table 2.4 Concrete mix properties at the age of 28 days

Mix Binder Materials

Cylinder

strength,

MPa

Tensile

strength,

MPa

Flexural 

strength,

MPa

Dynamic 

modulus,

GPa

1 100% OPC 61.8 6.58 9.34 46.5

2 62% slag + 38% OPC 40.3 4.28 5.47 39.9

3 20% fly ash + 80% OPC 67.2 5.08 6.94 45.7

4 6.6% silica fume+ 93.4% OPC 76.4 6.93 8.51 55.2

Source: R.S. Ravindrarajah, R. Lopez, and H. Reslan, ‖Effect of Elevated Temperature on Properties of HighStrength Concrete Containing Cement Supplementary Materials,‖ 9

th International Conference on 

Durability of Building Materials and Components, 9 p., Brisbane, Australia, 17-20 March 2002.
95

from 6 to 2 hours depending on maximum temperature). Figure 2.134 presents the normalized residual compressive, 

splitting-tensile, and flexural strength results for specimens subjected to elevated temperature and quenched in water 

prior to testing. The strengths of all concretes dropped significantly as the maximum exposure temperature 

increased. At an exposure temperature of 1000 ̊C all concretes experienced a significant drop in strength 

independent of binder type used. 

Figure 2.134 Normalized residual property results for specimens subjected to

elevated temperature and quenched in water prior to testing.

Source: R.S. Ravindrarajah, R. Lopez, and H. Reslan, ‖Effect of Elevated Temperature on Properties of HighStrength Concrete Containing Cement Supplementary Materials,‖ 9

th International Conference on 

Durability of Building Materials and Components, 9 p., Brisbane, Australia, 17-20 March 2002.

The effect of elevated temperature on the flexural and splitting-tensile strengths of high-strength concrete was 

investigated [2.127]. Four concrete mixes were used in the investigation: ordinary Portland cement having a 28-day 

compressive strength of 54 MPa, and three concrete mixes containing silica fume with different water-cement ratios 

that produced concretes having compressive strengths from 91 to 133 MPa. Granite coarse aggregate material 

having a 10-mm maximum size was used in the mixes to fabricate 100 x 100 x 400 mm3beam flexural specimens 

and 100-mm-diameter by 200-mm-long cylindrical splitting-tensile specimens. The specimens were cured at 27 ̊C 

for 27 days in water and then allowed to cure at 25 ̊C and 60% relative humidity for an additional two months. At 

an age of about 90 days, the specimens were heated to 200 ̊ or 400 ̊C where they remained until the temperature at 

the center of the specimen was within ±10 ̊C of the target temperature (about 2-3 hours from start of heat up). 

Specimens were then slowly cooled by letting the furnace cool to room temperature or quick cooled by removing 

them from the oven and immediately placing into a water bath. Figure 2.135 presents the variation of residual 

flexural strength with temperature for slow and quick cooling. The designations G50, G90, G110, and G130 

correspond to the nominal compressive strengths (grades) of the four concretes investigated. For each grade 

concrete the residual flexural strength was lower at 400 ̊ than 200 ̊C, with the loss in flexural strength increasing as 

the concrete grade increased. Residual flexural strengths were lower for specimens cooled rapidly relative to those 

that were slowly cooled. Figure 2.136 presents the variation of residual splitting-tensile strength with temperature 

for slow and quick cooling. All grades of concrete exhibited increasing loss of residual splitting-tensile strength 

with increasing temperature for both slow and quick cooling. Higher strength concretes (G90, G110, and G130) lost 
96

more residual splitting-tensile strength at elevated temperature than the lower strength concrete (G50). Also, 

specimens cooled quickly exhibited lower residual splitting-tensile strengths than that exhibited by specimens 

cooled slowly.

Figure 2.135 Effect of cooling rate on residual flexural strength.

Source: R.V. Balendran, A. Nadeem, T. Maqsood, and H.Y. Leung, ―Flexural and Split Cyliner Strengths of HSC 

at Elevated Temperature,‖ Fire Technology 39, pp. 47-61, 2003.

Figure 2.136 Effect of cooling rate on residual splitting-tensile strength.

Source: R.V. Balendran, A. Nadeem, T. Maqsood, and H.Y. Leung, ―Flexural and Split Cyliner Strengths of HSC 

at Elevated Temperature,‖ Fire Technology 39, pp. 47-61, 2003.

The effect of elevated temperature on the splitting-tensile strength of plain and fiber-reinforced concrete has been 

investigated [2.128]. The concrete utilized ordinary Portland cement, sand, 20-mm maximum size crushed gravel 

aggregate, and a water-cement ratio of 0.55 (by weight). Fiber reinforcement (0.0, 0.5, 0.7, or 1.0%, by volume) 

was included in the mixes and consisted of either hooked steel (50-mm long x 0.35-mm diam) or cement fil alkaliresistant glass fibers (50-mm long x 0.25–mm diam). Cylindrical test specimens 100-mm-diameter by 200-mm-long 

were cast from the mixes and cured for 7 days in water followed by air drying in the laboratory for an additional 

21 days before exposing to elevated temperature. The cylinders were exposed to temperatures of 100 ̊, 200 ̊, 350 ̊, 

600 ̊, 700 ̊, or 800 ̊C for 90 minutes. Following temperature exposure, six cylinders were removed from the oven 

and tested immediately, six cylinders were cooled at room temperature for two hours prior to testing, and six 

cylinders were cooled in water for two hours prior to testing. Figure 2.137 presents results for the plain concrete 

specimens for the three testing conditions. Results clearly indicate that the splitting-tensile strength of the plain 

concrete decreases with temperature with the decrease increasing as the temperature increases. Cooling in air and 

water cooling increased the loss of splitting-tensile strength with temperature. Similar results were obtained for the 

1% steel fiber concrete. Figure 2.138 presents the effect of fiber type and volume on the splitting-tensile strength of 

specimens tested hot. Results show that the splitting-tensile strength of the fiber-reinforced concrete was higher 

than that of the plain concrete for all temperatures and that the strength of the glass fiber concrete was lower than 

that of the steel fiber concrete. At temperatures greater than 350 ̊C, the effect of fiber content was not significant, 

probably as a result of fiber unbonding.
97

Figure 2.137 Effect of temperature on splitting-tensile strength of plain concrete.

Source: F.I. Faiyadh and M.A. Al-Ausi, ―Effect of Elevated Temperature on Splitting Tensile Strength of Fibre 

Concrete,‖ International Journal of Cement Composites and Lightweight Concrete 11(3), pp. 175-178, 

August 1989.

Figure 2.138 Effect of temperature and fiber reinforcement type and quantity on 

splitting-tensile strength of specimens tested at temperature.

Source: F.I. Faiyadh and M.A. Al-Ausi, ―Effect of Elevated Temperature on Splitting Tensile Strength of Fibre 

Concrete,‖ International Journal of Cement Composites and Lightweight Concrete 11(3), pp. 175-178, 

August 1989.

Table 2.5 provides additional results on the Effect of elevated temperature on residual splitting-tensile strength of 

fiber-reinforced concrete [2.129]. Cube specimens 100 x 100 x 100 mm3 were prepared using high-strength 

ordinary Type I Portland cement concrete with 19-mm maximum size basalt coarse aggregate, silica sand, and 

different volumetric mixtures of three types of fibers (i.e., brass-coated steel, hooked steel, and high performance 

polypropylene) with a total fiber volume of 2%. Reference concrete compressive strengths were 77.3, 89.8, 85.5, 
98

93.75, and 103.6 MPa for plain concrete, hooked steel fibrous concrete (HS), hooked steel- high performance 

polypropylene fibrous concrete (HSHP), hooked steel-brass coated steel fibrous concrete (HBCS), and brass coated 

steel fibrous concrete (BCS), respectively. Results in the table indicate that although the splitting-tensile strengths 

of the fiber-reinforced concrete were generally much higher than those for the plain concrete at temperatures below 

500 ̊C, the percentage loss was larger for the fibrous concretes probably due to differences in thermal expansion 

coefficient between the fibers and matrix leading to loss of bond. At higher temperatures, the presence of fibers 

reduced the extent of cracking and spalling.

Table 2.5 Residual splitting-tensile strengths of plain and fibrous concretes after temperature exposure

T ( ̊C) Plain HS HSHP HBCS BCS

23 11.8 (100%) 22.65 (100%) 20.45 (100%) 20.7 (100%) 22.5 (100%)

350 8.85 (75%) 13.4 (59.2%) 11.45 (56%) 13.35 (64.5%) 16.4 (72.9%)

500 5.9 (50%) 7.9 (34.9%) 7 (34.2%) 9.25 (44.7%) 7.8 (34.7%)

600 3.55 (30.1%) 5.18 (22.5%) 4.25 (20.8%) 4.9 (23.7%) 5.55 (24.7%)

700 Spall 1.7 (7.5%) 1.65 (8.1%) 1.7 (8.2%) 1.4 (6.2%)

Source: R.H. Haddad, R.J. Al-Saleh, and N.M. Al-Akhras, ―Effect of Elevated Temperature on Bond Between Steel 

Reinforcement and Fiber Reinforced Concrete,‖ Fire Safety Journal 43(5), pp. 334-343, July 2008.

The effect of silica fume addition on the splitting-tensile and compressive strengths has been investigated for a 

lightweight concrete [2.130]. The concrete mixes utilized ordinary Type I Portland cement with partial cement 

replacement by silica fume in amounts representing either 0 (mix M), 10 (mix H), 20 (mix A), or 30% (mix E) the 

total cementitious materials content. Crushed pumice (16-mm maximum size) was used as the coarse aggregate and 

the water-cement ratio was 0.77 (by weight). Compressive strength was determined using 100-mm cubes and 

splitting-tensile strength with cylindrical test specimens 100-mm-diameter by 200-mm-long. Specimens were cured 

in a water tank for 28 days prior to heating to temperatures of 200 ̊, 400 ̊, or 800 ̊C where they were maintained for 

one hour prior to permitting them to slowly cool to room temperature. Figure 2.139 presents the effect of elevated 

temperature on the residual compressive and splitting-tensile strengths. Highest initial compressive and splittingtensile strengths were obtained for the mix containing 20% cement replacement by silica fume. Starting at a 

temperature of about 200 ̊C, the compressive and splitting-tensile strengths started to drop for each of the mixes and 

exhibited a similar trend in strength reduction with increasing temperature.

Figure 2.139 Effect of temperature and silica fume content on 

residual compressive and splitting-tensile strengths of a lightweight concrete.

Source: H. Tanyildizi and A. Coskun, ―Performance of Lightweight Concrete with Silica Fume After High 

Temperature,‖ Construction and Building Materials 22, pp 2124-2129, 2008.
99

Examples of Relations for Concrete Tensile Strength at Elevated Temperature. Based on test results the 

following has been proposed for concrete tensile strength as a function of temperature [2.107]:

ft

T

= (1- 0.001T)ft (20 ̊ to 1000 ̊C) (2.54)

where and are the tensile strength at elevated temperature and room temperature, respectively. After exposure 

to elevated temperature the following has been proposed for estimating concrete tensile strength [2.131]: 

(2.55)

which has been simplified to a two-part expression

for 20 ̊ < T < 300 ̊C, and (2.56)

for 300 ̊ < T < 800 ̊C. (2.57)

An expression for the relationship of normalized residual tensile strength ( ) to temperature (T) exposure has 

been proposed [2.108]:

for 20 ̊ < T < 100 ̊C, (2.58)

for 100 ̊ < T < 200 ̊C, and (2.59)

for 200 ̊ < T < 800 ̊C. (2.60)

Summary. Conclusions from the limited tensile test data available in the literature are that the aggregate type and 

mixture proportions have a significant effect on the tensile strength vs temperature relationship; the decrease in 

tensile strength of calcareous aggregate concrete is twice as high as that of siliceous aggregate concrete at 500°C; 

concretes with lower cement content have lower reduction in tensile strength than those with higher cement content; 

the rate of heating has minimal effect on tensile strength at high temperature; and the residual tensile strength is 

somewhat lower than the tensile strength measured at elevated temperature.

2.2.1.7 Shrinkage and Creep

Information and Data. When freshly hardened concrete is exposed to the ambient environment and humidity it will 

generally undergo thermal shrinkage (shrinkage strain associated with cooling) and drying shrinkage (shrinkage 

strain associated with moisture loss). Generally for massive structures the drying shrinkage is less important than 

the thermal shrinkage. Length (or volume) change of concrete under these conditions is important because of its 

effect on movement of the structure and its tendency to induce cracking. 

Shrinkage occurs as a result of two effects: (1) drying or (2) autogeneous volume change. Drying shrinkage results 

from the loss of absorbed water and is generally the more predominant of the two effects. The rate and magnitude of 

drying shrinkage generally increase with temperature. Several factors affect concrete drying shrinkage: (1) cement 

and water contents (shrinkage varies directly with water-cement ratio) [2.132]; (2) composition and fineness of 

cement; (3) type and gradation of aggregate (shrinkage inversely proportional to size and amount of coarse 

aggregate - sandstone, slate, basalt and trap rock produce concretes having greater shrinkage than quartz, limestone, 

dolomite, granite, and feldspar aggregate concretes); (4) admixtures (those that reduce water requirement reduce 

shrinkage); (5) moisture and temperature conditions; and (6) amount and distribution of reinforcement. Although 

shrinkage can take place over long periods of time, the shrinkage rate decreases with time (e.g., 14 to 34% of the 

20-year shrinkage occurring in two weeks, 40 to 80% in three months, and 66 to 85% in one year). [2.133]. 

 

Long-term shrinkage may be due to carbonation.
100

Autogeneous shrinkage is more prevalent in mass concrete structures where the total moisture content remains 

relatively constant; it results from continued cement hydration reducing the free-water content (products of 

hydration occupy less volume than the sum of the separate volumes of the components) [2.134]. Autogeneous 

shrinkage (linear) varies between approximately 40 x 10-6at one month and 100 x 10-6at five years [2.122]. 

Figure 2.140 presents the effects of several variables influencing autogeneous shrinkage [2.122].

Figure 2.140 Effect of several factors on autogenous shrinkage of concrete.

Source: G. E. Troxell, H. E. Davis, and J. W. Kelly, Composition and Properties of Concrete, 2nd Ed., McGrawHill, New York, 1968.

In mass structures where the concrete is maintained below 100°C, shrinkage will not be a significant factor over the 

30- to 40-year design life of a structure such as a PCPV [2.135]. The shrinkage in the main body of mass concrete 

remote from the outside environment is caused by self-desiccation of the hardened cement paste component [2.136]. 

After the initial period where the heat of hydration results in thermal movement, even over long periods of time, the 

shrinkage is low in comparison to that of smaller members where moisture loss to the external environment can be 

large. Variation of shrinkage of unsealed concrete stored at different relative humidities for 30 years indicates that if 

the estimated external relative humidity of the mass concrete at normal temperatures does not drop below 85%, 

shrinkage after 30 years will be less than 400 microstrain as noted in Figure 2.141. High shrinkage begins when the 

main bulk of capillary-held water in concrete is lost [2.133], which is only likely to occur if the water-cement ratio 

of the concrete mix is less than 0.40 [2.137]. Other results indicate shrinkage values for sealed specimens at room 

temperature of 65 microstrain after 5 years and less than 35 microstrain for sealed specimens at 45 ̊C after 

9.5 months [2.136]. Shrinkage results obtained from eight 15.2-cm-diam by 30.5-cm-long sealed cylinders cast 

using the Wylfa concrete mix design are presented in Figure 2.142 and, on average, appear to be linear with log 

time. Initial high strains resulted from chemical reactions of the cement. Extrapolation of results over 30 years 

(Figure 2.143) indicates an expected minimum shrinkage of 100 microstrain. The maximum shrinkage limit shown 

in the figure was obtained by doubling the sealed shrinkage up to commissioning (i.e., when heating would 

commence) and then increasing the shrinkage to provide a value of 400 microstrain. This was to account for site 

mix variations and the uncertain effects of prolonged heating.
101

Figure 2.141 Variation of long-term shrinkage of concrete with relative humidity of storage.

Source: R.D. Browne, ―Properties of Concrete in Reactor Vessels,‖ Proceedings of the Conference on Prestressed 

Concrete Pressure Vessels, Group C, Paper 13, pp. 131-151, Institute of Civil Engineers, London, United 

Kingdom, 1967.

Figure 2.142 Shrinkage measurements obtained from sealed cylinders cast from Wylfa concrete mix.

Source: R.D. Browne, ―Properties of Concrete in Reactor Vessels,‖ Proceedings of the Conference on Prestressed 

Concrete Pressure Vessels, Group C, Paper 13, pp. 131-151, Institute of Civil Engineers, London, United 

Kingdom, 1967.
102

Figure 2.143 Shrinkage design limits for Wylfa PCPV.

Source: R.D. Browne, ―Properties of Concrete in Reactor Vessels,‖ Proceedings of the Conference on Prestressed 

Concrete Pressure Vessels, Group C, Paper 13, pp. 131-151, Institute of Civil Engineers, London, United 

Kingdom, 1967.

Creep can be defined as the increase in strain in a structural member with time due to sustained stress. Creep and 

drying shrinkage have a number of commonalities: (1) both originate from same source (hydrated cement paste), 

(2) stress-strain curves are very similar, (3) factors that influence one generally influence the other in the same way, 

(4) in concrete the magnitude of both (on the order of 400 to 1000 x 10-6 microstrain) is large and can not be ignored 

in design, and (5) both are partially reversible [2.91]. Because creep affects strains, deflections, and stress 

redistribution, it is important with respect to structural analysis. Creep may also be viewed from another standpoint: 

if a loaded specimen is restrained from movement (constant net strain), creep will manifest itself as a progressive 

decrease in stress with time (stress relaxation). Figure 2.144 summarizes the various combinations of loading, 

restraint, and humidity conditions associated with drying shrinkage and creep. The magnitude of creep and drying 

shrinkage is also affected by the size and shape of the concrete element [2.91].

Several theories for the creep mechanism have been proposed: viscous flow of the cement-water paste, closure of 

internal voids, crystalline flow of aggregate, and seepage into internal voids of colloidal (adsorbed) water formed by 

cement hydration [2.122]. Some investigators divide creep into two types: (1) basic and (2) drying creep [2.138]. 

Basic creep has been defined as the load-induced, time-dependent deformation of a specimen which is loaded after 

achieving thermal, hygral, chemical, and dimensional stability at first heating to a given temperature [2.139]. 

Drying creep is a function of the moisture loss from the concrete and is related to drying shrinkage. In sealed 

concrete, drying creep is absent and it is usually the practice to describe the creep during first heating simply as 

transitional thermal creep. Total creep is the sum of basic and drying creep. Under typical service conditions 

concrete is most likely drying while under load resulting in creep deformations greater than if the concrete had been 

dried prior to loading [2.90]. Figure 2.145 illustrates concrete creep under simultaneous drying and loading. A 

description of the strains that develop for loaded and unloaded concrete during first heat-up, at constant temperature, 

during cooling, and residual strains is available as well as a method for isolating individual strain components and 

performing an assessment of their magnitude [2.8]. Although creep is generally considered only for specimens 

loaded in compression, creep of concrete in tension also occurs and is on the same order of magnitude as creep in 

compression [2.140].

 

A method for predicting the long-term behavior of prestressed concrete containments is provided in Regulatory 

Guide 1.35.1, ―Determining the Prestressing Forces for Inspection of Prestressed Concrete Containments,‖ U.S. 

Nuclear Regulatory Commission, Washington, D.C., July 1990.

Also included in this reference is a method for isolating individual strain components and performing an 

assessment of their magnitude.
103

Figure 2.144 Interaction of loading, restraint, and humidity conditions.

Source: P.K. Mehta and P.J.M. Monteiro, Concrete – Microstructure, Properties, and Materials, McGraw Hill, 

New York, New York, 2006.
104

Figure 2.145 Illustration of concrete creep under simultaneous drying and loading.

Source: S. Mindess and J.F. Young, Concrete, Prentice-Hall, Inc., Englewood Cliffs, New Jersey, 1981.

Several physical and environmental parameters affect creep deformations of normal weight concrete. Physical 

parameters inherent to the particular concrete mix include: (1) cement type (degree of hydration); (2) cement paste 

proportions and content (creep proportional to volume fraction of cement paste in mix); (3) aggregate properties and 

volume fraction {(a) aggregate restrains creep, (b) mineral character effects are presented in Figure 2.146, and 

(c) creep tends to be inversely proportional to maximum aggregate size for uniformly graded mixes [2.122,2.141]}; 

(4) strength and stage of hydration [(a) creep decreases with degree of cement hydration of a mix, and (b) generally 

the amount of creep is inversely proportional to the concrete strength]; (5) moisture conditions of storing {creep is 

generally inversely proportional to the relative humidity of the medium surrounding the concrete (Figure 2.147) 

[2.133]}, and (6) size of mass (the larger the mass, the lower the creep). 

Figure 2.146 Effect of aggregate type on drying shrinkage and creep of concrete.

Source: G. E. Troxell, H.S. Davis, and J.W. Kelly, Composition and Properties of Concrete, 2nd Ed., 

McGraw-Hill, New York, 1968.
105

Figure 2.147 Creep of concrete stored at different relative humidities.

Source: A. M. Neville, Properties of Concrete, Pitman, London, United Kingdom, 1970.

Mechanical parameters affecting creep include: (1) state of stress: {(a) under uniaxial compressive stress for 

stress/strength ratio < 0.4, creep is proportional to applied stress; (b) at high stress-strength levels (>0.85), creep can 

lead to failure; (c) creep under multiaxial compression is less than under uniaxial compression of the same 

magnitude in the given direction (Figure 2.148); (d) creep occurs under hydrostatic compression [2.142-2.144]}; 

(2) age at loading [specific creep, or creep strain per unit stress, decreases for increased loading age]; and 

(3) temperature [(a) creep follows the same general pattern as creep at room temperature – being an exponential 

Figure 2.148 Typical creep-time curves under multiaxial compression: (a) biaxial, and (b) triaxial.

Source: A. M. Neville, Creep of Concrete: Plain, Reinforced and Prestressed, North-Holland Publishing Company, 

Amsterdam, The Netherlands, 1970.
106

function of time under load and a relatively linear function of stress up to a stress-strength ratio of ~0.4; (b) sealed 

specimens exhibit less creep than unsealed; (c) creep definitely increases with temperature up to at least 50°C and 

probably increases with temperature up to 150°C; and (d) the degree of creep recovery appears to be more 

dependent on stress level than temperature] [2.145,2.146]. 

Upon release of the sustained load, an initial elastic recovery of strain occurs followed by creep recovery that can 

continue for several days. The magnitude of creep recovery is greater for concrete specimens that were loaded later in 

their cure cycle, and is inversely proportional to the period of sustained stress. The reversibility of drying shrinkage 

and creep is illustrated in Figure 2.149. Both the drying shrinkage and creep of concrete exhibit a degree of 

reversibility. Figure 2.149a relates to drying shrinkage and Figure 2.149b represents a plain concrete specimen 

subjected to sustained uniaxial compression for 90 days and then unloaded resulting in instantaneous or elastic 

recovery approximately of the same order of magnitude as the elastic strain on first loading. Creep recovery occurs 

more rapidly than creep but is not complete. 

Figure 2.149 Reversibility of drying shrinkage and creep: (a) drying and rewetting, 

and (b) loading and unloading under uniaxial compression.

Source: P.K. Mehta and P.J.M. Monteiro, Concrete – Microstructure, Properties, and Materials, McGraw Hill, 

New York, New York, 2006.

Additional data and information on concrete shrinkage and creep under uniaxial sustained compressive stress is 

available in a comprehensive database on the subject that has been developed at Northwestern University in 

Evanston, Illinois [2.147]. The database contains results for 621 creep tests and 426 shrinkage tests. Potential 

applications of the database include development of prediction models, reevaluation, recalibration, and mutual 

comparisons of the existing creep and shrinkage prediction models (e.g., RILEM [2.148] and ACI Committee 209 

[2.149]). The database can be downloaded freely from http://www.iti.northwestern.edu.
107

The effects of temperature on the creep of hardened cement paste can be broadly classified as thermal and structural 

[2.150]. The thermal effect is that due to the temperature at loading, being seated in the molecular agitation caused by 

temperature. The structural effect will depend on the maximum exposure temperature, on the assumption that cooling 

down to the loading temperature does not reverse any structural changes caused by heating or cause structural 

changes of its own. The above assumption regarding the cooling will be true only if differential thermal strains within 

the specimen are minimized by a slow rate of cooling and if hydration is not allowed to take place. It was shown that 

(1) the thermal effect of temperature on creep can be modeled by an Arrhenius-type rate theory, (2) the structural 

effect of temperature on creep can be classified as that due to loss in strength and that due to a stabilization process, 

and (3) in the range 300°C to 635°C, the above two processes are opposed to each other (i.e., strength is reduced 

causing an increase in creep potential, while stabilization, reflected by percentage weight loss, is increased causing a 

decrease in creep potential) [2.150].

Like all solid materials, creep of concrete increases with temperature. An example of the effect of temperature 

(21.1 ̊, 71.1 ̊, and 96.1 ̊C) and stress level (35%, 60%, and 70%) on creep of is presented in Figure 2.150. Below 

100°C, concrete creep at moderate stress levels originates in the cement paste, probably because of the mutual 

approach of adjacent laminar particles of cement gel, which is facilitated by the presence of water in gaps between 

the particles [2.16]. In unsealed specimens the influence of drying becomes significant in the temperature range of 

70 to 100 ̊C [2.16]. At moderate temperature levels hydration (aging) accelerates, but as the temperature increases 

the reverse of this effect takes place (dehydration accelerates creep). Above 100°C drying of the concrete is very 

rapid with an associated increase in the creep rate until a stable moisture condition is reached [2.151]. After the 

moisture is lost and a stable moisture content at a given temperature is reached, the creep rate becomes less than that 

before loss of moisture [2.16]. In addition to the moisture content and drying effects, the size and shape of the 

structural member are also important in that they affect the time and rate of moisture loss [i.e., large section 

members such as walls or a base mat in a nuclear power plant require extended periods of time (days or years) for 

the moisture to diffuse out of the structure]. Results indicate that when the magnitudes of creep strains in unsealed 

concrete are compared with the magnitudes of sealed ones (mass concrete), the creep of sealed concrete is greater 

with it being 0.7 to 5 times greater than the corresponding values for unsealed concrete above a temperature of about 

71°C [2.152]. Figures 2.151 and 2.152 present examples of the effect of temperature on creep rate and creep. 

Temperature accelerates the diffusion of the solid components and water along gaps between the particles.

Figure 2.150 Relation between creep and logarithm of time under load for concretes stored at different 

temperatures and loaded to different stress levels. (Temperatures are in ̊F and data points for 205 ̊F not shown).

Source: K.W. Nasser and A.M. Neville, ―Creep of Concrete at Elevated Temperatures,‖ Journal of the American 

Concrete Institute, Title 62-87, pp. 1567-1579, Farmington Hills, Michigan, December 1965.
108

Figure 2.151 Variation of creep rate of Portland cement/porphyry aggregate concrete with temperature.

Source: J.C. Marechal, ―Creep of Concrete as a Function of Temperature,‖ Paper SP 34-30 in Concrete for Nuclear 

Reactors, pp. 547-564, American Concrete Institute, Farmington Hills, Michigan, 1972.

Figure 2.152 Creep of Portland cement/porphyry concrete at various temperatures.

Source: J.C. Marechal, ―Creep of Concrete as a Function of Temperature,‖ Paper SP 34-30 in Concrete for Nuclear 

Reactors, pp. 547-564, American Concrete Institute, Farmington Hills, Michigan, 1972.

Tests have been carried out to investigate the effect of age of concrete at loading (i.e., at 3, 8, 12, 18, 32, 90, 300, or 

365 days) and its water-cement ratio (i.e., 0.5 or 0.7) [2.153]. Eight-day compressive strengths for the two mixes 

were 11.9 and 29.1 MPa for the higher and lower water-cement ratio mixes, respectively. Prismatic concrete 

specimens 20 x 20 x 60 cm3fabricated from Type II Portland cement and granite aggregate were used in the study. 

Specimens were either water soaked at room temperature or simulated mass concrete. Test temperatures were either 

room temperature (21.5 ̊C) or 45 ̊C. Figure 2.153 presents specific creep strains (total) for specimens tested at either
109

Figure 2.153 Specific creep strains at room temperature and 45 ̊C for young (left) and mature (right) concretes.

 

Source: A.F. da Silveira and C.A. Florentino, ―Influence of Temperature on the Creep of Mass Concrete,‖ Paper SP 

25-7 in Temperature and Concrete, pp. 173-189, American Concrete Institute, Farmington Hills, Michigan, 

1970.

room temperature or 45 ̊C and either loaded early (i.e., 3 or 8 days) or at later ages (i.e., 28, 90, or 365 days). 

Specific creep strains in concrete loaded at age at 45 ̊C could be fit reasonably well by an expression developed by 

McHenry [2.154]:

, (2.61)

where t is the time after loading, and A, B, , and are characteristic parameters determined experimentally. 

Conclusions of the study were that creep strains and creep recovery strains were of the same nature and could be 

described by the same analytical expression for the temperature range of the study, temperature rise increases creep 

strains, the greater the temperature the greater the creep recovery strains, influence of temperature is greater for 

concretes having a higher water-cement ratio, creep strains are on the same order for water-soaked concrete and a 

similar mass concrete, and ratio of transverse to longitudinal strains remains approximately constant during a creep 

test and has a value equal to the Poisson’s ratio.

Most testing of creep of concrete at temperatures above 100 ̊C utilized specimens that were unsealed (i.e., freely 

lose moisture). A review of the literature on the effect of elevated temperature on the time-dependent volume 

change due to load (i.e., creep) of concrete has been conducted [2.155]. Data from several investigators [2.156-

2.162] were compiled and normalized in terms of specific creep or specific creep rate (average slope of creep curve 

during specified period in semi-log plot), Figures 2.154-2.156. Conclusions derived from these results were: 

(1) Creep at elevated temperature follows the same general pattern as creep at room temperatures (i.e., it is 

approximately an exponential function of the time under load and a fairly linear function of the stress applied 

at least up to a stress/strength ratio of about 0.50). Sealed or water-stored specimens generally exhibit less 

creep than unsealed specimens, and creep decreases with increasing maturity and increases with increasing 

moisture content of the specimen at loading. Poisson's ratio in creep appears to be unaffected by elevated 

temperatures. 
110

(2) The effect of elevated temperatures (at least up to 50 ̊ C) is to increase creep, creep at 50 ̊ C being 

approximately two to three times as great as creep at room temperature. 

(3) For temperatures of 50 ̊ to about 100 ̊ C, some controversy exists about whether or not there is a further 

increase of total creep with increasing temperature. Some investigators have found a definite maximum of 

total creep in the range of 50 ̊ to 80 ̊ C, but most have not and have concluded that creep of concrete increases 

with temperature up to around 100 ̊ C, the creep at 100 ̊ C being on the order of four to six times as great as 

the creep at room temperature (at the end of a 60- to 100-day loading period). 

(4) Conversely, some investigators also have found a definite maximum for the creep rate between 50 ̊ and 80 ̊ C 

if the creep rate is computed for some period between 1 and 107 days under load (Figure 2.156). This seems 

to indicate that as the temperature increases, a larger portion of the total creep deformation occurs during the 

first few hours under load with the effect that the creep values at the end of a l00-day loading period, for 

instance, increase steadily with temperature, in spite of a creep rate maximum at about 50 ̊ to 80 ̊ C for the 1-

to 100-day loading period.

(5) Few data are available concerning the creep at temperatures exceeding 100 ̊ C. Tests on unsealed specimens 

showed no appreciable change in total creep within the temperature range of about 100 ̊ to 140  ̊C; the creep 

rate for a 1- to l00-day loading period appeared to decline. Beyond 110 ̊ C both creep rate and creep 

magnitude increase with temperature (unsealed specimens) [2.162 and 2.163]. 

(6) Some controversy apparently exists concerning creep recovery. Although References [2.156,2.157.2.164] 

found creep recovery to be essentially independent of temperature and stress, Reference [2.165] reported a 

significantly higher creep recovery at 45 ̊ C than at room temperature. 

(7) The described experimental results concerning the effect of' temperature appears to further support the 

"seepage theory," while casting some further doubt upon the validity of other concepts. 

Figure 2.154 Influence of temperature on creep for sealed or water-stored 

specimens at up to 100 days loading.

Source: H. G. Geymayer, ―Effect of Temperature on Creep of Concrete: A Literature Review,‖ Paper SP-34-31 in 

Concrete for Nuclear Reactors, pp. 565-589, American Concrete Institute, Farmington Hills, Michigan, 

1972.

Two series of creep tests have been conducted using quartz aggregate concretes having compressive strengths of 

37.9 and 46.2 MPa [2.166]. Cylindrical test specimens 76.2-mm diameter by 152.4-mm long were loaded at an age 

of 58 days to stress-strength ratios of either 40 or 60% and heated to temperatures of 93.3 ̊, 204.5 ̊, 315.6 ̊, or 

426.7 ̊C. The specimens were first loaded to their test values for a period of one day and then heated to the desired 

temperature at 1.7 ̊C/hr. Heating continued for 326 days at which time it was terminated and the specimens were 
111

Figure 2.155 Influence of temperature on creep for unsealed specimens at up to 107 days loading.

Source: H. G. Geymayer, ―Effect of Temperature on Creep of Concrete: A Literature Review,‖ Paper SP-34-31 in 

Concrete for Nuclear Reactors, pp. 565-589, American Concrete Institute, Farmington Hills, Michigan, 

1972.

Figure 2.156 Influence of temperature on creep rate (fc/fc’ < 0.5).

Source: H. G. Geymayer, ―Effect of Temperature on Creep of Concrete: A Literature Review,‖ Paper SP-34-31 in 

Concrete for Nuclear Reactors, pp. 565-589, American Concrete Institute, Farmington Hills, Michigan, 

1972.

permitted to slowly cool to room temperature. Figures 2.157 and 2.158 present creep results for the 37.9 and 

46.2 MPa concrete, respectively, for the two stress-strength levels investigated. Conclusions of the investigation 

were that: (1) the shape of the creep-time curves of concrete at elevated temperature was the same as that at room 

temperature; (2) creep rate is higher when concrete is subjected to a high temperature and a high stress-strength 

ratio, with the latter affecting creep more; (3) concrete with lower water-cement ratio has less creep than concrete 

with higher water-cement ratio, and (4) a nonlinear relationship generally exists between creep and stress-strength 

ratio.
112

Figure 2.157 Relation between creep and time under load: 37.9 MPa concrete.

Source: C.H. Wang, ―Creep of Concrete at Elevated Temperatures,‖ Paper SP 27-19 in Designing for Effects of 

Creep, Shrinkage, Temperature in Concrete Structures,‖ American Concrete Institute, Farmington Hills, 

Michigan, 1971.

Figure 2.158 Relation between creep and time under load: 46.2 MPa concrete.

Source: C.H. Wang, ―Creep of Concrete at Elevated Temperatures,‖ Paper SP 27-19 in Designing for Effects of 

Creep, Shrinkage, Temperature in Concrete Structures, American Concrete Institute, Farmington Hills, 

Michigan, 1971.

Creep data under elevated-temperature conditions (e.g., >100 ̊C) using sealed specimens to represent mass concrete 

are limited because of the difficulty in sealing the specimens to prevent moisture loss. Several steady-state creep 

tests have been conducted in support of development of prestressed concrete reactor vessels for high-temperature 

gas-cooled reactor applications. Test temperatures were generally limited to the ASME Code limit of 65°C, 

although some data are available at higher temperatures.

Creep data have been developed for limestone or dolerite aggregate concretes loaded to 14.55 MPa at temperatures 

between 20 ̊ and 95 ̊C and ages from 7 and 400 days [2.167]. Although data initially indicated that creep versus log 

time from loading was linear, the data deviated upward from the straight line for longer times under load. A log-log 

plot of results showed that for the sealed specimens the experimental creep curves remained linear up to six years. 

Incorporating data from other investigators [2.156,2.158,2.159,2.162], led to the following expression for the creep 

curve:

= a(t)n or log = log a + n log t (2.62)

where

= specific creep strain,

a = factor decreasing with age at loading, k, and increasing with absolute temperature , 

t = time under load in days, and 

n = factor decreasing with age at loading, k, and varying with absolute temperature.
113

Relative to the slope of the straight lines in a log-log plot, it was hypothetically suggested that n can be expressed in 

terms of a modified Arrhenius activation energy equation such as:

(2.63)

where

C = a constant,

= stress,

E = activation energy, and

R = Boltzman constant.

In a program to observe the creep in mass concrete (sealed specimens) at high temperatures (up to 232.2 ̊C), 7.62-

cm-diam by 22.9-cm-long concrete cylinders were subjected to various stress-strength ratios (20%, 30%, and 60% 

of 14-d compressive strength) for periods greater than 6 months [2.145]. A Type III cement with 1.9l-cm maximum 

size aggregate (dolomite and hornblende) and a water-cement ratio of 0.60 were used in all tests. After curing the 

specimens at test temperature for 13 days, the specimens were loaded and creep strains measured for 6 months. The 

specimens were then unloaded and creep recovery monitored for about 70 days. At temperatures of 121 ̊C and 

above the axial stresses were corrected to account for the steam pressure. Figure 2.159 presents an example of creep 

versus log time results obtained at a test temperature of 176.7 ̊C and stress levels of 8.27, 12.41, and 17.24 MPa. 

The results indicate that two straight lines, one covering 1 to 21 days and one covering 21 to 180 days, can be used 

to fit the data at each stress level. Figure 2.160 presents an example of the affect of temperature on creep versus 

time results for specimens loaded to a stress-strength ratio of 20%. Although the shape of the curves was found to 

be similar, the creep at 148.9 ̊C was found to be the largest from the 40th day onward (i.e., after 180 days, the creep 

at temperatures of 176.7 ̊C and 232.2 ̊C was less than creep at 148.9 ̊C). Creep recovery was found to be smaller 

than the corresponding creep, independent of temperature, dependent on stress, and having a maximum recovery 

strain of 390 x 10-6. Additional data indicating the effect of load level and temperature on creep of a 42 MPa quartz 

aggregate concrete loaded to stress-strength levels ( ) of either 0.2 or 0.4, and a 39.8 MPa quartz aggregate concrete 

loaded to a stress-strength level of 0.3 are presented in Figures 2.161 and 2.162, respectively.

Figure 2.159 Effect of stress level on creep versus log time in days at 176.7 ̊C.

Source: K.W. Nasser and R.P. Lohtia, ―Creep of Mass Concrete at High Temperatures,‖ Journal of American 

Concrete Institute 68(4), pp. 276-281, Farmington Hills, Michigan, April 1971.
114

Figure 2.160 Effect of temperature on creep versus time in days for a stress-strength ratio of 20%.

Source: K.W. Nasser and R.P. Lohtia, ―Creep of Mass Concrete at High Temperatures,‖ Journal of American 

Concrete Institut, 68(4), pp. 276-281, Farmington Hills, Michigan, April 1971.

Figure 2.161 Influence of load level and temperature on creep of a quartz aggregate concrete.

Source: H. Gross, ―On High Temperature Creep of Concrete,‖ Paper H6/5 in Proceedings of 2nd International 

Conference on Structural Mechanics in Reactor Technology, Elsevier Science Publishers, North-Holland, 

The Netherlands, 1973.
115

Figure 2.162 High temperature creep of ordinary concrete with quartz aggregate.

Source: U. Schneider, Behaviour of Concrete at High Temperature, HEFT 337, Deutscher Ausschuss für 

Stahlbeton, Wilhelm Ernst & Sohn, Munich, Germany, 1982.

A 5-year investigation on creep and other properties of concrete for prestressed concrete reactor vessels (PCRVs) 

has been conducted [2.168]. The creep tests were made on sealed concrete specimens to simulate mass concrete. 

The test parameters included: temperatures of 23 ̊, 43 ̊, and 71°C; nominal stress levels of 30%, 45%, and 60% of 

the reference compressive strength; and ages at loading of 28, 90, and 270 d. Additional tests were conducted to 

determine autogeneous length change and drying shrinkage of the concrete, the effect of testing temperature (23 ̊, 

43 ̊, and 71°C) on compressive strength, and the influence of up to five thermal cycles (23° C to 71° C to 23° C) on 

compressive strength, splitting-tensile strength, modulus of elasticity, and Poisson’s ratio. All creep specimens were 

cast in the vertical position with sealing of the specimen against loss of moisture accomplished by means of steel 

end plates and a wrapping of 1.59-mm-thick butyl rubber around the specimen. All specimens were stored at 23° C 

prior to testing. In Figure 2.163 the total strains obtained for Mix No. 1 are plotted for the 13 test conditions 

investigated. Making a direct comparison of strain data obtained at different levels of stress requires that values of 

total strain per unit of applied constant stress be computed. The effect of temperature on total strain of concrete per 

unit of applied stress was for the strain to increase with an increase in temperature. Concrete tested at higher stress 

levels achieved higher total strains and thus exhibited the highest strains per unit of applied constant stress. Age at 

loading had a more significant effect on strains of concretes tested at 23°C than for concretes tested at 43°C or 71°C. 

For the seven groups of specimens for which creep recovery at 23°C or 71°C was obtained, about 39% of the 90-d 

creep recovery occurred during the first day and 65% within 10 d after unloading, independent of the previous 

applied stress level. At 71°C creep recovery of only one group of specimens was observed. This group experienced 

31% and 52% of the 90-d creep recovery at 1 and 10 d, respectively, after unloading. In general, the drying 

shrinkage strains leveled off between 400 and 600 microstrains at all test conditions, with the higher drying 
116

shrinkage strains occurring at the higher temperatures where humidities were lowest. Modulus of elasticity of the 

creep specimens was determined during loading, subsequent unloading, and when testing the creep specimens to 

failure on completion of the creep phase of the program. The modulus of elasticity of the concretes ranged from 40.7 

to 44.8 GPa, with an average Poisson’s ratio of 0.22. 

Figure 2.163 Total strains for a number of test parameters.

Source: J. Komendant. V. Nicolayeff, M. Polivka, and D. Pirtz, ―Effects of Temperature, Stress Level, and 

Age at Loading on Creep of Sealed Concrete,‖ Paper SP 55-3 in Douglas McHenry International 

Symposium on Concrete and Concrete Structures, American Concrete Institute, Farmington Hills, 

Michigan, 1978.

Creep of concrete under multiaxial loading conditions has been investigated [2.169,2.170]. The objective of the 

study was to develop information on the time-dependent deformation behavior of concrete in the presence of 

temperature, moisture, and loading conditions similar to those encountered in a prestressed concrete pressure vessel. 

Four primary factors were addressed in the study: (1) temperature during curing history (23.9 ̊ and 65.6 ̊C), 

(2) curing history (as air-dried or as-cast), (3) several multiaxial load combinations with stress levels varying from 0 

to 24.8 MPa, and (4) time covering a period of 18 months after casting. Cylindrical concrete specimens 15.2-cmdiameter by 40.6-cm-long were fabricated from a 41±4.1 MPa concrete mix. One day after casting the as-cast 

specimens were sealed by coating them with epoxy. One day later a second epoxy coating was applied and the 

specimens placed into copper jackets that were sealed by soldering. At an age of 83 days, a neoprene jacket was 

applied to the specimens. Air-dried specimens at an age of 48 hours were submerged and cured in a lime-saturated 

water solution for five days after which they were stored in a laboratory environment at 22.8 ̊C and 60% relative 

humidity until an age of 81 days they were sealed using the same procedure as for the as-cast specimens. Strains 

were measured using vibrating wire gages. Loading was initiated at an age of 90 days. Compressive loads were 

applied along the three principal axes of the cylindrical specimens with loads along the longitudinal axis varied 

independently of the loads applied along the radial axes permitting triaxial, biaxial, and uniaxial states of stress. 

Figure 2.164 presents a schematic of the multiaxial testing system. Time-creep strain curves illustrating the 

influence of temperature during loading and time on axial and radial creep strains are presented in Figures 2.165 and 
117

Figure 2.164 Schematic of multiaxial testing system.

Source: G.P. York, T.W. Kennedy, and E.S. Perry, Experimental Investigation of Creep in Concrete Subjected to 

Multiaxial Compressive Stresses and Elevated Temperature, Research Report 2864-2, Department of Civil 

Engineering, The University of Texas at Austin, June 1970.

2.166, respectively. Results indicate that creep strains for curing histories and temperatures utilized in the study 

were generally unaffected by time at later ages after loading indicating there was no significant interactions between 

time and temperature during loading or between time and curing history. Each of the factors addressed affected 

creep. Compressive and tensile creep strains were generally larger for: (1) a test temperature of 65.6 ̊ as compared 

to 23.9 ̊C, (2) an air-dried concrete than an as-cast concrete (except for low-tensile creep where the opposite was 

true), (3) increased time after loading, and (4) higher stresses for uniaxial and biaxial states of stress. The effects of 

major interactions indicated that increasing the stress level increased the creep strains but that the increase was 

larger for specimens at the higher temperature relative to the lower temperature and for air-dried specimens relative 

to as-cast specimens, except for specimens exhibiting small tensile creep strains. Also, this increase was larger after 

longer periods under load. 
118

Figure 2.165 Effect of time after loading and temperature on axial creep strain for a variety of stress conditions.

Source: G.P. York, T.W. Kennedy, and E.S. Perry, Experimental Investigation of Creep in Concrete Subjected to 

Multiaxial Compressive Stresses and Elevated Temperature, Research Report 2864-2, Department of Civil 

Engineering, The University of Texas at Austin, June 1970.

Figure 2.166 Effect of time after loading and temperature on radial creep strain for a variety of stress conditions.

Source: G.P. York, T.W. Kennedy, and E.S. Perry, Experimental Investigation of Creep in Concrete Subjected to 

Multiaxial Compressive Stresses and Elevated Temperature, Research Report 2864-2, Department of Civil 

Engineering, The University of Texas at Austin, June 1970.
119

As the confining pressure increased, the overall creep strains decreased. The effect of temperature and time after 

loading on creep Poisson’s ratio are presented in Figure 2.167 for as-cast and air-dried specimens at comparable 

stress conditions. Within the test conditions of the study, creep Poisson’s ratio values ranged from approximately 

39% to 84% of the elastic Poisson’s ratio values with an overall average value throughout the testing program of 

0.16. Creep Poisson’s ratio values for air-dried concrete were approximately 30% less than those from as-cast 

specimens. The magnitude of stress and state of stress influenced creep Poisson’s ratio although this influence was 

less at higher temperature and for air-dried concrete. Creep recovery results indicated from 1 to 46% of the creep 

strain (average = 17%) at time of unloading was recovered 84 days after release of load with the percentage of creep 

recovery depending primarily on the curing history (e.g., as-cast specimens recovered more than air-dried 

specimens). Percentage of creep recovery was not significantly affected by temperature during loading, elastic 

strain recovery, magnitude of stress, or type of stress applied.

Figure 2.167 Effect of temperature and time after loading on creep Poisson’s ratio.

Source: G.P. York, T.W. Kennedy, and E.S. Perry, Experimental Investigation of Creep in Concrete Subjected to 

Multiaxial Compressive Stresses and Elevated Temperature, Research Report 2864-2, Department of Civil 

Engineering, The University of Texas at Austin, June 1970.
120

In a related study [2.171], the time-dependent deformation of concrete was investigated by subjecting it to various 

stress conditions and elevated temperatures using a multiaxial testing system similar to that presented in 

Figure 2.164. Cylindrical test specimens were 15.2-cm diameter by 40.6-cm long. Factors addressed in the study 

included modulus of elasticity of aggregate [Tennessee limestone (main), Alabama greywacke (low), and Alabama 

chert (high)], curing history prior to loading (as-cast and air-dried), temperature during loading (22.7 ̊C and 65.6 ̊C), 

and state of stress during loading (uniaxial, biaxial, hydrostatic, and triaxial stress state) with axial and radial states 

of stress ranging from 0 to 16.54 MPa. Examples of axial and radial creep strain versus time results for triaxiallyloaded specimens ( A = 16.54 MPa, R = 4.14 MPa) are presented in Figures 2.168 to 2.171. Results are presented 

for cure history (as-cast and air-dried specimens), aggregate modulus (high, main, and low), and test temperature 

(22.7 ̊C and 65.6 ̊C). The effect of time after loading on creep Poisson’s ratio is presented in Figure 2.172. In the 

investigation it was found that compressive and tensile total creep strains were generally larger for (1) a test 

temperature of 65.6°C than for 23.9°C, (2) an air-dried concrete than an as-cast concrete (except for low tensile 

creep), (3) increased time after loading, and (4) uniaxial than biaxial states of stress, and that the axial creep strains 

decreased with increasing confining stress (Figure 2.173). Compressive and tensile creep strains were generally 

proportional to the inverse of the elastic modulus of both the concrete and aggregate. Creep Poisson’s ratio was 

generally smaller than elastic Poisson’s ratio. The percentage of creep recovered was essentially the same for 

variations in curing, temperature, concrete modulus, and stress level investigated, with multiaxial stress states 

averaging only about 5% more creep recovery than for uniaxial stress states.

Figure 2.168 Axial and radial creep strain vs time: as-cast specimens at 22.7 ̊C.

Source: J.E. McDonald, Time-Dependent Deformation of Concrete Under Multiaxial Stress Conditions, Technical 

Report C-75-4, U.S. Army Engineer Waterways Experiment Station, Vicksburg, Mississippi, October 

1975.
121

Figure 2.169 Axial and radial creep strain vs time: air-dried specimens at 22.7 ̊C.

Source: J.E. McDonald, Time-Dependent Deformation of Concrete Under Multiaxial Stress Conditions, Technical 

Report C-75-4, U.S. Army Engineer Waterways Experiment Station, Vicksburg, Mississippi, October 

1975.

Figure 2.170 Axial and radial creep strain vs time: as-cast specimens at 65.6 ̊C.

Source: J.E. McDonald, Time-Dependent Deformation of Concrete Under Multiaxial Stress Conditions, Technical 

Report C-75-4, U.S. Army Engineer Waterways Experiment Station, Vicksburg, Mississippi, October 

1975.
122

Figure 2.171 Axial and radial creep strain vs time: air-dried specimens at 65.6 ̊C.

Source: J.E. McDonald, Time-Dependent Deformation of Concrete Under Multiaxial Stress Conditions, Technical 

Report C-75-4, U.S. Army Engineer Waterways Experiment Station, Vicksburg, Mississippi, October 

1975.

Figure 2.172 Effect of time after loading on creep Poisson’s ratio.

Source: J.E. McDonald, Time-Dependent Deformation of Concrete Under Multiaxial Stress Conditions, Technical 

Report C-75-4, U.S. Army Engineer Waterways Experiment Station, Vicksburg, Mississippi, October 

1975.
123

Figure 2.173 Relative effect of type of loading on compressive creep.

Source: J.E. McDonald, Time-Dependent Deformation of Concrete Under Multiaxial Stress Conditions, Technical 

Report C-75-4, U.S. Army Engineer Waterways Experiment Station, Vicksburg, Mississippi, October 

1975.

Summary. Investigations of creep behavior at elevated temperature produced the common observation that 

increasing temperature results in substantially higher creep strains. In general, the specific creep of sealed 

specimens is shown to be less than that for specimens subjected to some degree of drying. Also, it is apparent that 

creep will be substantially greater for young concrete in both sealed and unsealed conditions. The phenomenon of a 

"creep maximum" has been observed by several investigators. This term is somewhat of a misnomer, because the 

observation is that the specific creep rate reaches a maximum with increasing temperature and is not necessarily 

accompanied by a corresponding maximum in actual creep strain. The observed maximum has been reported to 

occur at 50 to 100°C and in one case for a 20% stress-strength ratio at up to 150°C. In fact, not all studies have 

reported the maximum creep rate effect. Also, in general, most studies have reported that the shape of the curves for 

creep versus time at high temperatures is similar to those at room temperature. Creep recovery has been observed to 

be less than the associated creep strain. The degree of creep recovery appears to be independent of temperature but 

dependent on stress. In addition, shrinkage strains of concrete are reported to be very low for sealed specimens, and, 

in fact, high temperature exposure has been shown to result in expansion. With regard to the stress-strength ratio, 

increasing ratios increase creep substantially. Considering the reports on deterioration of compressive strength at 

temperatures over 100°C for sealed specimens, the stress at which creep becomes structurally significant is 

substantially decreased. For sustained temperatures above 100°C, the variations in experimental techniques, 

concrete mixtures, curing, and loading histories prohibit the development of a reliable general conclusion on longterm behavior. 

2.2.1.8 Concrete–Steel Reinforcement Bond Strength

Information and Data. In addition to the effect of elevated temperature on the properties of the steel and concrete, 

knowledge of the effects of temperature on the bond strength between concrete and steel is important for 

understanding the response and structural capacity, including residual, of reinforced concrete to thermal loadings. 

Bond arises primarily from friction and adhesion between concrete and steel and may be affected by the relative 

magnitude of concrete shrinkage. It is a function of (1) the concrete properties (cement type, admixtures, watercement ratio), (2) the mechanical properties of the steel (size and spacing of lugs), and (3) rebar position within the 

concrete member (bond is greater for vertical bars than for horizontal bars). Permissible bond stresses are generally 

specified as percentages of concrete’s compressive strength. Although considerable research has been conducted 

investigating the bond between concrete and steel at room temperature, results indicating the effect of elevatedtemperature exposure are somewhat limited. The general trend for the effect of elevated temperature on the bond 

strength of concrete to mild steel is shown in Figure 2.174. Concrete aggregates used in the tests were sandstone, 

andesite, serpentine, and basalt [2.50].

 

Appendix B provides information on the effects of elevated temperature on steel reinforcement.
124

Figure 2.174 Indication of effect of elevated temperature on relative bond strengths 

of concretes fabricated using different types of conventional aggregate materials.

Source:

Relative residual bond stress was evaluated by testing specimens fabricated from concrete mixes using river gravel 

and ordinary Portland cement containing embedded plain, round steel bars [2.172]. The specimens were fabricated 

from four different mixes, moist cured for 28 days at 20 ̊C, and then stored at 21 ̊C and 85% relative humidity until 

tested at an age of 90 days. The specimens were subjected to temperatures from 20 ̊C to 300 ̊C for 90 days, 

permitted to slowly cool to room temperature, and then tested. It was found that the residual bond stress (evaluated 

at free-end slip of 0.025 mm) after subjecting the specimens to 300°C for 90 d and then cooling to room temperature 

was only about 50% the reference value before heating (Figure 2.175). 

Figure 2.175 Relative residual bond strengths of heated concretes.

Source: H. Kasami, T. Okuno, and S. Yamane, ―Properties of Concrete Exposed to Sustained Elevated 

Temperature,‖ Paper Hl/5 in Proc. 3rd International Conference on Structural Mechanics in Reactor 

Technology, Elsevier Science Publishers, North-Holland, The Netherlands, 1975.
125

In a similar investigation, normal weight concrete specimens 140 x 140 x 300 mm3and fabricated from ordinary 

Portland cement and river gravel coarse aggregate were tested to investigate the effect of reinforcement type (i.e., 

20-mm-diameter plain round bars or ribbed bars) [2.173]. The specimens were heated to temperatures up to 450 ̊C 

and then permitted to cool to room temperature prior to testing. Slip at both the loaded and nonloaded ends was 

measured as a function of load. Figure 2.176 presents bond stress versus temperature results for the ribbed and plain 

round bars and shows that the ribbed bars experienced a loss of bond strength only above 400°C, but the smooth 

bars lost strength after only a small temperature increase. 

Figure 2.176 Relative residual bond strength of ribbed and plain round bars.

Source: A. F. Milovanov and G. D. Salmanov, ―The Influence of High Temperature Upon the Properties of 

Reinforcing Steels and Upon Bond Strength Between Reinforcement and Concrete,‖ Issledovanija po 

zharoupornym betonu I zhelezobetonu, pp. 203–223, 1954.

A comparison of the effects of elevated temperature on the bond of 14-mm-diameter ribbed and plain round bars 

embedded in 15 x 15 x 45 cm3prisms for two different concrete strengths (17 and 33 MPa) is presented in 

Figure 2.177 where results are presented relative to those obtained from unheated specimens [2.174]. The concrete 

mixes consisted of ordinary Portland cement in combination with an aggregate mixture of 60% river gravel and 40% 

crushed granite. Residual bond strengths after exposure to 500 ̊C and 700 ̊C and then cooling for 24 hours at room 

temperature ranged from about 80% to 50% the unheated values for the ribbed steel bars and about 45% to 25% for 

the plain round bars. 

Figure 2.177 Relative residual bond strength of ribbed and plain round bars 

for different concrete compressive strengths.

Source: H.V. Reichel, ―How Fire Affects Steel-to-Concrete Bond,‖ Building Research and Practice 6(3), pp. 176–

186, May/June 1978.
126

Figure 2.178 presents the effect of bar diameter on bond strength after elevated temperature exposure for ribbed and 

plain round bars [2.175]. The results were obtained from a series of tests involving 280 specimens with varied bar 

diameters from 8 to 25 mm that were heated to temperatures up to 800 ̊C, maintained at temperature for 2 hours, and 

then permitted to slowly cool to room temperature prior to testing. The concrete was fabricated from Danish sea 

gravel (mixture of quartz, granite, and limestone) and had a compressive strength of 20 MPa. The bond strength as a 

percentage of the concrete compressive strength was higher for the ribbed bars than the plain round bars, and at 

temperatures up to 400 ̊C was about 65% the concrete compressive strength. 

Figure 2.178 Effect of bar diameter on bond strength after 

elevated-temperature exposure for ribbed and plain round bars.

Source: K. Hertz, ―The Anchorage Capacity of Reinforcing Bars at Normal and High Temperatures,‖ Magazine of 

Concrete Research 34(121), pp. 213–220, December 1982.
127

In another investigation the bond strength at high temperature for ribbed steel bars, plain round bars, and deformed 

prestressing bars was evaluated by testing cylindrical concrete pull-out specimens 191-mm-high by 172-mmdiameter [2.176]. The 16-mm maximum size siliceous aggregate concrete mix had cube compressive strengths of 

48.0 to 60.9 N/mm2at 28-days and 50.7 to 63.9 N/mm2at testing (150 to 600 days age). The specimens were heated 

to the desired test temperature, held at temperature for 3 hours, and then tested. Slip at both the loaded and 

nonloaded ends was measured as a function of applied load. Bond stress-slip results as a function of temperature are 

presented in Figure 2.179 for 16-mm diameter cold deformed steel and 7.5-mm diameter prestressing steel 

reinforcement. Figure 2.180 presents bond stress-slip results for heavily rusted plain 8-mm diameter round bars of 

cold-deformed steel, and relative bond strength-temperature results for plain round mild steel bars, cold deformed 

steel, and shaped prestressing steel. Results of the investigation indicate that loss of bond strength for ribbed bars at 

constant elevated temperature is of the same order of magnitude as loss of high-temperature compressive strength 

and at same temperatures round plain bars exhibit a sharper decrease in bond strength than ribbed bars. Results for 

the rusted plain mild steel round bars were not significantly affected by elevated temperatures up to 200 ̊C, but at 

temperatures of 300 ̊C and above the bond stress at a given slip decreased significantly. 

Figure 2.179 Bond-slip relationship at elevated temperature for

cold deformed steel and prestressing steel.

Source: U. Diederichs and U. Schneider, ―Bond Strength at High Temperature,‖ Magazine of Concrete Research 

33(115), pp. 75–83, June 1981.
128

Figure 2.180 Relative bond strength as a function of temperature.

Source: U. Diederichs and U. Schneider, ―Bond Strength at High Temperature,‖ Magazine of Concrete Research

33(115), pp. 75–83, June 1981.

Concrete specimens 15 x 15 x 20 cm3 were used to evaluate rebar bond strength [2.177]. The concrete mix used 

ordinary Type I Portland cement and had a compressive strength of 49.2 MPa after water curing for 28 days. Two 

No. 3 reinforcing bars were embedded vertically in each specimen from both ends with embedded lengths of 6 and 

8 cm, respectively. After water curing the specimens were air dried for a week and then subjected to elevated 

temperatures up to 550 ̊C. After attaining the desired soak temperature, the specimens were maintained at this 

temperature for different periods ranging from 30 minutes to 180 minutes, after which they were permitted to slowly 

cool to room temperature. Testing was initiated 24 hours after reaching room temperature. Figure 2.181 presents 

relative residual bond strength results as function of exposure time for several different soak temperatures. A 

substantial decrease in bond strength was observed at temperatures greater than 200°C, and the bond strength 

decreased with increasing exposure time. A procedure for experimentally and numerically relating the ratio of 

residual bond strength to temperature and exposure time, and a procedure for integrating changes in bond strength 

due to temperature fluctuations to predict final residual bond strength was proposed.
129

Figure 2.181 The effect of elevated-temperature exposure time 

on the relative residual bond strength of No. 3 bars embedded in a concrete cube.

Source: C-H. Chiang and C-L. Tsai, ―Time-Temperature Analysis of Bond Strength of a Rebar After Fire 

Exposure,‖ Cement and Concrete Research 33, pp. 1651–1654, 2003.

Figure 2.182 presents the influence of type aggregate on the relative bond strength [2.178]. The type of aggregate is 

one of the main factors determining the high temperature bond strength. The lower the thermal strain of concrete 

the higher the bond strength at elevated temperature.

Figure 2.182 Bond between concrete and deformed bars exposed to high temperature.

Source: H. Sager and F.S. Rostasy, ―High Temperature Behavior of Reinforcing and Prestressing Steels,‖ 

Sonderforschungs-Bereich 148, Part II, pp. 51–53, Technical Universitat Braunschweig, Germany, 1980.

Figure 2.183 presents the affect of curing conditions (e.g., in water unsealed, sealed, in air unsealed) prior to 

elevated-temperature exposure on the relative variation in bond strength at start of pull-out for a hard sandstone 

aggregate concrete after various heating periods at 175°C [2.97]. Specimens 150 x 150 x 150 mm3containing a D25 

rebar embedded 10 cm were used for the tests. The bond strength of specimens unsealed during heating exhibited 

up to about a 30% decline for heating exposures of 3 d or less, but specimens that were sealed exhibited practically 
130

no reduction as a result of temperature exposure. Prolonged elevated-temperature exposure for sealed specimens 

exhibited a positive effect on the bond strength, and for the unsealed specimens a recovery in bond strength to the 

point that after 91-d exposure it was reduced by 10 to 15% relative to control results. Other results also indicate that 

the loss in bond strength between concrete and steel reinforcement at temperatures < 65°C is small (≤15%) 

[2.176,2.179].

Figure 2.183 Relative variation in bond strength at start of pull-out

for a hard sandstone aggregate concrete after various heating periods at 175°C.

Source: K. Hirano, K. Ohmatsuzawa, T. Takeda, S. Nakane, T. Kawaguchi, and K. Nagao, ―Physical Properties of 

Concrete Subjected to High Temperature for MONJU,‖ Paper P2-25, Power Reactor and Nuclear Fuel 

Development Corporation, Tokyo, Japan.

The effect of elevated temperature on the residual bond strength of a siliceous aggregate ordinary Portland cement 

concrete was evaluated by testing 10-cm-diameter by 20-cm-long cylindrical bond pull-out specimens [2.54]. The 

concrete strength just prior to initiation of heating (age of about 26 months) was 46.7 MPa. The specimens were 

heated to soak temperatures of 100 ̊, 300 ̊, or 450 ̊C, maintained at temperature for 72 hours, and then slowly cooled 

to room temperature prior to testing. Figure 2.184 presents residual bond strength results relative to those for 

unheated specimens, as well as residual results for weight, compressive strength, dynamic modulus of 

Figure 2.184 Effect of elevated temperature on residual strength ratios.

Source: T. Harada, T. Takeda, S, Yamane, and F, Furumura, ―Strength, Elasticity and Thermal Properties of 

Concrete Subjected to Elevated Temperatures,‖ Paper SP-34-21in Concrete for Nuclear Reactors, pp. 377-

406, American Concrete Institute, Farmington Hills, Michigan, 1972.
131

elasticity, and static modulus of elasticity. The relative residual bond strength at time of free-end slippage of 0.05 

mm was approximately 44% at 100 ̊ to 300 ̊C, and 10% at 450 ̊C. The relative residual maximum bond strength 

was approximately 50 to 60% at 100 ̊ to 300 ̊C, and 10% at 450 ̊C. It was concluded that the percentage reduction 

in bond strength due to elevated temperature exposure was greater than that for the compressive strength.

The response of bond in reinforced concrete to high temperature was investigated for four test conditions: 

(1) stressed during heating (3.7 MPa) and loaded to failure while hot, (2) stressed during heating (3.7 MPa) and 

loaded to failure after cooling, (3) no applied stress during heating and loaded to failure while hot, and (4) no stress 

during heating and loaded to failure after cooling [2.180]. The effect of cover depth was also investigated. 

Cylindrical specimens 300 mm in length by various diameters according to cover depth were fabricated from a 

19-mm maximum size natural gravel aggregate, ordinary Portland cement, and had a cube strength of 35.0 MPa 

after curing in air for 3 months. Bonded within the central portion of each specimen over a 32-mm depth was a 16-

mm diameter Tor bar (high yield strength deformed bar). A standard cover depth of 55 mm was used except cover 

depths of 25, 32, and 46 mm were also investigated for test condition 2. Specimens were heated to temperatures up 

to 750 ̊C, held at temperature for one hour, and then tested or permitted to cool to room temperature where they 

remained for 24 hours prior to testing. Figure 2.185 presents bond stress-slip curves for various temperatures for 

specimens tested using test condition 2. A comparison of hot and residual maximum bond stress and temperature 

Figure 2.185 Residual bond stress vs slip relationships at various temperatures: 

16-mm diameter bars with 55 mm cover and stressed at 3.7 MPa while heated.

Source: P.D. Morley and R. Royles, ―Response of the Bond in Reinforced Concrete to High Temperatures,‖ 

Magazine of Concrete Research 35(123), pp. 67-74, June 1983.

results is presented in Figure 2.186. These results tend to indicate that specimens heated with a steady-state bond 

stress applied provided a slightly improved performance relative to results obtained from specimens heated under a 

no-stress condition due to restraint applied by loading. The effect of cover on the ratio of residual maximum bond 

stress to maximum bond stress at 20 ̊C with a steady-state stress of 3.7 MPa held for 24 hours while heated is 

presented in Figure 2.187. The maximum bond stress and the maximum bond slip both decreased as the cover depth 

decreased. The large difference in slip between smaller and larger cover depths was caused by the difference in 

mode of failure experienced. Specimens with larger cover depths exhibited a greater resistance to splitting 

permitting greater slip to take place prior to concrete failure. General conclusions of the study were that the 

reduction in bond strength with temperature was greater than that for the concrete compressive strength, crushing of 

concrete immediately beneath the rebar ribs was the cause of a critical value in the bond stress-slip curve, and the 

bond stress was dependent on concrete strength.
132

Figure 2.186 Stressed and unstressed hot and residual bond stress results at various temperatures.

Source: P.D. Morley and R. Royles, ―Response of the Bond in Reinforced Concrete to High Temperatures,‖ 

Magazine of Concrete Research 35(123), pp. 67-74, June 1983.

Figure 2.187 Variation of maximum bond stress with temperature for different cover depths.

Source: P.D. Morley and R. Royles, ―Response of the Bond in Reinforced Concrete to High Temperatures,‖ 

Magazine of Concrete Research 35(123), pp. 67-74, June 1983.

The effect of load cycling on the ratio of maximum residual bond stress to maximum bond stress at 20 ̊C was 

investigated [2.181]. Cylindrical test specimens 126-mm diameter by 300-mm long containing a 15-mm diameter 

rebar embedded over a length of 32 mm within the central portion of the cylinder were used. Before loading to 

failure, the effect of cycling the bond stress twenty times between 1.0 and 3.7 MPa was examined. Figure 2.188 
133

presents the effect of load cycling on residual bond strength for deformed rebar. Increasing the temperature caused 

the bond stress-slip relationship to decrease and produced irreversible slip, particularly in the temperature range 250 ̊ 

to 450 ̊C indicating a significant decrease in bond performance. The dashed line in the figure to 720 ̊C indicates that 

these specimens failed before the sequence of twenty load cycles and loading to failure had been completed.

Figure 2.188 Effect of load cycling on the residual bond strength for 16-mm-diameter deformed rebar.

Source: R. Royles and P.D. Morley, ―Further Responses to the Bond in Reinforced Concrete to High 

Temperatures,‖ Magazine of Concrete Research 35(124), pp. 157-163, September 1983.

The residual performance of bond between high-strength pozzolanic concrete and reinforcing steel has been 

investigated [2.182]. Cylindrical pullout specimens either 82- or 100-mm diameter by 150-mm long were cast with 

an 18-mm diameter Grade 60 deformed steel rebar embedded along the entire specimen length. Concrete mixes 

used a 19-mm maximum size limestone coarse aggregate and ordinary Portland cement or a combination of ordinary 

Portland cement and natural pozzolan. Four concrete mixes were used in which the ordinary Portland cement was 

replaced with natural pozzolan in weight percentages of either 0, 10, 15, or 25. Concrete compressive strengths for 

these mixes were 73.1, 66.0, 67.0, and 66.1 MPa, respectively. After curing in water at 23 ̊C for 40 days, the 

specimens were heated to temperatures of either 600 ̊ or 800 ̊C where they remained for 1 hour prior to permitting 

them to slowly cool to room temperature over about a 26 hour period. The bond behavior of the pullout specimens 

was evaluated using a fixture that applied a constant displacement rate of 0.01 mm/s with slippage monitored by two 

LVDTs. Compressive strength results obtained from companion 100-mm-cube test specimens exhibited almost 

equal strength reduction percentages (i.e., 39.2 to 40.8%) at 600 ̊C, except for the specimens in which 25% of the 

ordinary Portland cement was replaced by natural pozzolans (48.2% reduction). At 800 ̊C, however, the percentage 

reduction in compressive strength increased as the percentage of pozzolan increased. Post-test examination of 

specimen surfaces indicated that cracking and the maximum crack widths at 800 ̊C were greater than that observed 

at 600 ̊C. Figure 2.189 presents residual bond stress versus free-end slip results obtained after exposure to either 

600 ̊ or 800 ̊C for concretes having different natural pozzolan contents. The curves corresponding to unheated 

specimens exhibited almost linear response at low slip ranges before becoming nonlinear near failure. Bond 

strength was significantly reduced after exposure to temperatures of 600 ̊ and 800 ̊C, with the reduction higher for 

the 800 ̊C value (e.g., up to 74% versus up to 24%). Curves for the specimens that had been heated exhibited 

nonlinear behavior over the entire slip range with a varying rate of stress change. This was attributed to cracks that 

formed at low to medium pullout loads that due to heating tended to close permitting greater slippage. Once the 

cracks closed resistance to slippage increased as did the pullout load required for further slippage. Near failure local 

crushing of the concrete occurred near the rebar ribs resulting in significant slippage followed by a sudden splitting 

failure of the concrete along the rebars, Relative to natural pozzolan content, at 600 ̊C the percentage reduction in 

bond strength increased as the natural pozzolan content increased.
134

Figure 2.189 Effect of natural pozzolan content on residual bond stress versus slip results.

Source: R.H. Haddad and L.G. Shannis, ―Post-Fire Behavior of Bond Between High Strength Pozzolanic Concrete 

and Reinforcing Steel,‖ Construction and Building Materials 18, pp. 425-435, 2004.

The bond behavior between fiber-reinforced concrete and 20-mm diameter reinforcing steel bars under elevated 

temperature exposure has been investigated [2.129]. Modified pullout specimens 100 x 100 x 400 mm3 were 

prepared using high-strength ordinary Type I Portland cement concrete with 19-mm maximum size basalt coarse 

aggregate, silica sand, and different volumetric mixtures of three types of fibers (i.e., brass-coated steel, hooked 

steel, and high performance polypropylene) with a total fiber volume of 2%. Reference concrete compressive 

strengths were 77.3, 89.8, 85.5, 93.75, and 103.6 MPa for plain concrete, hooked steel fibrous concrete (HS), 

hooked steel-high performance polypropylene fibrous concrete (HSHP), hooked steel-brass coated steel fibrous 

concrete (HBCS), and brass coated steel fibrous concrete (BCS), respectively. Pullout specimens contained Grade 

60 deformed rebars having a diameter of 20 mm. Specimens were cured for 28 days at 40 ̊C and then placed into 

laboratory air at 60% relative humidity and 23 ̊C prior to heat treatment. Heat-treated specimens were then 

subjected to temperatures ranging from 350 ̊ to 700 ̊C for 2 hours and then permitted to slowly cool to room 

temperature prior to testing. The presence of fibers in the concrete mix reduced the length and intensity of cracks in 

the temperature range of 350 ̊ to 500 ̊C, but had limited effect at higher temperatures because of loss of bond to the 

concrete matrix. The contribution of fibers to maintaining residual compressive and tensile strengths was limited in 

the temperature range of 350 ̊ to 600 ̊C but was significant at 700 ̊C. Figure 2.190 presents bond stress versus freeend slip results for plain concrete, hooked steel fiber concrete, hooked steel-high performance polypropylene fibrous 

concrete, hooked steel-brass coated steel fibrous concrete, and brass coated steel fibrous concrete. Results exhibit

linear behavior up to about 40% of ultimate bond stress, and from then on a nonlinear response. Exposure of the 
135

modified bond pullout specimens fabricated using different combinations of fibers to temperatures in the range of

500 ̊ to 700 ̊C resulted in softening of the bond stress versus free-slip curves. Response was determined by the state 

of the fibrous concrete materials with negligible effect provided by steel rebars. At elevated temperatures cracks 

increased to reduce the confinement around the rebars resulting in a reduction in bond stress and an increase in freeend slip relative to lower temperatures (e.g., 350 ̊C). Figure 2.191 summarizes relative residual bond strength versus 

temperature results obtained for the plain and fibrous concrete mixes. As noted in the figure, the residual bond 

strength decreases with increasing exposure temperature with the largest decrease occurring at 500 ̊C. The decrease 

in bond strength resulted from an increase in intensity, width, and extension of cracks with temperature resulting in 

reduced confinement of steel rebar. In the temperature range of 350 ̊ to 600 ̊C the fibrous concrete mixes exhibited 

greater relative residual bond strength relative to the plain concrete mix.

Figure 2.190 Effect of fiber type on residual bond stress versus slip results.

Source: R.H. Haddad, R.J. Al-Saleh, and N.M. Al-Akhras, ―Effect of Elevated Temperature on Bond Between Steel 

Reinforcement and Fiber Reinforced Concrete,‖ Fire Safety Journal 43(5), pp. 334-343, July 2008.
136

Figure 2.191 Effect of fiber type on residual bond stress versus temperature results.

Source: R.H. Haddad, R.J. Al-Saleh, and N.M. Al-Akhras, ―Effect of Elevated Temperature on Bond Between Steel 

Reinforcement and Fiber Reinforced Concrete,‖ Fire Safety Journal 43(5), pp. 334-343, July 2008.

Examples of Constitutive Relationships for Bond Stress Between Concrete and Steel. The bond strength 

between concrete and steel decreases with increasing temperature. The magnitude of the loss is a function of the 

type of concrete and the reinforcement sttl surface condition (i.e., smooth, deformed, or degree of rusting) [2.183]. 

An estimate of the ultimate bond stress as a function of temperature can be obtained from [2.184]:

(2.64)

where and u represent the ultimate bond strength at elevated and room temperature, respectively.

An analytical model has been developed for bond stress versus free-end slip [2.129]. It is proposed that the 

ascending part of the bond stress versus free-end slip curve can be described by:

, (2.65)

where 1 is the ultimate bond strength, S1 the corresponding slip, and n is a curve-fitting parameter that must be less 

than 1. A nonlinear empirical model has been developed that relates the percentage residual bond to exposure 

temperature for the data presented in Figure 2.191 [2.129]:

, (2.66)

where RBS is the residual bond strength, T the exposure temperature for T > 23 ̊C, and k a constant depending on 

the type of concrete mixture.

Summary. Conclusions from results presented in the literature specific to concrete-reinforcing steel bond at elevated 

temperature are that the test procedure and shape of specimen have a significant influence on test results, ribbed bars 
137

exhibit improved performance relative to plain round bars, surface roughness increases the performance for plain 

round bars, the bond strength decreases as the exposure temperature increases, at high temperatures (≥200°C) the 

time at temperature affects the bond strength, the diameter of ribbed steel reinforcement (8 to 25 mm) does not have 

a significant effect on bond strength, residual bond strengths of specimens sealed during temperature exposure 

perform better than unsealed specimens (≤175°C), curing conditions are important at moderate elevated-temperature 

exposures (≤400°C), residual bond strength is lower than bond strength obtained at temperature, a clear influence of 

the water-cement ratio and concrete strength on bond strength at elevated-temperature exposure has not been 

observed, the type of aggregate has a significant effect on the high-temperature bond strength, at temperatures 

< 65°C the bond strength is relatively unaffected, the smaller the aggregate size in relation to the bar diameter the 

more workable will be the concrete and the greater the chances of obtaining good adhesion between the cement 

paste and steel, the maximum bond stress and the maximum bond slip decrease as the cover depth decreases (tensile 

strength of the concrete surrounding a reinforcing bar is enhanced as the concrete cover increases because a larger 

amount of concrete can be compacted more readily and the large difference in slip between smaller and larger cover 

depths is caused by the difference in mode of failure experienced), load cycling causes a decrease in the maximum 

bond stress, reduction in bond strength due to elevated temperature exposure appears to be greater for concretes 

having ordinary Portland cement contents replaced by natural pozzolans in amounts greater than 10% the weight, 

incorporation of fibers into the concrete mix has been shown to provide improved bond stress versus free-end slip 

results by limiting concrete cracking and reducing spalling, and the reduction in bond strength with temperature is 

greater than that for the concrete compressive strength.

2.2.1.9 Fracture Energy and Fracture Toughness 

Information and Data. The effects of elevated temperature related to exposure period and curing age on the residual 

fracture properties of a normal- and a high-strength concrete were investigated using three-point bending tests of 

preheated beams [2.185]. The 500 x 100 x 100 mm3beams were notched with a diamond saw to a depth of 50 mm 

prior to heating. Rapid-hardening Portland cement in conjunction with a 20-mm maximum size graded gravel and 

sand were used to fabricate the test specimens. Water-cement ratios (by weight) used for the normal- and highstrength concretes were 0.54 and 0.30, and 29-day compressive strengths were 57.4 and 77.6 MPa, respectively. 

Specimens were cured in water for seven days and then placed into the laboratory environment until testing. Most 

beams were exposed to temperatures from between 100 ̊ and 600 ̊C for 12 hours at 14 days, while some of the 

normal-strength beams were heated either for various exposure periods up to 168 hours at 14 days age or for 

12 hours at ages of 7, 28, and 90 days. Residual fracture energy (Gf), defined as total energy dissipated over a unit 

crack area, as a function of exposure temperature is presented in Figure 2.192 for both concretes (heating for 

12 hours at an age of 14 days). Results show that the residual fracture energy increases up to about 300 ̊C and then 

decreases with increasing temperature. The initial increase was attributed to additional cement hydration and the

Figure 2.192 Effect of temperature on fracture energy (Gf).

Source: B. Zhang, N. Bicanic, C.J. Pearce, and G. Balabanic, ―Residual Fracture Properties of Normal- and HighStrength Concrete Subject to Elevated Temperatures,‖ Magazine of Concrete Research 52(2), pp. 123-136, 

April 2000.
138

decrease in fracture energy at temperatures above 300 ̊C to increased microcracking and dehydration of cement. 

Figure 2.193 presents the effect of exposure period (th) starting at a concrete age of 14 days. When the heating 

temperature was below 300 ̊C, Gf monotonically increased with exposure period as a result of additional cement

Figure 2.193 Effect of exposure period (th) on fracture energy (Gf) of normal strength concrete.

Source: B. Zhang, N. Bicanic, C.J. Pearce, and G. Balabanic, ―Residual Fracture Properties of Normal- and HighStrength Concrete Subject to Elevated Temperatures,‖ Magazine of Concrete Research 52(2), pp. 123-136, 

April 2000.

hydration. At the higher temperatures Gfincreased initially, but as the exposure period increased (t > ~8 hours) it 

decreased as shown in the figure. The effect of different curing times on the fracture energy is shown in 

Figure 2.194 for specimens that were unheated and specimens that had been heated at 300 ̊C for 12 hours at the 

designated curing time. As shown in the figure, Gf continued to increase as the curing age increased due to 

continued cement hydration, but appeared to level off at 90 days.

Figure 2.194 Effect of curing period on fracture energy (Gf).

Source: B. Zhang, N. Bicanic, C.J. Pearce, and G. Balabanic, ―Residual Fracture Properties of Normal- and HighStrength Concrete Subject to Elevated Temperatures,‖ Magazine of Concrete Research 52(2), pp. 123-136, 

April 2000.

Figure 2.195 presents temperature dependence of residual fracture energy for an ordinary concrete and a highstrength concrete [2.186]. Both concretes were fabricated using 20-mm maximum size calcareous aggregate. Silica 

fume (5.7% by weight of total cementitious materials) was incorporated into the high-strength concrete mix. 

Compressive strengths for the ordinary and high-strength concretes at 28 days were 30 and 75 MPa, respectively. 

Tests were conducted using three-point bend specimens 100 x 100 x 400 mm3containing a 50-mm-deep notch. 

Prior to testing, specimens were heated to temperatures of 120 ̊, 250 ̊, or 400 ̊, held at temperature for three hours, 
139

and then permitted to slowly cool to room temperature. Results for the ordinary and high-strength concretes 

exhibited similar trends in that generally the fracture energy increased with increasing exposure temperature. 

Fracture energies for both concretes at 400 ̊C were about two and a half times the reference unheated values.

Figure 2.195 Effect of exposure temperature on residual fracture energy of ordinary and high-strength concretes.

Source: A. Menou, G. Mounajed, H. Boussa, A. Pineaud, and H. Carre, ―Residual Fracture Energy of Cement Paste, 

Mortar and Concrete Subject to High Temperature,‖ Theoretical and Applied Fracture Mechanics 45, 

pp. 64-71, 2006.

Beams 150 x 150 x 750 mm3 were tested under three-point bending to investigate the effect of elevated temperature 

on concrete fracture toughness [2.187]. Half the beams had a 25-mm initial notch at midspan and half the beams 

had a 60-mm notch at midspan. The notch thickness was 2.5 mm and was cast into the beams by placing a piece of 

Plexiglas with the desired crack dimension in the beam mold prior to casting the concrete. The 36 MPa compressive 

strength concrete was fabricated using Type I Portland cement, gravel coarse aggregate having a maximum 

aggregate size of 12.7 mm, and a water content corresponding to a water-cement ratio of 0.45 (by weight). Prior to 

testing, the beams were placed into a curing room at 21 ̊C and 95% relative humidity for 28 days. After curing the 

beams were removed from the curing room and heated to temperatures of 50 ̊, 100 ̊, 150 ̊, 200 ̊, 250 ̊, or 300 ̊C 

where they remained for 24 hours prior to slowly cooling to room temperature and testing. Figure 2.196 presents the 

variation of residual fracture toughness with temperature for the beams. Also shown in the figure are analytical 

results obtained from an expression presented in the reference. Results indicate a steady decline in residual fracture

Figure 2.196 Effect of temperature on residual fracture toughness of a concrete.

Source: H. Abdel-Fattah and S.A. Hamoush, ―Variation of the Fracture Toughness of Concrete with Temperature,‖ 

Construction and Building Materials 11(2), pp. 105-108. 1997.
140

toughness with an increase in exposure temperature with the decrease increasing for exposure temperatures above 

50 ̊C. The effect of number of thermal cycles, one to three from room temperature to exposure temperature to room 

temperature, on the concrete residual fracture toughness is shown in Figure 2.197. Results indicate that each 

thermal cycle leads to a reduction in the residual fracture toughness with the reduction being about 20% for each 

thermal cycle at temperatures above 50 ̊C. The effect of thermal cycling becomes more significant at temperatures 

greater than 100 ̊C.

Figure 2.197 Effect of thermal cycling on residual fracture toughness of a concrete.

Source: H. Abdel-Fattah and S.A. Hamoush, ―Variation of the Fracture Toughness of Concrete with Temperature,‖ 

Construction and Building Materials 11(2), pp. 105-108. 1997.

Summary. The residual fracture energy of both normal- and high-strength concretes increased as the exposure 

temperature increased for temperatures up to about 300 ̊ or 400 ̊C due to increased cement hydration; however, as 

the temperature further increased the residual fracture energy decreased due to increased microcracking and cement 

dehydration. The residual fracture energy of normal-strength concrete tended to increase with thermal exposure 

period for temperatures below 300 ̊C; however, at higher exposure temperatures the residual fracture energy 

increased at first for a few hours (th < 10 hours) and then decreased. The residual fracture energy increased with 

increased curing age in the first 28 days due to additional cement hydration, but after that it tended to be stable. The 

residual fracture toughness of concrete decreased as the exposure temperature increased with the decrease being 

relatively significant at temperatures greater than 50 ̊C. As the number of heating cycles increased the residual 

fracture toughness decreased, with the decrease being about 20% for each additional cycle (up to three total thermal 

cycles). 

2.2.1.10 Long-Term Exposure (Aging) 

Information and Data. During the nominal 40-year design life of nuclear power plants, certain concrete 

components may be subjected to moderately elevated temperatures that could affect the concrete’s properties. In 

conducting safety evaluations for these components, the significance of concrete component aging needs be taken 

into account. Unfortunately, only limited data have been identified related to the effects of long-term elevatedtemperature exposure on concrete properties. Surprisingly, although Portland cement concretes have been in 

existence for more than 160 years, relatively little documented information is available on the aging of concrete 

structures that are being acted upon by environmental stressors under well-defined conditions [2.188]. When 

concrete is fabricated with close attention to the factors related to the production of good concrete (e.g., material 

selection, production control, desirable properties, and costs), the concrete will have infinite durability unless 

subjected to extreme external influences (e.g., overload, elevated temperature, industrial liquids, and gases) [2.122].

An investigation has been conducted to determine the changes in mechanical properties of a limestone aggregate 

concrete after exposure to temperatures of 75°C and 300°C for periods up to 8 months and 600°C for 1 month 

[2.123]. After 8-month’s exposure to 75°C, compressive and splitting-tensile strengths were 98 and 94%, 

 

Service lives of 60 years and beyond are currently being addressed.
141

respectively, of their reference values. However, after exposure to 600°C for just 1 month, compressive and 

splitting-tensile strengths were only 23 and 38%, respectively, of their reference values. In companion mixes where 

either fly ash or blast furnace slag was used, improvement in retention of mechanical properties occurred after 

exposure to sustained high temperatures as a result of partial replacement of the cement. 

The effect of long-term exposure (up to 13 years) at moderate elevated temperature (65°C) on the mechanical 

properties of a limestone aggregate concrete was investigated [2.189]. Tests conducted during this study were 

somewhat unusual because the specimens were first subjected to a simulated temperature-vs-time cement hydration 

cycle. Also, because the concrete mix was being evaluated for an application that experienced exposure to sulfatebearing groundwater at elevated temperatures (~65°C), both ordinary and sulfate-resistant Portland cements were 

investigated. Specimens, after being subjected to the simulated cement hydration cycle, were stored either in water 

at 19°C (control specimens) or in a sodium sulfate solution (2000 ppm) at 65°C. Frequently during the test program, 

the sodium sulfate solution was changed, which required cooling to room temperature; the specimens were therefore 

also subjected to thermal cycling. Results of the study indicated that there was no evidence of long-term 

degradation in compressive strength for any of the concrete mixes and heat treatments utilized, and that for a given 

compressive strength the dynamic modulus of elasticity was lower for the concrete that had been heated. Cooling 

down and reheating the limestone and flint aggregate mixes for a total of 87 cycles did not appear to cause 

degradation in strength. 

A 5-year testing program was conducted to determine the effects of long-term exposure to elevated temperature on 

the mechanical properties of concrete used in construction of the radioactive underground storage tanks at Hanford 

Engineering Development Laboratory (HEDL) [2.190]. Tests were conducted using specimens fabricated from the 

same mix proportions and materials specified for the concrete used to fabricate the tanks (20.7- and 31.0-MPa 

design compressive strengths). Concrete strength, modulus of elasticity, and Poisson’s ratio values were determined 

from specimens subjected to 121 ̊, 177 ̊, or 232°C for periods of up to 33 months. The effect of thermal cycling was 

also investigated. Results showed that the compressive strength in general tended to decrease with increasing 

temperature and also with length of exposure; however, with the exception of the cylinders exposed to 232°C, all 

compressive strength results obtained after a 900-d exposure period exceeded the design values noted previously. 

Splitting-tensile strength results also decreased somewhat with increasing temperature and length of exposure. 

Modulus of elasticity was affected the most by the elevated-temperature exposure; after 920 d of heating at 232°C, it 

had a value that was only 30% the value obtained from an unheated control specimen. Poisson’s ratio, although 

exhibiting somewhat erratic values, was relatively unaffected by either the magnitude or the length of elevatedtemperature exposure. Thermal cycling (~18 cycles) to 177°C produced moderate reductions in compressive 

strength (5 to 20%), significant reductions in modulus (30 to 50%), and slight reductions in Poisson’s ratio (0 to 

20%). Time-dependent (creep) and physical property data were also obtained using specimens cast from the 

concrete mixes.

Associated with the laboratory investigation described in the previous paragraph was a study to confirm the 

laboratory results by testing samples removed from the underground storage tanks and process buildings at HEDL 

[2.191]. Cores 76-mm in diameter were obtained over the length of the haunch wall and footing of a single-shell 

tank that was built in 1953; contained waste for about 8 years; reached temperatures in the range of 127°C to 138°C; 

and experienced a radiation field of 0.10 to 0.13°C/kg/h (400 to 500 R/h). Although considerable scatter was 

obtained from the data because of different concrete pours and different environmental exposures, after about 

29 years of exposure only one data point fell below the 20.7-MPa design compressive strength. Figure 2.198 

presents compressive strength results obtained from cores removed from structures at HEDL and compares the 

results to values based on the laboratory results.

A study has been carried out to examine the effect of temperature on sealed and unsealed air-entrained concrete 

containing fly ash, conventional water reducer, and superplasticizer [2.192]. The properties of compressive strength 

and modulus of elasticity were studied at seven different temperatures ranging from –11°C to 232°C and at seven 

different exposure periods from 1 to 180 d. Local crushed aggregates of 19-mm maximum size consisting primarily 

of dolomite and hornblende were used in the concrete mixtures. Figure 2.199 presents the relative strength (ratio of 

compressive strength at temperature to that obtained from unheated control specimens) as a function of temperature 

and exposure time that was obtained from testing sealed specimens. The results indicate that up to a temperature of 

121°C there was no degradation in compressive strength for exposures up to 180 d. With increasing temperature, 
142

Figure 2.198 Comparison of laboratory and actual sample long-term compressive strength data.

Source: M.P. Gillen et al., ―Strength and Elastic Properties of Concrete Exposed to Long-Term Moderate 

Temperatures with High Radiation Fields,‖ RHO-RE-SA-55 P, Rockwell Hanford Operations, Richland, 

Washington, 1984.

Figure 2.199 Relationship of relative strength and temperature of mass concrete (sealed). 

Source: K.W. Nasser and M. Chakraborty, ―Temperature Effects on Strength and Elasticity of Concrete Containing 

Admixtures,‖ Proceedings of Symposium Temperature Effects on Concrete, ASTM Special Technical 

Publication 858, American Society for Testing and Materials, West Conshohocken, Pennsylvania, 1985.

the strength decreased with the extent of strength reduction generally proportional to the exposure temperature and 

time at temperature (e.g., at 232°C and 180-d exposure the strength was about 50% its reference value). The 

corresponding relationship of the relative modulus of elasticity (ratio of modulus of elasticity at temperature to that 

obtained from unheated control specimens) as a function of temperature and exposure time obtained from sealed 

specimens is presented in Figure 2.200. The modulus of elasticity started to decline monotonically at temperatures 

≥ 71°C with the decline in modulus proportional to the exposure temperature and time at temperature (e.g., at 232°C 

and 180-d exposure the modulus was about 25% its reference value). Relative compressive strength and modulus of 
143

elasticity results for unsealed specimens are presented in Figure 2.201 and indicate improved performance relative to 

the sealed specimens. An explanation of the greater effect of elevated-temperature exposure on sealed (mass 

concrete) specimens was that in a closed system saturated steam pressure develops at high temperatures, that causes 

deterioration in structural properties of the cement gel.

Figure 2.200 Relationship of elasticity ratio and temperature of mass concrete (sealed).

Source: K.W. Nasser and M. Chakraborty, ―Temperature Effects on Strength and Elasticity of Concrete Containing 

Admixtures,‖ Proceedings of Symposium Temperature Effects on Concrete, ASTM Special Technical 

Publication 858, American Society for Testing and Materials, West Conshohocken, Pennsylvania, 1985.

Figure 2.201 Relationship of strength and elasticity and temperature of unsealed concrete.

Source: K.W. Nasser and M. Chakraborty, ―Temperature Effects on Strength and Elasticity of Concrete Containing 

Admixtures,‖ Proceedings of Symposium Temperature Effects on Concrete, ASTM Special Technical 

Publication 858, American Society for Testing and Materials, West Conshohocken, Pennsylvania, 1985.
144

A laboratory study has been conducted to evaluate the relative performance of limestone and dolostone aggregate 

ordinary Portland cement concretes under sustained exposure to high temperature [2.193]. After 28-d moist cure 

followed by 26 weeks of room temperature curing, the test specimens were exposed for up to 4 months to 

temperatures ranging from 76°C to 450°C, and 1 month for a 600°C exposure. Figure 2.202 presents residual 

compressive strength versus conditioning temperature (up to 450°C) after 4 months exposure for the limestone 

concrete.* The loss of compressive strength of specimens exposed to elevated temperature was proportional to the 

exposure temperature. At temperatures of 150°C and higher, an increase in length of exposure from 48 h to 

4 months resulted in further decreases in strength. In all cases, any major loss in strength was found to occur within 

the first month of exposure. In general the leaner concretes (water-cement ratio = 0.6) were slightly less affected 

than the richer concretes in terms of relative strength loss after exposure.

Figure 2.202 Relative residual compressive strength of limestone concrete after

4-month exposure to various temperatures (up to 450°C).

Source: G. G. Carrette and V. M. Malhotra, ―Performance of Dolostone and Limestone Concretes at Sustained High 

Temperatures,‖ Proceedings of Symposium Temperature Effects on Concrete, ASTM Special Technical 

Publication 858, pp. 38–67, American Society for Testing and Materials, West Conshohocken, 

Pennsylvania, 1985.

The effect of elevated-temperature exposure at 65°C, 90°C, or 110°C for periods up to 3.5 years was investigated in 

Japan in support of nuclear power plant facilities [2.194]. Either basalt or sandstone coarse aggregates were utilized 

in the concrete mixtures. Cementitious materials studied included Class B fly ash, moderate heat cement plus fly 

ash, and normal Portland cement. Heating conditions adopted were: (1) long-term heating tests [allowable 

temperature except for local areas (long-term) (65°C), allowable local temperature (long-term) (90°C), temperature 

 

*The dolostone aggregate results are not discussed because pyrite was contained in some of the aggregate particles, 

and it underwent slow oxidation that produced a disintegrating expansion of the aggregate and cracking of the 

concrete.
145

at which water is considered to evaporate rapidly (110°C)]; (2) short-term heating tests [allowable temperature 

(short-term) (175°C)]; and (3) thermal cycling tests for up to 120 cycles [cycled heating temperatures (20°C to 

110°C to 20°C) to simulate temperature variations during operational periods]. Three cylindrical specimens were 

prepared for each test condition and put under either sealed conditions, where evaporation of water was prevented, 

or unsealed conditions, where evaporation was allowed. Figure 2.203 presents the residual compressive strength and 

modulus of elasticity results, and Figure 2.204 the effect of thermal cycling for the sealed and unsealed concretes. 

Figure 2.203 Long-term (3.5-year) heating effect on compressive strength and modulus.

Source: T. Suzuki, M. Tabuchi, and K. Nagao, ―Study on the Degradation of Concrete Characteristics in the High 

Temperature Environment,‖ Concrete Under Severe Conditions: Environment and Loading, Vol. 2, pp. 

1119–1128, E & FN Spon Publishers, 1995.
146

Figure 2.204 Effect of thermal cycling on compressive strength and modulus.

Source: T. Suzuki, M. Tabuchi, and K.Nagao, ―Study on the Degradation of Concrete Characteristics in the High 

Temperature Environment,‖ Concrete Under Severe Conditions: Environment and Loading, Vol. 2, 

pp. 1119–1128, E & FN Spon Publishers, 1995.

Several conclusions were derived from the heating tests results: 

1. Under long-term heating at 65°C, 90°C, and 110°C, compressive strength after heating was greater than before 

heating, under both sealed and unsealed conditions. This was especially true under sealed conditions at 110°C 

in which the compressive strength increased for 1.5 years reaching approximately 200% that of unheated 

control specimens. These results are considered to be the effect of autoclave curing. Under unsealed conditions 
147

it was considered that microcracking in concrete, resulting from moisture migration and evaporation under high 

temperatures, caused a degradation of compressive strength. However, the acceleration of hydration at high 

temperatures in any nonhydrated sections of the concrete increased the compressive strength more than 

degradation caused by microcracking. The results after 3.5 years were relatively unchanged from those 

measured after 1 year. Therefore, it was found that the compressive strength ceased to fluctuate at an early 

stage. 

2. Under sealed conditions during heating, elastic modulus had a tendency to increase slightly, though not as much 

as the compressive strength. The elastic modulus under sealed conditions remained relatively unchanged under 

heating, even after 3.5 years. Under unsealed conditions during heating, the modulus of elasticity at 

temperature relative to that obtained from unheated control specimens was reduced about 50%. The reduction 

in the elastic modulus under unsealed conditions can be explained by the closing of microcracks at an early 

stage of stress. Thus, the elastic modulus of concrete heated to high temperatures, with moisture migration and 

evaporation present, was markedly reduced. In addition, it was found that the elastic modulus became 

stabilized at an early stage, not changing much from 91 d through 3.5 years, even under heating. 

3. During the thermal-cycle heating test, compressive strength after heating was greater than before heating under 

both sealed and unsealed conditions. However, the ratio of increase was smaller than under constant heating, 

suggesting the influence of thermal cycling. For the same number of thermal cycles, the compressive strength 

was consistently higher for the sealed specimens relative to that obtained from unsealed specimens. Under 

unsealed conditions specimens exhibited little influence of number of cycles on compressive strength for 

thermal cycle numbers greater than five (i.e., little change in compressive strength value for cycles greater than 

five). Under unsealed conditions the modulus of elasticity exhibited a similar trend to that obtained for constant 

heating in that it was reduced by about 50%. A major part of the reduction occurred in the early stages of the 

thermal cycling. Under sealed conditions during thermal cycling the elastic modulus showed a tendency similar 

to that of the compressive strength, but the modulus of elasticity did not increase as much as the compressive 

strength of the sealed specimens. 

4. The greater the weight reduction became, the greater the decrease in the elastic modulus tended to be, which 

indicated that moisture migration and evaporation during heating affected the reduction of elastic modulus. 

Therefore, to estimate the properties of massive concrete structures subjected to high temperatures accurately, it 

is necessary to study moisture migration in mass concrete members that are subjected to high temperatures over 

long periods of time.

A series of studies was conducted to evaluate the drying effect of elevated-temperature exposure on the properties of 

concrete [2.172]. Specimens made from four concrete mixes of ordinary Portland cement and river-gravel 

aggregate (Table 2.6) were tested to investigate the compressive, tensile, and bond strengths ; moduli of elasticity; 

and weight loss after 90-d exposure to temperatures of 35 ̊, 50 ̊, 65 ̊, 80 ̊, 110 ̊, 200 ̊, and 300°C. Moisture in the 

specimens was allowed to evaporate freely. Residual strength tests of unheated and heated concretes were 

conducted at room temperature on both dried and wet specimens presoaked in water for 2 d. Exposure to sustained 

elevated temperature higher than 35°C showed remarkable deteriorating effects on the physical properties of 

concrete when moisture in concrete was allowed to evaporate. Greater reduction of strengths and weights after 

exposure were associated with the mixes having higher mix water contents. Strengths did not decline linearly as the

Table 2.6 Mix proportions for concrete mixtures

Concrete Mixture

Mix Proportions (kg/m3)

Water Cement Sand River Gravel

Mix 1 165 330 836 1045

Mix 2 195 390 779 974

Mix 3 164 273 895 1035

Mix 4 190 317 849 1036

Source: H. Kasami, T. Okuno, and S. Yamane, ―Properties of Concrete Exposed to Sustained Elevated 

Temperature,‖ Paper H1/5 in Proc. 3rd International Conference on Structural Mechanics in Reactor 

Technology, Elsevier Science Publishers, North-Holland, The Netherlands, 1975.

 

Bond strength results were discussed in Section 2.2.1.8.
148

temperature rose and were minimal at around 50°C. Dry compressive strengths of heated concretes indicated 

approximately 10% loss when heated at 50°C with practically no change when heated at higher temperatures up to 

110°C, 20% loss at 200°C, and 30% loss at 300°C. Wet compressive strengths of heated concretes indicated larger 

losses than dry strengths, and more than 20% loss when heated at only 35°C. Reductions in tensile and bond 

strengths and modulus of elasticity of heated concretes were greater than that obtained for compressive strengths. 

Tensile strength tests indicated approximately 30% loss upon exposure to temperatures of 50°C to 65°C, and less 

strength loss at higher temperatures. Bond strengths showed approximately 15% loss when heated at only 35°C, and 

smaller loss for higher temperatures of 50°C to 80°C, 25% loss at 110°C, and more than 50% loss at 300°C. Moduli 

of elasticity of heated concretes indicated a tendency to decline linearly with increasing temperature, experiencing a 

15% loss at 50°C, 25% loss at 110 ̊C, and 50% loss at 300 ̊C. The unusual deterioration at around 50 ̊C in 

compressive, tensile, flexural and shear strengths was also indicated in subsequent investigations regardless of the 

kind of aggregates. However, the influence of aggregate on the properties of heated concrete was significant. 

Sandstone and basalt aggregate concretes indicated smaller reductions, while limestone, andesite, and serpentine 

aggregate concretes showed greater reductions in strengths after exposure. Changes in chemical composition in the 

cement paste were not noticeable below 100°C; however, the porosity was found to be affected by the exposure 

temperature. The unusual deterioration in strengths at around 50°C can be due to either the expansion of cement

paste or to the change in porosity caused by evaporation of free water. Figure 2.205 presents the effect of exposure 

temperature on residual compressive strength, tensile strength, and modulus of elasticity results after 90-day 

Figure 2.205 Effect of exposure temperature on residual compressive strength, 

tensile strength, and modulus of elasticity after 90-d exposure.

Source: H. Kasami, T. Okuno, and S. Yamane, ―Properties of Concrete Exposed to Sustained Elevated 

Temperature,‖ Paper Hl/5 in Proc. 3rd International Conference on Structural Mechanics in Reactor 

Technology, Elsevier Science Publishers, North-Holland, The Netherlands, 1975.
149

exposure for each of the concrete mixtures investigated. Figures 2.206 presents weight loss results for the different 

concrete mixtures. Residual compressive strength and moduli of elasticity results showing the effect of aggregate 

material type are presented in Figure 2.207. The effect of aggregate type and temperature exposure level on shear 

strength* is presented in Figure 2.208.

Figure 2.206 Weight loss of heated concrete.

Source: H. Kasami, T. Okuno, and S. Yamane, ―Properties of Concrete Exposed to Sustained Elevated 

Temperature,‖ Paper Hl/5 in Proc. 3rd International Conference on Structural Mechanics in Reactor 

Technology, Elsevier Science Publishers, North-Holland, The Netherlands, 1975.

Figure 2.207 Compressive strength and modulus of elasticity of heated concretes.

Source: H. Kasami, T. Okuno, and S. Yamane, ―Properties of Concrete Exposed to Sustained Elevated 

Temperature,‖ Paper H1/5 in Proc. 3rd International Conference on Structural Mechanics in Reactor 

Technology, Elsevier Science Publishers, North-Holland, The Netherlands, 1975.

 

*Shear is the action of two equal and opposite parallel forces applied in planes a short distance apart. Shear stresses cannot exist 

without accompanying tensile and compressive stresses (pure shear is applied only through torsion).
150

Figure 2.208 Shear strength of heated concrete. 

Source: H. Kasami, T. Okuno, and S. Yamane, ―Properties of Concrete Exposed to Sustained Elevated 

Temperature,‖ Paper H1/5 in Proc. 3rd International Conference on Structural Mechanics in Reactor 

Technology, Elsevier Science Publishers, North-Holland, The Netherlands, 1975.

An experimental study has been conducted to rationalize the design method for facilities used to store spent fuel and 

to obtain fundamental data for estimating the long-term safety of the facilities under elevated temperature exposure 

conditions [2.195]. Four concrete mixes, listed in Table 2.7, were employed in the study: two ordinary Portland 

cement concretes having design compressive strengths of 24 and 40 MPa, and two concrete mixes having design 

compressive strength of 40 MPa in which 55% (by weight) of the ordinary Portland cement was replaced by blast 

furnace slag. Shale and limestone coarse aggregate were used in the concrete mixes. Cylindrical test specimens 

10-cm diameter by 20-cm long were cured by soaking for 28 days in water and then air drying at 20±3 ̊C and 60% 

Table 2.7 Concrete mixture proportions

Concrete

Mixture

Mix Proportions (kg/m3) Reference Values*

Water Cement Blast 

Furnace

Slag

Coarse

Aggregate

Fine

Aggregate

Water

Reducer Compressive

Strength 

(MPa)

Modulus of

Elasticity

(GPa)

No. 1 160 370 0 1103 789 0.925 56.5 35.3

No. 2 162 250 0 1082 908 0.625 32.3 28.6

No. 3 160 167 203 1103 789 0.416 58.6 33.3

No. 4 160 167 203 1103 789 0.416 54.9 34.3

*Room temperature (unheated, unsealed) reference values for mixes used in 110 ̊C test series.

Source: T. Kanazu, T. Matsumura, and T. Nishiuchi, ―Changes in the Mechanical Properties of Concrete Subjected 

to Long-Term Exposure to High Temperatures,‖ Report No. U95037, Abiko Research Laboratory, Central 

Research Institute of Electric Power Industry, Japan, March 1996.

 

Cylindrical test specimens 20-cm diameter by 40-cm long were used for splitting-tensile strength tests.
151

relative humidity for about two months. Both sealed and unsealed specimens were used in the study. Sealed 

specimens were sealed by placing into metal containers and soldering the seams. At an age of about 3 months,

heating of the specimens initiated. Exposure temperatures investigated were 65 ̊, 85 ̊, and 110 ̊C. After thermal 

exposures of 1 month, 1 year, 1.5 years, 2 years, 3 years, and 8 years specimens were removed from the ovens and 

permitted to cool to room temperature prior to testing. Compressive strength, modulus of elasticity, weight change, 

and splitting-tensile strength data were obtained and compared to reference 3-month baseline results. Figures 2.209 

and 2.210 present residual compressive strength results for unsealed and sealed specimens, respectively. Residual 

moduli of elasticity results for unsealed and sealed specimens are presented in Figures 2.211 and 2.212, respectively. 

Change in specimen weight with exposure time is presented in Figures 2.213 and 2.214 for unsealed and sealed 

specimens, respectively. Figure 2.215 presents splitting-tensile strength results. Conclusions derived from the study 

were that: the decrease in compressive strength at temperatures of 65 ̊ and 110 ̊C basically ended after three months 

under unsealed conditions and after 1 to 2 years exposure for sealed conditions as the moisture condition stabilized, 

the decrease in compressive strength was somewhat greater for sealed specimens than for unsealed specimens, 

decreases in compressive strength were somewhat greater for specimens containing blast furnace slag, modulus of 

elasticity for sealed and unsealed specimens exhibited similar decreases with increasing exposure temperature, the 

decrease in modulus for unsealed specimens tended to stabilize after about 2 years exposure, the decrease in 

modulus was somewhat larger for specimens containing blast furnace slag, weight loss increased as the temperature 

increased, weight loss for unsealed specimens tended to stabilize after 1 month exposure while for sealed specimens 

the weight loss increased for about 1.5 years before stabilizing, weight loss results seemed to correlate well with 

decrease in modulus of elasticity in that as the weight loss increased the modulus decreased, specimens containing 

blast furnace slag tended to exhibit a slightly higher weight loss than that exhibited by reference concrete, and the 

tensile strength ratio tended to decrease for exposures up to 2 years and then increase slightly. The study 

recommended that consideration should be given to increasing the current temperature limit (65 ̊C) for concrete 

utilized in facilities for storage of spent nuclear fuel. A limit of up to about 110 ̊C for the physical properties is 

possible as long as the temperature dependence of the mechanical properties is taken into consideration. Since 

evaporation of gel water increases markedly at temperatures greater than 110 ̊C and elevated temperature 

considerations for design of reinforced concrete structures (e.g., creep and bond between concrete and steel 

reinforcement) become more pronounced at temperatures greater than 100 ̊C, it was considered reasonable to raise 

the limit to 85 ̊C. 

Figure 2.209 Residual compressive strength vs exposure time: unsealed specimens.

Source: T. Kanazu, T. Matsumura, and T. Nishiuchi, ―Changes in the Mechanical Properties of Concrete Subjected 

to Long-Term Exposure to High Temperatures,‖ Report No. U95037, Abiko Research Laboratory, Central 

Research Institute of Electric Power Industry, Japan, March 1996.

 

No 8-year results were obtained for the sealed specimens.
152

Figure 2.210 Residual compressive strength vs exposure time: sealed specimens.

Source: T. Kanazu, T. Matsumura, and T. Nishiuchi, ―Changes in the Mechanical Properties of Concrete Subjected 

to Long-Term Exposure to High Temperatures,‖ Report No. U95037, Abiko Research Laboratory, Central 

Research Institute of Electric Power Industry, Japan, March 1996.

Figure 2.211 Residual modulus of elasticity vs exposure time: unsealed specimens.

Source: T. Kanazu, T. Matsumura, and T. Nishiuchi, ―Changes in the Mechanical Properties of Concrete Subjected 

to Long-Term Exposure to High Temperatures,‖ Report No. U95037, Abiko Research Laboratory, Central 

Research Institute of Electric Power Industry, Japan, March 1996.
153

Figure 2.212 Residual modulus of elasticity vs exposure time: sealed specimens.

Source: T. Kanazu, T. Matsumura, and T. Nishiuchi, ―Changes in the Mechanical Properties of Concrete Subjected 

to Long-Term Exposure to High Temperatures,‖ Report No. U95037, Abiko Research Laboratory, Central 

Research Institute of Electric Power Industry, Japan, March 1996.

Figure 2.213 Change in weight vs exposure time: unsealed specimens.

Source: T. Kanazu, T. Matsumura, and T. Nishiuchi, ―Changes in the Mechanical Properties of Concrete Subjected 

to Long-Term Exposure to High Temperatures,‖ Report No. U95037, Abiko Research Laboratory, Central 

Research Institute of Electric Power Industry, Japan, March 1996.
154

Figure 2.214 Change in weight vs exposure time: sealed specimens.

Source: T. Kanazu, T. Matsumura, and T. Nishiuchi, ―Changes in the Mechanical Properties of Concrete Subjected 

to Long-Term Exposure to High Temperatures,‖ Report No. U95037, Abiko Research Laboratory, Central 

Research Institute of Electric Power Industry, Japan, March 1996.

Figure 2.215 Change in splitting-tensile strength ratio vs exposure time: sealed specimens.

Source: T. Kanazu, T. Matsumura, and T. Nishiuchi, ―Changes in the Mechanical Properties of Concrete Subjected 

to Long-Term Exposure to High Temperatures,‖ Report No. U95037, Abiko Research Laboratory, Central 

Research Institute of Electric Power Industry, Japan, March 1996.

Summary. Results for exposure of concrete to elevated temperature for relatively long periods of time are limited 

and tend to indicate that strength and modulus of elasticity decrease with increasing exposure time at elevated 

temperature (T > 110 ̊C), with most of the decrease occurring relatively soon. Modulus of elasticity is affected more 

than compressive strength. At relatively low elevated temperatures (T < 110 ̊C) the elevated temperature exposure 

can result in an increase in compressive strength, especially for sealed conditions due to autoclave curing, with the 

modulus of elasticity being relatively unaffected for sealed conditions but dropping for unsealed conditions. At 
155

moderate elevated temperatures (e.g., 230 ̊C) unsealed specimens exhibited improved performance relative to sealed 

specimens, possibly due to saturated steam pressure that develops in a closed system to cause deterioration of 

structural properties of cement gel. Partial replacement of the cement with fly ash or blast furnace slag appears to 

impart improved performance at elevated temperature. Sealed specimens performed better under thermal cycling 

than unsealed specimens, and thermal cycling has a greater effect on modulus of elasticity than compressive 

strength.

2.2.1.11 Radiation Shielding Effectiveness

Information and Data. Portland cement concrete possesses many of the physical qualities of an ideal radiation 

shield. It is a polyphase material consisting of particles of aggregate contained in a matrix of Portland cement paste. 

Gamma rays are absorbed by the high-density-aggregate materials and neutrons are attenuated by hydrogen atoms in 

the cement paste. A concrete shield is exposed to two sources of heat: heat transferred from hot parts of the reactor 

system and heat produced internally by the attenuation of neutrons and gamma rays [2.196]. Energy captured from 

slowed down fast neutrons and gamma rays entering the shield from the reactor core is deposited within the shield 

material and liberated as heat. The total amount of heat generated can be considerable. The heat generated may 

have detrimental effects on the physical, mechanical, and nuclear properties of the concrete. Different types of 

concrete perform differently under radiation exposure, although if heated to relatively high temperatures they all will 

lose waters of crystallization and become somewhat weaker and less effective in neutron attenuation [2.197].

Provided below is a brief summary of the effect of elevated temperature exposure on shielding effectiveness. 

The effectiveness of concrete as a shield may be reduced under service conditions (elevated temperature) as drying 

reduces the hydrogen content or cracking occurs. Figure 2.216 presents results of elevated-temperature exposure on 

shielding of heavyweight aggregate (iron limonite and magnetite limonite) concretes [2.198]. Significant changes in 

Figure 2.216 Thermal neutron distribution in ordinary concrete as a function of temperature.

Source: E. G. Peterson, ―Shielding Properties of Ordinary Concretes as a Function of Temperature,‖ HW-65572, 

Hanford Atomic Products Operation, Richland, Washington, August 2, 1960.
156

attenuation effectiveness were found as the concrete was heated to 100°C and 200°C, with little additional change as 

a result of heating to 300°C. Despite the loss of neutron and gamma attenuation efficiency with increasing 

temperature, it was concluded that the concrete would serve as a satisfactory shield material. If increasing efficiency 

were required at higher temperatures, it could be accounted for in the design. 

Shielding effectiveness of concrete is also reduced if through-cracks develop. The effect of straight and crooked 

cracks through a concrete shield on gamma ray attenuation has been investigated [2.199]. In the immediate vicinity 

of the concrete surface, leakage of gamma rays through a slit contributed significantly to -dose rate, but diminished 

rapidly with distance from the surface as a result of shield thickness and scattering effects. The shielding 

effectiveness of cracked concrete has been investigated and formulas developed to define the resulting effects 

[2.200]. Guidelines developed to compensate for cracking note that it might be economically advantageous to allow 

a concrete shield to crack and then shield the resulting radiation by other means.

Additional information on attenuation, materials for radiation-shielding concrete, nuclear properties of concrete, 

water content and its effect on nuclear properties of concrete, and basic design of radiation shields is available 

[2.21].

Summary. The presence of hydrogen in a concrete radiation shield is important, particularly with respect to neutron 

attenuation. Thermal loadings that can affect the concrete water content can reduce the concrete’s radiation 

shielding effectiveness. Temperature limits have been recommended to protect ordinary and heavyweight concrete

shields against water loss, the effect of elevated temperature, radiation-absorption effects, and thermal stress 

conditions [2.201]. For neutron shields maximum internal and ambient temperatures are noted as 88 ̊ and 71 ̊C, 

respectively, and for gamma shields the maximum internal and ambient temperatures are noted as 177 ̊ and 149 ̊C, 

respectively,

2.2.1.12 Multiaxial Conditions

Information and Data. In large structures such as prestressed concrete pressure vessels, once the post-tensioning is 

applied, the concrete is stressed either biaxially or triaxially, with the biaxial case, particularly tension-compression, 

being of most concern. Regulatory documents in the United States, United Kingdom, France, and Germany 

generally consider multiaxial stress states by allowing use of stresses that are somewhat higher than would be 

permitted for uniaxial loading conditions. A summary of guidelines provided in these codes is available [2.202]. 

As most of the multiaxial testing of concrete has been conducted at room temperature, a brief review of this 

information is presented first to provide an indication of the effect of multiaxial loading on concrete behavior. More 

details on these results are available [2.103,2.202]. This is followed by information on more recent results and 

multiaxial loading of concrete at elevated temperatures.

A number of studies have been conducted investigating the behavior of concrete at room temperature under 

multiaxial loading, primarily using unsealed specimens [2.203-2.217]. A review of information contained in many 

of these documents is available [2.103]. In the review it is noted that it is generally accepted that the ultimate 

strength of concrete increases when sustaining multiaxial stresses are imposed. The method of load application is a 

primary topic of research, and results obtained with steel-brush platens, ball-bearing platens, fluid-cushion platens, 

and deformable platens have shown promise for reducing end restraints to tolerable levels so that the true ultimate 

strength can be measured (see Figure 2.217). The most reliable data indicate a maximum biaxial compressive 

strength of 1.25 times the uniaxial strength when measured with the same apparatus. Triaxial tests indicate strengths 

up to four to six times the uniaxial strength. Tests with sealed specimens are very limited. Results indicate that the 

sealed specimens are weaker than dry concrete, but that wet concrete gives a greater relative increase in strength 

under multiaxial stress conditions. An important observation concerns that of a supposed critical stress level under 

biaxial loading, a stress that may occur as low as 50% of the ultimate and at which severe permanent damage takes 

place within the specimen. Since uniaxial strength is generally used as a material property for design, one may 

suppose that the actual vessel conditions involving biaxial stresses will make that data conservative by 25%. 

However, tests conducted with steel platens, the common method, indicate a strength measurement that is 25% too 

high, meaning that there may be, essentially, no additional margin for strength data obtained with steel platens. The 

additional observation of a critical stress below ultimate for short- and long-term loading may be more realistic for 

design purposes. The review concludes noting that additional work needs to be done to develop a better 
157

understanding of multiaxial behavior as well as development of improved methods of providing accurate strength 

data, especially with regard to sealed concrete and elevated temperatures. 

Since the above review, an international program involving researchers from the Federal Republic of Germany, 

Italy, United Kingdom, and United States has been conducted [2.218]. The program’s objective was to provide 

insight into the significance of systematic effects on concrete multiaxial test results. All participants in the program 

used specimens from the same concrete and mortar mix that were: (1) mixed, cast, and cured at one laboratory; 

(2) shipped under controlled conditions to the other laboratories; and (3) tested at an identical age. The testing 

program was divided into two parts: (1) deformations and strength of concrete under biaxial loading, applied 

monotonically and proportionally, with stress ratios 2/ 1 = 0/3, 1/3, 2/3, and 3/3; and (2) deformations and strength 

of concrete under triaxial loading with stress increased hydrostatically up to one of several specified octahedral 

shear planes ranging from 75 to 200% of the uniaxial strength and then deviated within that plane along one of three 

stress paths - triaxial compression, constant intermediate principal stress, or triaxial extension to failure. Uniaxial 

mortar cylinder compressive strengths ranged from 29.9 to 37.2 MPa and concrete compressive strengths from 29.7 

to 34.7 MPa. Figure 2.217 presents schematics of the different multiaxial test methods used in the investigation. 

Figure 2.217 Multiaxial test methods.

Source: K.H. Gerstle, D.L. Linse, P. Bertacchi, M.S. Kotosovos, H-Y. Ko, J.B. Newman, P. Rossi, G. Schickert, 

M.A. Taylor, L.A. Traina, R.M. Zimmerman, and R. Bellotti, ―Strength of Concrete Under Multiaxial 

Stress States,‖ Paper SP 55-5 in Douglas McHenry International Symposium on Concrete and Concrete 

Structures, pp. 103-131, American Concrete Institute, Farmington Hills, Michigan, 1978.

Biaxial and triaxial results for concrete cube specimens are presented in Figure 2.218. Figure 2.218b presents 

triaxial failure envelopes within the octahedral plane 0 = 34.5 MPa. Most of the triaxial tests were performed by 

first subjecting the specimen to a specified hydrostatic pressure, ranging from 20.7 to 55.2 MPa, followed by 

deviating within the octahedral plane along one of the three stress paths shown in the figure. The following 

conclusions were derived from the investigations: (1) a systematic relationship exists between the degree of 

constraint of the loading system and the uniaxial and multiaxial strengths (differences in strength found in previous 

investigations at least partially can be ascribed to differences in loading systems) and (2) decomposition of the 

failure stress state into hydrostatic and deviatoric portions appears to offer a systematic approach to the development 
158

of failure criteria for concrete and mortar under mu1tiaxia1 stresses (establishment of a common failure criterion for 

uniaxial, biaxial, and triaxial loading conditions that is path-independent seems possible on the basis of test results).

(a) Biaxial loading. (b) Triaxial loading.

Figure 2.218 Multiaxial test results: (a) strength of concrete under biaxial loading and 

(b) triaxial failure envelopes within octahedral plane 0 = 34.5 MPa.

Source: K.H. Gerstle, D.L. Linse, P. Bertacchi, M.S. Kotosovos, H-Y. Ko, J.B. Newman, P. Rossi, G. Schickert, 

M.A. Taylor, L.A. Traina, R.M. Zimmerman, and R. Bellotti, ―Strength of Concrete Under Multiaxial 

Stress States,‖ Paper SP 55-5 in Douglas McHenry International Symposium on Concrete and Concrete 

Structures, pp. 103-131, American Concrete Institute, Farmington Hills, Michigan, 1978.

The behavior of high-performance concrete under biaxial loading has been investigated [2.219]. The concrete mix 

(fc’ = 71.3 MPa at 28d) incorporated limestone coarse aggregate, siliceous sand, and three cementitious materials: 

Type I cement, fly ash, and ground granulated blast furnace slag. The test specimens were 150 x 150 x 50 mm3. 

Interfaces between the loading platens and specimens were lubricated to avoid restraint. Figure 2.219 presents 

typical complete stress-strain curves for biaxial compression loading and tension-compression loading for different 

nominal strain ratios. The shape of the biaxial compression stress-strain curves was found to be similar to results for 

the uniaxial curves in that they both had ascending and descending parts. The biaxial tension-compression stressstrain curves in the tensile direction were found to be similar to those observed under uniaxial tension loading. The 

multiaxial strength and the strength envelope are an important basis for the analysis and design of concrete 

structures. Specimens under tension-compression loading were found to be only damaged in the tensile direction 

with a split developing parallel to the compression direction. Figure 2.220 presents biaxial strength envelopes in 

stress space and strain space. The biaxial peak strains for most cases were found to be lower than those obtained for 

uniaxial compression loading, primarily due to the strength increase in the compression region under biaxial loading 

and an associated reduction in deformation capacity. It was concluded that the strength increase in the biaxial 

compression region is accompanied by strain softening which needs to be considered in the analysis and design of 

concrete structures.
159

Biaxial compression

(a) = 1, (b) = 0.3, (c) = 0.1

Tension-compression

(a) = -0.167, (b) = -0.25, (c) = -0.5

Figure 2.219 Complete stress-strain curves for biaxial compression for different strain ratios ( ).

Source: W. Ren, W. Yang, Y. Zhou, and J. Li, ―Behavior of High-Performance Concrete Under Uniaxial and 

Biaxial Loading,‖ Title No. 105-M62, ACI Materials Journal, American Concrete Institute, Farmington 

Hills, Michigan, November-December 2008.
160

Stress Space Strain Space

Figure 2.220 Biaxial strength envelopes in stress space and strain space.

Source: W. Ren, W. Yang, Y. Zhou, and J. Li, ―Behavior of High-Performance Concrete Under Uniaxial and 

Biaxial Loading,‖ Title No. 105-M62, ACI Materials Journal, American Concrete Institute, Farmington 

Hills, Michigan, November-December 2008.

In another study, an investigation was conducted to provide experimental failure behavior characteristics under 

biaxial stresses for plain concrete of a Korean nuclear containment building [2.220]. Type I Portland cement 

concretes of two strength were investigated: (1) 37.8 MPa corresponding to design strength of wall and dome 

concrete of containment and (2) 27.5 MPa corresponding to base concrete. Three combinations of biaxial loading 

(compression-compression, compression-tension, and tension-tension) and four stress-ratio levels ( = f2/f1) were 

considered in each loading regime as noted in Figure 2.221. Concrete plate specimens 200 x 200 x 60 mm3 were 

used in the study. Teflon pads were placed at the interfaces between the loading system and specimen to minimize 

frictional effects. Tensile loading was applied by epoxying the platens to the specimen surfaces. Biaxial tests were 

performed under constant stress ratio. Figure 2.222 presents the biaxial strength envelopes for the two concretes 

Figure 2.221 Stress ratios considered(f2/f1).

Source: S-K. Lee, Y-C. Song, and S-H. Han, ―Biaxial Behavior of Plain Concrete of Nuclear Containment 

Building,‖ Nuclear Engineering and Design 227, pp. 143-153, 2004.
161

Figure 2.222 Biaxial strength envelopes.

Source: S-K. Lee, Y-C. Song, and S-H. Han, ―Biaxial Behavior of Plain Concrete of Nuclear Containment 

Building,‖ Nuclear Engineering and Design 227, pp. 143-153, 2004.

investigated. Figure 2.223 presents biaxial stress-strain curves that were developed for the three different loading 

conditions for the wall and dome concrete materials. The stress-strain curve for specimens subjected to equal 

biaxial compression (f2/f1 = -1/-1) tends to have an initial slope that is relatively steep whereas for uniaxial 

compression (f2/f1 = 0/-1) the slope is much less indicating the effect of increased confinement under biaxial 

loading. Stress-strain curves for combined tension-compression show that the principal stress in the compressive 

direction (f1) decreases as the absolute value of the stress ratio (f2/f1) increases due to the principal tensile stresses in 

the orthogonal direction increasing more due to the increase of absolute value of f2/f1 (principal compressive stress 

decrease). Typical crack patterns and failure modes are presented in Figure 2.224 and are similar to those obtained 

by other researchers. The fracture of concrete under uniaxial compression, Figure 2.224a, was determined by 

formation of a major crack. As the stress ratio increased, Figures 2.224b-d, the major cracks become inclined at 

increasing angles to the direction of applied loading. In the compression-tension region three different fracture 

mode patterns occurred. In Figure 2.224e a major crack formed at a 45 ̊ angle to the direction of applied 

compressive load, as well as many tensile cracks perpendicular to the direction of applied load. As the stress ratio 

increased, Figure 2.224f, fewer cracks formed perpendicular to the direction of applied load. Figures 2.224g and h 

exhibit only one apparent tensile crack. Figure 2.224i presents results for uniaxial tension in which the specimen 

failed by development of one major crack perpendicular to the direction of applied load. Figures 2.224j to l show a 

major crack developed at an approximate 45 ̊ inclination to the direction of maximum principal stresses as the stress 

ratio decreases to 1.0. 
162

Figure 2.223 Stress-strain relationships for wall and dome concrete under: (a) biaxial compression, 

(b) combined compression and tension, and (c) biaxial tension.

Source: S-K. Lee, Y-C. Song, and S-H. Han, ―Biaxial Behavior of Plain Concrete of Nuclear Containment 

Building,‖ Nuclear Engineering and Design 227, pp. 143-153, 2004.
163

Figure 2.224 Failure modes under biaxial loading for different stress ratios.

Source: S-K. Lee, Y-C. Song, and S-H. Han, ―Biaxial Behavior of Plain Concrete of Nuclear Containment 

Building,‖ Nuclear Engineering and Design 227, pp. 143-153, 2004.

Other researchers indicated that concrete under different biaxial loading conditions (e.g., compression-compression, 

compression-tension, and tension-tension) could be described by relatively simple mathematical relations [2.213].

The relations developed for the different loading regions are:
164

(compression-compression region) (2.67)

(compression-tension region) (2.68)

(tension-tension region) (2.69)

where,

= principal stresses,

= uniaxial compressive strength, and 

= uniaxial tensile strength.

The equations were obtained by breaking down the stress and strain states into hydrostatic and deviatoric 

components that could be used in conjunction with finite-element methods and force-deformation relations to 

provide a means to perform nonlinear analyses of structures. 

Biaxial tests of a quartzite concrete (fc’ = 35.5 to 45.0 MPa) and mortar (fc’ = 35.0 MPa) using plate-shaped 

specimens (200 200 50 mm3) were conducted [2.221]. After a 2-h hold time at temperatures of 150 ̊, 300 ̊, 450 ̊, 

or 600 C, the unsealed specimens were loaded to failure at a constant displacement rate using brush platens. 

Specimen heating was applied at the free surfaces. A comparison of uniaxial (x = 0) and biaxial (x = 0.7) results for 

tests at 300 C and 600 C is shown in Figure 2.225. Results have been normalized with respect to 

Figure 2.225 Influence of aggregate content and water-cement ratio on 

uniaxial and biaxial results for tests at 300°C and 600°C.

Source: K.-Ch. Thienel and F. S. Rostásy, ―Influences of Concrete Composition on Strength and Deformations 

Under Uniaxial and Biaxial Loading at Elevated Temperature,‖ Paper H04/6 in Transactions of the 12th

International Conference on Structural Mechanics in Reactor Technology, pp. 145–150, 1993.

corresponding strength values at room temperature and the same stress ratio. The tests exhibited a smaller decrease 

in compressive strength for concrete with higher aggregate contents at both temperatures. Under biaxial loading, an 

increase in strength even appears at 300°C for the leaner mix. The effect of water-cement ratio on strength for 

uniaxially and biaxially loaded specimens is also presented in Figure 2.225. These results also have been 
165

normalized and indicate that initially the water-cement ratio did not affect the temperature-dependent strength losses

at 300 ̊C. After decomposition of the Portlandite (450 ̊ to 550 ̊C) the decrease in strength was at a lower rate for the 

concrete (QB3) made with a higher water-cement ratio. As the temperature increased, the influence of water-cement 

ratio decreased resulting in similar strength values for both concretes at 600 ̊C. The uniaxial strength of mortar was 

reduced less than that of concrete at temperatures above 300°C. The influence of aggregate content on residual 

modulus of elasticity at 300 ̊ and 600 ̊C is presented in Figure 2.226. At 300 ̊C the modulus is reduced more for the 

leaner concrete mix, however, at 600 ̊C the differences are on the same order for the two mixes. If a higher watercement ratio is used, Figure 2.226, a similar trend is observed but the difference in modulus for the two concretes is 

greater at 300 ̊C. For ultimate strain in the direction of the main loading axis it was noted that the ultimate strain 

increases with increasing temperature, and at temperatures up to 300 ̊C there was a tendency for lower strain values 

for the concrete with higher aggregate contents. Biaxial compression was noted to result in higher strain values 

throughout the temperature range. The influence of stress ratio on the compressive strength for various temperatures 

is shown in Figure 2.227. The stress ratio of the compressive strength was varied between 0 and 1, where the stress 

ratio represents the ratio of the applied stresses in the two principal directions. The solid and dashed lines in the 

figure represent the mean biaxial behavior (predicted) at each temperature noted. The failure envelopes are similar 

for each temperature with increasing temperature producing a larger loss of biaxial strength. The difference between 

uniaxial and biaxial strength increased as the temperature increased. The temperature-dependent decrease in 

strength was affected by the composition of the concrete in the entire range of biaxial compressive stress. The 

maximum aggregate size had a significant influence on behavior, while the aggregate content and water-cement 

ratio were less influential.

Figure 2.226 Influence of aggregate content and water-cement ratio on 

modulus of elasticity results for tests at 300°C and 600°C.

Source: K.-Ch. Thienel and F. S. Rostásy, ―Influences of Concrete Composition on Strength and Deformations 

Under Uniaxial and Biaxial Loading at Elevated Temperature,‖ Paper H04/6 in Transactions of the 12th

International Conference on Structural Mechanics in Reactor Technology, pp. 145–150, 1993.

Experimental investigations at temperatures to 600 ̊C were made to determine the strengths and deformations of 

concretes subjected to biaxial compressive stresses [2.68]. Figure 2.37 presents the test frame used to conduct 

biaxial tests of 200 by 200 by 50 mm3gravel (quartz) concrete specimens [62.0 N/mm2(cube strength)] [2.68,2.69]. 

Load was applied to the specimen using four servohydraulic jacks. The tests were conducted at temperature using 

unsealed specimens. Figure 2.228 presents the effect of loading platen (brush or rigid) on the concrete biaxial 

strength. The brush platens provided a reduction in strain inhibition and a corresponding decrease in strength 

compared to rigid platen results. The effect of elevated temperature on concrete biaxial strength is presented in 
166

Figure 2.227 Biaxial compressive strength at different temperatures.

Source: K.-Ch. Thienel and F. S. Rostásy, ―Strength of Concrete Subjected to High Temperature and Biaxial Stress; 

Experiments and Modeling,‖ Materials and Structures 28, pp. 575–580, 1995.

Figure 2.228 Effect of loading platen type on concrete biaxial strength.

Source: C. Ehm and U. Schneider, ―The High Temperature Behaviour of Concrete Under Biaxial Load,‖ Cement

and Concrete Research 25, pp. 27-34, 1985.
167

Figure 2.229 where the data are reported as a fraction of the unconfined uniaxial compressive strength at 20 ̊C. 

Results show that the strength of concrete under biaxial compression is higher than under uniaxial compression. At 

20 ̊C the strength ratio reaches a maximum of 1.34 with the relative increase in ultimate strength at this stress ratio 

being more marked at higher temperatures. In uniaxial loading the ultimate failure occurred by splitting in the plane 

parallel to the load and perpendicular to the specimen face, while in biaxial compression loading the ultimate failure 

occurred by splitting along planes parallel to the load and parallel to the specimen free surface. Deformation results 

were presented previously in Section 2.2.1.1.

Figure 2.229 Effect of temperature on concrete biaxial strength.

Source: C. Ehm and U. Schneider, ―The High Temperature Behaviour of Concrete Under Biaxial Load,‖ Cement

and Concrete Research 25, pp/ 27-34, 1985.

Biaxial tension-compression tests (residual) at elevated temperature (i.e., 20 ̊, 200 ̊, 300 ̊, 500 ̊, and 600 ̊C) have 

been conducted [2.70]. Specimens 100 by 100 by 100 mm3 were fabricated using a 20-mm-maximum-size crushed 

limestone aggregate concrete having reference compressive and tensile strengths of 32.5 and 3.14 MPa, respectively. 

The specimens were held at temperature for six hours and then permitted to slowly cool to room temperature prior to 

testing. Stress ratios ( 2/ 3) investigated included 0 (uniaxial), -0.1, -0.25, -0.5, and -0.75. Tensile loads were 

applied by gluing the specimen surface to the tensile loading platens. Figure 2.230 presents the change in tensile 

strength with temperature for the stress ratios investigated. For each stress ratio all tensile strengths follow the same 

pattern. A regression equation between tensile strength and temperature, also shown in the figure, was developed:

, (2.70)

where, is the tensile strength for any stress ratio corresponding to temperature t,

is tensile strength at 20 ̊C and for any stress ratio, and

t = temperature.
168

Figure 2.230 Change in tensile strength with temperature.

Source: Y. Song, A. Zhang, L. Qing, and C. Yu, ―Biaxial Tensile-Compressive Experiment on Concrete at High 

Temperatures,‖ Frontiers of Architecture and Civil Engineering in China 1(1), pp. 94-98, 2007.

The relationship between tensile strength and stress ratio and temperature is presented in Figure 2.231 and shows 

that at each temperature the tensile strength decreases with the decline in absolute values of the stress ratios. 

Figure 2.231 Effect of temperature on the relationship between tensile strength and stress ratios.

Source: Y. Song, A. Zhang, L. Qing, and C. Yu, ―Biaxial Tensile-Compressive Experiment on Concrete at High 

Temperatures,‖ Frontiers of Architecture and Civil Engineering in China 1(1), pp. 94-98, 2007.

A two-linear regression equation between tensile strength, stress ratio, and temperature was developed:

-0.4177 < < - 0.1 (2.71)

where, is the tensile strength at t ( ̊C) for various stress ratios. The failure envelopes under biaxial tensilecompressive stresses derived from the experimental data are presented in Figure 2.232. For every temperature all 
169

envelopes are somewhat concave to the axis. A linear regression failure equation has been developed 

from the data: 

, where (2.73)

and are compressive and tensile strengths, respectively, and is the uniaxial compressive strength 

for various temperatures whose value can be estimated from

, where (2.74)

is the concrete uniaxial compressive strength at room temperature. Conclusions of the study for concrete under 

biaxial tension-compression loading were that: at high temperature sudden tensile failure occurs for stress ratios less

than or equal to -1.0; both the absolute value of tensile and compressive strengths decrease with increasing 

temperature with a drop in absolute values of tensile strain ( 1) and compressive strain ( 3) being similar to that of 

strength; and the tensile strength increases with a decrease in stress ratio and reaches a maximum value at the 

uniaxial tensile stress state, but the absolute value of compressive strength (and compressive strain 3) decreases 

with a decrease in stress ratio and reaches a maximum value at the uniaxial compressive stress state.

Figure 2.232 Effect of temperature on the concrete biaxial tension-compression failure envelopes.

Source: Y. Song, A. Zhang, L. Qing, and C. Yu, ―Biaxial Tensile-Compressive Experiment on Concrete at High 

Temperatures,‖ Frontiers of Architecture and Civil Engineering in China 1(1), pp. 94-98, 2007.

Elsewhere for biaxial testing of concrete at elevated temperature it was noted that when the principal stress ratio 

remains constant, the biaxial strength of the concrete decreased with an increase in temperature [2.222]. Loss of 

strength was observed at 150°C under biaxial conditions but was much lower than that obtained under uniaxial 

loading. Under the same elevated temperature, the biaxial strength of concrete varied with different values of 

principal stress ratio. When the principal stress ratio equaled 0.5, the biaxial strength of concrete at high temperature 

reached a maximum.

Triaxial testing of concrete at elevated temperature was conducted using solid and hollow cylinders [2.223]. The 

specimens were either in a sealed or unsealed condition after exposure to temperatures up to 150 ̊C. Triaxial stress 

was imposed on the specimens by applying a hydrostatic pressure of 3.31 MPa during axial loading. The moisture 

loss varied directly with the strength after heat exposure. The sealed concrete strengths were reduced to about 70% 

and 60% of the reference strength at 100 ̊ and 150 ̊C, respectively.
170

An apparatus for testing concrete under multiaxial compression loading at elevated temperature has been recently 

developed at the University of Sheffield [2.224]. The apparatus (Figure 2.233) can subject 100 mm cubic concrete 

specimens to true multiaxial compression (s1 ≠ s2 ≠ s3) up to 400 MPa at temperatures to 300 ̊C. Forces are delivered 

through three independent loading frames equipped with servo-controlled hydraulic actuators creating uniform 

displacement boundary conditions via rigid platens. Displacements are measured to an accuracy of 10-6 m using a 

system of six laser interferometers. Complex experiments can be conducted involving (1) load and temperature 

cycling, (2) small stress probes, and (3) arbitrary (pre-defined) loading paths. The apparatus has been used to 

investigate uniaxial, biaxial, and hydrostatic compression in partially sealed conditions, at temperatures up to 250 ̊C 

of an ordinary Portland cement quartz-diorite concrete [2.225]. The elevated temperature tests were performed in 

two phases: (1) conditioning and (2) deviatoric loading. In the conditioning phase the specimens were

heated to different heat-load regimes [e.g., heated without load and held under constant temperature (HS), loaded 

then heated and held under steady-state conditions (LHS), heated without load then loaded and held under constant 

temperature and load (HLS), or loaded first and subjected to a thermal cycle followed by a steady-state period 

(LHCHS)]. Results of the conditioning phases were used to determine the load-induced thermal strain defined as 

the difference between the strains recorded in LHS and HS tests. The effects of the different conditioning regimes 

on the properties of the material under multiaxial compression loading were investigated during the deviatoric 

loading phase performed at a constant temperature of 250 ̊C. The reference only presents results of the conditioning 

phase (i.e., load-induced thermal strain for concrete heated at two different rates under uniaxial, biaxial, and 

hydrostatic compression of either 13 MPa or 26 MPa). Figure 2.234 provides a summary of the averaged strains 

recorded in all the tests (i.e., load-induced thermal strain and shrinkage, and load-induced thermal strain). Results 

indicate that: the load-induced thermal strain is a function of load, temperature, and moisture movement; and only 

occurs on first heating of loaded concrete to a given temperature and not during cooling or upon second heating to 

the same temperature, but if the second heating cycle exceeds the first the strains start to increase again at a similar 

rate to that of the first heating cycle.

Figure 2.233 mac2T apparatus for multiaxial compression of concrete at elevated temperature.

Source: M. Petkovski, R.S. Crouch, and P. Waldron, ―Apparatus of Testing Concrete Under Multiaxial 

Compression at Elevated Temperature (max2T),‖ Experimental Mechanics 46, pp. 387-398, 2006.
171

Figure 2.234 Average temperature-strain relationships from all LHS and HS tests: 

(a) load-induced thermal strain and shrinkage, (b) load-induced thermal strain.

Source: M. Petkovski and R.S. Crouch, ―Strains Under Transient Hygro-Thermal States in Concrete Loaded in 

Multiaxial Compression and Heated to 250 ̊C,‖ Cement and Concrete Research 38, pp. 586-596, 2008.

Summary. Several conclusions can be derived from the multiaxial strength results. Multiaxial strength and strength 

envelopes are an important basis for the analysis and design of concrete structures. The platen-interface constraint 

influences test results (i.e., strength increases as the constraint increases). Ultimate strength of concrete at room 

temperature increases under multiaxial compressive loading with maximum biaxial compressive strength being 

about 1.25 times uniaxial compressive strength and triaxial compressive strength being four to six times uniaxial 

compressive strength. Under essentially biaxial conditions where stresses in two directions are substantially larger 

than the third, there is little increase in strength and if the stress in the third direction becomes tensile the biaxial 

strength is reduced below the uniaxial strength [2.135]. Sealed specimens are weaker than dry concrete, but wet 

concrete gives a greater relative increase in strength under multiaxial stress conditions. For biaxial loading at 

elevated temperature the maximum aggregate size had a significant influence on results, the aggregate content and 

water-cement ratio generally had second-order effects, the failure envelopes although similar for each temperature as 

the temperature increased the loss of biaxial strength increased, and when the principal stress ratio remained 

constant the biaxial strength of the concrete decreased with an increase in temperature. Difference between uniaxial 

and biaxial strength increased as the temperature increased. Under elevated temperature biaxial tensioncompression loading, the tensile strength decreased with the decline in absolute values of the stress ratios at each 

temperature.
172

2.2.2 Physical Properties and Thermal Effects

The physical, or thermal, properties of concrete are important both in the planning of mass concrete construction 

(thermal changes) and the dissipation of heat build up during operation. The thermal expansion coefficient and 

modulus of elasticity data are prerequisite to computing the thermoelastic stresses that result from nonuniform 

heating. Desirable thermal properties identified for a prestressed concrete pressure vessel (PCPV) are summarized 

in Table 2.8.

Table 2.8 Preferred concrete physical properties

Preferred property Reason

High density To provide good neutron and gamma ray absorption properties

Low elastic and creep deformation 

under load

To reduce movements and the redistribution of stresses under varying 

load and temperature cycles

To reduce prestress losses

Low drying shrinkage To reduce movements and temperature stresses

Low thermal expansion To reduce movements and temperature stresses

Resistant to thermal shock To prevent damage to structure under rapid heating application (i.e., 

adjacent to steam penetrations)

High thermal conductivity To minimize the cooling system requirements to keep vessel concrete at a 

permissible level (PCPVs)

Source: R.D. Browne, ―Properties of Concrete in Reactor Vessels,‖ Proceedings of the Conference on Prestressed 

Concrete Pressure Vessels, Group C, Paper 13, pp. 131-151, Institute of Civil Engineers, London, United 

Kingdom, 1967.

As water is added to cement, an exothermic chemical reaction takes place. If the heat is generated at a faster rate 

than it can be dissipated, a temperature rise occurs. Factors affecting the amount and rate of heat generated during 

this reaction are the cement type, temperature at placement, water-cement ratio, and cement content. In mass 

concrete structures where there can be significant heat build up, cracking can occur upon cooling because the 

exterior of the structure will cool faster than the interior. However, by using low (Type IV) or moderate (Type II) 

heat of hydration cements and following the procedures recommended in Ref. 2.226, this problem can be 

minimized. 

Under normal conditions, most concrete structures are subjected to a range of temperature no more severe than that 

imposed by ambient environmental conditions. However, there are important cases where these structures may 

experience much higher thermal loadings. For example, local temperature loadings (hot spots) can develop at 

internal faces of a structure due to local loss of insulation, jet impingement (following pipe rupture), or hot 

penetrations. Internal gas temperatures under accident conditions operate on the lined faces of containments and 

internal structures resulting in a temperature distribution set up across a wall that can lead to significant internal 

forces and moments as well as affecting material properties of concrete that in turn can affect concrete creep and 

relaxation of post-tensioning tendon forces (where applicable). 

Concrete’s thermal properties are more complex than for most materials because not only is the concrete a 

composite material whose components have different properties, but also its properties depend on moisture content 

and porosity. Temperature variations produce expansions or contractions of concrete structures. If movement of a 

structure is restrained, significant internal stresses can develop, thus leading to cracking, distortions, or even 

destruction. In general, the density, conductivity, and diffusivity of concrete will increase with an increase in 

temperature. The coefficient of thermal expansion is used as a measure of the volume change of a material 

subjected to a temperature differential. Dissipation of heat is important to nuclear power plant structures such as a 

PCPV because it affects the development of thermal gradients and the resulting thermal stresses. The basic 

quantities involved are (1) the coefficient of thermal conductivity , (2) the thermal diffusivity D, and (3) the 

specific heat cp. These quantities are related by the term D = / cp (m2/s). 

 

Concrete structures that experience radiation also can develop thermal gradients as a result of radiation absorption.
173

2.2.2.1 Porosity and Density

Information and Data. The loss of evaporable water does not affect porosity, however, dehydration creates 

additional pore space to effectively increase the porosity as the temperature increases [ 2.24]. The average pore size 

increases with temperature as the ultrafine gel structure of the calcium silicate hydrate is progressively destroyed by 

dehydration [2.16]. At 850 ̊C porosity is about 40% greater than it was at 105 ̊C [2.24], as noted in Figure 2.235. 

As a result of the pore size increase the bulk mass density of the cement paste decreases with temperature (i.e., from 

1.45 g/cm3at 105 ̊C to 1.3 g/cm3at 850 ̊C [2.24]).

Figure 2.235 Variation of Portland cement paste porosity with temperature.

Source:

As the volume of a given mass of concrete changes on exposure to elevated temperature, the mass density will also 

change. Changes in mass density result from thermal expansion and drying shrinkage, diffusion of water or released 

gases, and dehydration, melting, or sintering. The density of concrete depends on the density of its aggregate 

materials and its moisture content in the temperature range from 20°C to 150°C. Figure 2.236 presents variations of 

true density, bulk density, and porosity for a cement paste [2.24]. In the unsealed condition results reflect the

Figure 2.236 True density, bulk density, and porosity of cement paste versus temperature.

Source:
174

influences of chemical transformations that include water dilation up to about 80 ̊C, the loss of free and physically 

bound water at 100 ̊ to 200 ̊C depending on section size and heating rate, followed by loss of chemically-combined 

water at temperatures above 100 ̊C, the dissociation of calcium hydroxide at 400 ̊ to 500 ̊C, and decarbonation 

above 600 ̊C [2.19]. In concrete the aggregate plays an important role in terms of thermal dilation and the 

dissociation of some aggregates such as carbonate aggregates that display a significant reduction in density above 

600 ̊C and a marked increase in porosity [2.19]. Figure 2.237 presents the effect of aggregate type on the density of 

concretes in the temperature range from room temperature to 1000°C [2.19]. Storage (curing or preconditioning) 

conditions at lower temperature are also important because moist specimens will lose water and thus experience a 

greater decrease in density upon heating than specimens that have experienced drying. At temperatures from 150°C 

to 600°C the density of limestone concrete is relatively constant. At temperatures from 600 ̊ to 900 ̊C decarbonation 

of limestone commences and weight loss occurs [2.227] as the concrete porosity increases until at higher 

temperatures sintering takes place and the density may increase slightly [2.15]. Siliceous aggregates exhibit a

somewhat steeper decrease in density with heating in the range from room temperature to 700°C due to the large 

thermal expansion of the quartz. Basalt exhibits the smallest decline in density because of its lower thermal 

expansion [2.15]. Figure 2.238 presents bulk density as a function of temperature for a 38-MPa basalt aggregate 

concrete utilized for Korean nuclear power plants [2.228] and Figure 2.239 provides additional data on the density 

of normal and lightweight concretes as a function of temperature.

Figure 2.237 Density of concretes having different coarse aggregate types.

Source:

Summary. The loss of evaporable water does not affect porosity, however, as the temperature increases dehydration 

creates additional pore space to effectively increase the porosity. The average pore size increases with temperature 

as the ultrafine gel structure of the calcium silicate hydrate is progressively destroyed by dehydration, with porosity 

being about 40% greater at 850 ̊C than it was at 105 ̊C. Changes in density are related to weight changes, thermal 

dilation, and changes in porosity. In temperature range 20 ̊ to 150 ̊C the density of concrete depends on the density 

of the aggregates and their moisture content, and is affected by storage conditions, with water-stored specimens 

changing most. At temperatures greater than 150 ̊ to ~600 ̊C there is a minor change in density of limestone 

aggregate concretes, however, at temperatures greater than 600 ̊C decomposition of limestone leads to a porous 

material. Siliceous aggregate concretes exhibit a steeper decline in density than calcitic aggregate concretes for 

temperatures from 20 ̊ to 700 ̊C. Basalt aggregate concretes exhibit a smaller decrease in density with temperature 

because they have relatively low thermal expansion. 
175

Figure 2.238 Density of a nuclear power plant concrete vs temperature.

Source: K-Y. Shin, S-B. Kim, J-H. Kim, M. Chung, and P-S. Jung, ―Thermo-Physical Properties and Transient 

Heat Transfer of Concrete at Elevated Temperatures,‖ Nuclear Engineering and Design 212, pp. 233–241, 

2002.

Figure 2.239 Density of normal and lightweight concretes versus temperature.

Source: T. Z Harmathy. and L.W. Allen, ―Thermal Properties of Selected Masonry Unit Concretes,‖ Journal 

of American Concrete Institute 70, pp. 132–142, 1973.
176

2.2.2.2 Coefficient of Thermal Expansion

Information and Data. The coefficient of thermal expansion represents the volume change of a material due to 

temperature change and is expressed as a change in length per degree of temperature change. The coefficient is 

important as a measure of the structural movement and thermal stresses resulting from a temperature change that can 

lead to cracking and spalling. 

Concrete’s thermal expansion is a complicated phenomenon because of the interaction of its two main components -

cement paste and aggregate - which each have their own coefficients of thermal expansion. When differences in the 

thermal expansion between the hardened cement paste and aggregate are large, heating causes microstresses and 

microcracking that can disrupt the concrete microstructure. Observations of thermal expansion are complicated by 

various extraneous effects that accompany the temperature change (e.g., additional volume changes caused by 

changes in moisture content, chemical reactions leading to dehydration and conversion, and creep and 

microcracking resulting from non-uniform thermal stresses) [2.5].

Figure 2.240 presents data on the change of length of hardened Portland cement paste at elevated temperature. 

Initially the hardened cement paste expands on heating up to approximately 150 ̊C, the maximum expansion being 

on the order of 0.2% [2.5]. Expansion ceases between 150 ̊ and 300 ̊C and then between 300 ̊ and 800 ̊C the 

hardened cement paste shrinks with the shrinkage being 1.6 to 2.2% at 800 ̊C. The initial expansion has been 

Figure 2.240 Length change of Portland cement paste specimens at elevated temperature.

Source:

attributed to kinetic molecular movements in the cement paste plus swelling pressures caused by a decrease in 

capillary tension of water as the temperature rises [2.229]. Up to about 300 ̊C thermal shrinkage and thermal 

expansion both occur to produce a net expansion of the cement paste and at higher temperatures thermal shrinkage 

exceeds thermal expansion. Shrinkage of the cement paste results from loss of evaporable and chemically-combined 

water. 

Because the aggregate generally constitutes a major proportion of the mix, it primarily influences the resultant 

coefficient of thermal expansion. Selection of an aggregate with a low coefficient of thermal expansion may help in 

crack prevention in mass concrete. The main factor affecting the coefficient of thermal expansion of aggregate 
177

materials is the percentage by weight of silica present in the aggregate [2.230]. Aggregates with little or no silica 

(e.g., limestone) have the lowest coefficients of thermal expansion. Table 2.9 presents typical values for coefficients 

of thermal expansion for different rocks and concretes at normal temperatures. Scatter within an aggregate (rock) 

type is due to the anisotropic nature of most rocks relative to exhibiting differing expansions in different directions. 

Processed lightweight aggregates (e.g., expanded slag, shale, or clay) have lower coefficients of thermal expansion 

than naturally-occurring rocks. Table 2.10 presents the effect of temperature on the coefficient of thermal expansion 

values and Figure 2.241 presents examples of linear thermal expansion of various rocks with temperature.

Table 2.9 Coefficients of thermal expansion of different rocks and concrete at normal temperature

Rock Group

Typical silica 

content by wt. 

%

Coefficient of thermal expansion (10-6 ̊C-1)

Rock Concrete

Range Average Range Average

Chert

Quartzite

Quartz

Sandstone

Marble

Siliceous limestone

Granite

Dolerite

Basalt

Limestone

Gravel

94

94

84

Negligible

45

66

50

51

Negligible

5-95

7.4-13.0

7.0-13.2

4.3-12.1

2.2-16.0

3.6-9.7

1.8-11.9

4.5-8.5

4.0-9.7

1.8-11.7

11.8

10.3

9.3

8.3

8.3

6.8

6.8

6.4

5.5

11.4-12.2

11.7-14.6

9.0-13.2

9.2-13.3

4.1-17.4

8.1-11.0

8.1-10.3

-

7.9-10.4

4.3-10.3

9.0-13.7

-

12.1

-

11.4

10.7

10.7

9.6

9.6

9.3

8.6

Source: R.D. Browne, ―Properties of Concrete in Reactor Vessels,‖ Proceedings of the Conference on Prestressed 

Concrete Pressure Vessels, Group C, Paper 13, pp. 131-151, Institute of Civil Engineers, London, United 

Kingdom, 1967.

Table 2.10 Effect of elevated temperatures on the coefficient of thermal expansion on selected rocks

Temperature range 

( ̊C)

Coefficient of thermal expansion (10-6 ̊C-1)

Sandstone Limestone Granite Anorthsite

20-100

100-300

300-500

500-700

10.0

15.0

21.5

25.0

3.0

9.0

17.0

33.0

4.0

13.5

26.0

47.5

4.0

8.5

10.0

12.5

Source:

The resultant coefficient of thermal expansion for concrete is dominated by the aggregate material, however, the 

resulting coefficient tends to be somewhat higher due to the larger thermal expansion coefficient of the hardened 

cement paste, and will increase in proportion to the cement content [2.231]. It has been indicated that up to 700 ̊C 

the concrete expansion is very similar to that for the aggregate used in the concrete [2.54]. In theory, the coefficient 

of thermal expansion of concrete can be calculated based on knowledge of the relative volumetric proportions, the 

relative elasticities, and the thermal expansion coefficients of its constituents [2.136]. Figure 2.242 presents thermal 

coefficient of expansion values for neat cements, mortars, and concretes [2.232]. As shown, values of the 

coefficient for concretes range from ~2.2 10–6 to 3.9 10–6 per C with 3.1 10–6 per C being a typical value. 

The coefficient of thermal expansion of hardened cement paste and concrete tend to initially increase slightly with 

age due to a decrease in internal relative humidity and then decrease slowly [2.233]. The coefficient is influenced 

by the moisture condition (applies to paste component) and has minimum values for the two extremes: dry and 

saturated [2.227].
178

Figure 2.241 Linear thermal expansion of various rocks with temperature.

Source:

Figure 2.242 Coefficients of thermal expansion of neat cements, mortars, and concretes.

Source: Concrete Manual, 7th Ed., U.S. Bureau of Reclamation, Denver, Colorado, 1963.
179

Figure 2.243 presents a comparison of thermal strain results for Portland cement paste, mortars, and concretes for 

temperatures to 871 ̊C [2.234]. The coefficient of linear expansion increases with increasing temperature due to the 

aggregate expansion dominating over contraction of the cement paste, however, the effects of specimen moisture 

Figure 2.243 Thermal strain of Portland cement, mortar ,and concrete on heating.

Source: C.R. Cruz and M. Gillen, ―Thermal Expansion of Portland Cement Paste, Mortar, and Concrete at High 

Temperature,‖ Fire and Materials 4(2), 1980.

condition at test initiation (i.e., the number of thermal cycles that have been applied to the specimen) also has to be 

taken into consideration in determining the net specimen length change with temperature [2.235]. Table 2.11 

presents coefficient of thermal expansion values at elevated temperatures for concretes made with different 

aggregates, including heavyweight aggregate materials. Results have been presented that indicate that lightweight 

aggregate concretes (e.g., pumice and expanded shale aggregates) may shrink at temperatures greater than 300 ̊C 

[2.18]. Tests of sealed concrete indicate that under thermal cycling (20 ̊ to 149 ̊ to 20 ̊C) permanent expansive set 

occurs after one cycle and the permanent expansion increases with increasing number of cycles, but at a decreasing 

rate [2.65]. Also the coefficient decreased as the number of cycles increased and specimens permitted to dry after 

initial heating showed less expansion during subsequent thermal cycles.

Table 2.11 Coefficients of thermal expansion at elevated temperature

for concretes made with different aggregates

Type of aggregate in 

concrete

Coefficient of thermal expansion (10-6 ̊C-1)

Below 300 ̊C 300 ̊-600 ̊C 600 ̊-800 ̊C

Granite

Serpentine

Limonite

Haematite

Steel shot

Iron and steel scrap

0.71

4.14

4.86

5.94

4.20

5.1

10.4

4.1

4.5

11.5

8.5

7.2

15.9

1.3

4.2

16.2

16.2

8.6

Source: A.N. Komarovskii, Design of Nuclear Plants, 2nd Edition, Atomizdat, Moscow, Chapter 7 (Translated from 

Russian by Israel Program for Scientific Translations, Jerusalem), 1968.
180

Additional thermal expansion coefficients for limestone and siliceous aggregate concretes are presented in 

Figs. 2.244 and 2.245, respectively. Results indicate an almost monotonic increase in thermal expansion 

Figure 2.244 Thermal expansion of limestone aggregate concretes.

Source: U. Schneider, Behaviour of Concrete at High Temperature, HEFT 337, Deutscher Ausschuss für 

Stahlbeton, Wilhelm Ernst & Sohn, Munich, Germany, 1982.

Figure 2.245 Thermal expansion of siliceous aggregate concretes.

Source: U. Schneider, Behaviour of Concrete at High Temperature, HEFT 337, Deutscher Ausschuss für 

Stahlbeton, Wilhelm Ernst & Sohn, Munich, Germany, 1982.
181

coefficient for the limestone concrete until decarbonation (CaCO3 CaO + CO2) leads to a decrease in the 

coefficient. The thermal expansion of the siliceous concrete is greater than that for the limestone concrete. Basalt, 

due to its fine crystalline structure, exhibits a lower expansion than the siliceous concrete. The presence of load 

reduces the thermal expansion as noted in Figure 2.246. 

Figure 2.246 Variation of thermal strain with temperature and load for Portland 

cement concretes: 0 = no load, 0.1 = load corresponding to 10% fc’.

Source: P.J.E. Sullivan, G.A. Khoury, and B.N Grainger, ―Strain Behavior Under Uniaxial Compression of Three 

Concretes During First Heating to 600 ̊C,‖ Proceedings of 6th International Conference on Structural 

Mechanics in Reactor Technology, Vol. G-H, Paper 1/3, Paris, France, 1981.

Examples of Relations for Concrete Thermal Expansion. Traditionally the free thermal strain is expressed as a 

linear function of temperature by employing a thermal expansion coefficient such as 

(2.75)

where is the thermal expansion coefficient and T0 is the initial temperature. For normal weight concrete with 

siliceous aggregate can be approximated by [2.72]:

. (2.76)

This same expression has also been recommended for high strength ( > 55MPa) siliceous and carbonate aggregate 

concretes [2.118].

Additional relationships for concrete thermal expansion at high-temperature of siliceous and carbonate aggregate 

high-strength concretes have been developed [2.236].
182

For siliceous aggregate concrete

for 0 ̊ < T < 450 ̊C (2.77)

for 450 ̊ < T < 650 ̊C (2.78)

for 650 ̊ < T < 1000 ̊C (2.79)

For carbonate aggregate concrete

for 0 ̊ < T < 450 ̊C (2.80)

for 450 ̊ < T < 920 ̊C (2.81)

for 920 ̊ < T < 1000 ̊C. (2.82)

Summary. Thermal expansion represents the volume change of a material due to temperature change and is 

important in that it can result in thermal stresses and structural movement that can lead to concrete cracking and 

spalling. Thermal expansion is a non-linear function of temperature. The main factors influencing the coefficient of 

thermal expansion of concrete are the aggregate type and coarse aggregate fraction. Moisture content, water-cement 

ratio, and type cement only affect thermal expansion at relatively low temperatures (e.g., T < 200 ̊C). The 

coefficient of thermal expansion decreases slightly with age. Although Portland cement paste experiences 

contraction at temperatures above 150 ̊ to 400 ̊C, naturally-occurring aggregates always expand with increasing 

temperature to produce a net expansion. When differences in thermal expansion between Portland cement paste and 

aggregate materials are large, microstresses can occur resulting in microcracking of the concrete. Considerable 

differences exist relative to expansion of aggregate materials with the largest expansion occurring for aggregates 

having the highest percentage of silica by weight (e.g., quartz expansion > limestone expansion). Processed 

lightweight aggregate materials have lower thermal expansion coefficients than naturally-occurring aggregate 

materials and can show residual contraction after heating. Under elevated temperature conditions there is no 

significant difference for concretes that are dry and those that are completely saturated. The presence of load reduces 

the thermal expansion with increasing temperature. The thermal expansion is not fully reversible on subsequent 

cooling due to irreversible chemical changes and changes in microstructure. At high temperatures (600°C to 800°C) 

most concretes no longer exhibit an expansion and in some cases contract.

2.2.2.3 Thermal Conductivity

Information and Data. Thermal conductivity gives the heat flux transmitted through a unit area of a material under a 

unit temperature gradient (i.e., ability to conduct heat). For prestressed concrete pressure vessels, concrete with a 

high thermal conductivity is generally desirable so thermal gradients through the thickness will be minimal. At 

normal temperatures the thermal conductivity of concrete depends primarily on the thermal conductivity of the 

aggregate and the moisture content at the time of heating (e.g., increase in aggregate-to-cement ratio and decrease in 

water-cement ratio tends to increase the coefficient). Table 2.12 presents typical thermal conductivity values for 

concrete constituents and saturated concretes. Other factors influencing the thermal conductivity of concrete are the

Table 2.12 Thermal conductivities at ambient temperature of concrete constituents

Material Thermal conductivity (W/m ̊C)

Aggregates

Saturated concretes

Saturated hardened cement paste

Water

Air

0.7-4.2

1.0-3.6

1.1-1.6

0.515

.0034

Source: G.A. Khoury, Transient Thermal Creep of Nuclear Reactor Pressure Vessel Type Concretes, Ph.D. 

Dissertation, University of London, Vol. 1, pp. 1126; Vol. 2, pp. 418; Vol. 3, pp. 895, 1983.
183

hardened cement paste, and the pore volume and distribution. Age does not appear to affect the coefficient of 

thermal conductivity [2.136]. At low temperatures and with moist concrete very high values for thermal 

conductivity exist [2.15]. At higher temperatures the thermal conductivity increases slightly, but decreases as it 

approaches 100 C. Up to 300 C to 400 C, the thermal conductivity decreases further, and as the temperature 

increases beyond 300 C increasing cracking develops. Figure 2.247 presents the effect of temperature on thermal 

conductivity of an initially saturated concrete.

Figure 2.247 Effect of temperature on thermal conductivity of initially saturated concrete.

Source: R. Blundell, C. Diamond, and R. Browne, The Properties of Concrete Subjected to Elevated Temperatures, 

Report No. 9, Construction Industry Research and Information Association, Underwater Engineering 

Group, London, United Kingdom, June 1976.

Although thermal conductivity of concrete depends on all its constituents, it is largely determined by the type of 

aggregate used since aggregate makes up 60 to 80% the concrete volume. Table 2.13 lists thermal conductivity 

values for concrete with different aggregate types. Lightweight aggregates have lower thermal conductivities due to 

their porosity. Because the conductivity of water is approximately half that of cement paste, the lower the mixwater content, the higher the conductivity of the hardened concrete [2.24]. As shown in Figure 2.248, thermal 

conductivities of concrete made with highly crystalline aggregate decrease with temperature up to 1000 C, while 

those of concrete made with amorphous aggregate are essentially constant [2.18]. Figure 2.249 presents thermal 

conductivities for normal concretes having different aggregate types. Figures 2.250 and 2.251 presents thermal 

conductivities of concrete as a function of density and moisture content and the variation of thermal conductivity

Table 2.13 Typical values of thermal conductivity

Aggregate Type

Thermal conductivity

Btu in./hr•ft2•F W/m•K

Quartzite

Dolomite

Limestone

Granite

Rhyolite

Basalt

24

22

18-23

18-19

15

13-15

3.5

3.2

2.6-3.3

2.6-2.7

2.2

1.9-2.2

Source: P.K. Mehta and P.J.M. Monteiro, Concrete – Microstructure, Properties, and Materials, McGraw Hill, 

New York, New York, 2006.
184

as a function of temperature for several mortars and concretes, respectively. The thermal conductivity on cooling 

depends on the temperature level at the start of cooling, and after cooling to ambient conductivity falls to a level still 

below the original unheated level [2.19].

Figure 2.248 Thermal conductivity of normal and lightweight concretes.

Source: T. Z. Harmathy and L.W. Allen, ―Thermal Properties of Selected Masonry Unit Concretes,‖ 

J. American Concrete Institute 70, pp. 132–142, 1973.

Figure 2.249 Thermal conductivity of ordinary concretes with different aggregates.

Source: U. Schneider, C. Diererichs, and C. Ehm, ―Effect of Temperature on Steel and Concrete for PCRV’s,‖ 

Nuclear Engineering and Design 67, 245–258, 1981.
185

Figure 2.250 Thermal conductivity as a function of density and moisture content.

Source: K-Y. Shin, S-B. Kim, J-H. Kim, M. Chung, and P-S. Jung, ―Thermo-Physical Properties and Transient 

Heat Transfer of Concrete at Elevated Temperatures,‖ Nuclear Engineering and Design 212, pp. 233–241, 

2002.

Figure 2.251 Thermal conductivity as a function of temperature.

Source: K-Y. Shin, S-B. Kim, J-H. Kim, M. Chung, and P-S. Jung, ―Thermo-Physical Properties and Transient 

Heat Transfer of Concrete at Elevated Temperatures,‖ Nuclear Engineering and Design 212, pp. 233–241, 

2002.
186

Figure 2.252 presents the influence of temperature on thermal conductivity of a siliceous aggregate concrete from 

the French Penly Nuclear Power Plant [2.237]. Results presented in the figure are based on 40 measurements at 

each of the test temperatures (30 ̊, 60 ̊, 90 ̊, 150 ̊, and 200 C) and are comparable to those obtained by other 

researchers for siliceous aggregates in this temperature range [2.24,2.238].

Figure 2.252 Thermal conductivity of a nuclear power plant concrete.

Source: F. Vodak, R. Cerny, J. Drchalová, S. Hosková, O. Kapicková, O. Michalko, P. Semerák, and J. Toman, 

―Thermophysical Properties of Concrete,‖ Cement and Concrete Research 27(3), pp. 415-426,1997.

Examples of Relations for Concrete Thermal Conductivity. Constitutive relationships for concrete thermal 

conductivity at high-temperature have been developed. The American Society of Civil Engineers has proposed the 

following relationships for normal-strength concrete [2.87]:

Siliceous aggregate concrete

20 ̊C < T < 800 ̊C (2.83)

800 ̊C < T (2.84)

Carbonate aggregate concrete

20 ̊C < T < 293 ̊C (2.85)

293 ̊C < T (2.86)

This relationship has been modified for high-strength concrete [2.88]: 

Siliceous aggregate concrete

20 ̊C < T < 1000 ̊C (2.87)

Carbonate aggregate concrete

20 ̊C < T < 300 ̊C (2.88)

300 ̊C < T (2.89)
187

Finally, Eurocode has developed relationships for both normal- and high-strength concretes [2.46]:

Upper limit all concrete types

20 ̊C < T < 1200 ̊C (2.90)

Lower limit all concrete types

20 ̊C < T < 1200 ̊C (2.91)

Figure 2.253 provides a comparison of thermal conductivity values predicted by ASCE and Eurocode models and 

test data for siliceous and carbonate aggregate normal-strength concretes [2.88]. Experimental results indicate large 

variation in thermal conductivity values for normal-strength concretes. The standard deviation of the compiled data 

ranged from 0.22 W/m• ̊C at room temperature to 0.18 W/m• ̊C at 800 ̊C for siliceous aggregate concretes, and from 

0.4 W/m• ̊C at room temperature to 0.16 W/m• ̊C at 800 ̊C for carbonate aggregate concretes. Variation in test data 

was attributed to differences in test methods, conditioning procedures, and measurement techniques.

Figure 2.253 Comparison of experimental results and thermal conductivity 

predicted by different models for normal-strength concretes.

Source: V.K.R. Kodur, M.M.S. Dwaikat, and M.B. Dwaikat, ―High-Temperature Properties of Concrete for Fire 

resistance Modeling of Structures,‖ ACI Materials Journal 105(5), pp. 517-527, American Concrete 

Institute, Farmington Hills, Michigan, September-October 2008.

Summary. Major factors that influence the concrete thermal conductivity are moisture content, hardened cement 

paste, pore volume and distribution, and amount and type of aggregate material. The conductivity varies linearly 

with moisture content. Increasing the water-cement ratio increases the concrete porosity and correspondingly 

reduces the thermal conductivity, especially when dry, since both water and air have lower thermal conductivity 

values than cement paste. Type of cement can influence thermal conductivity relative to temperature at which water 

is released (e.g., occurs at higher temperature for blast furnace slag than ordinary Portland cement concretes). 

Leaner normal weight concrete mixes have higher thermal conductivities, while the opposite is true for lightweight 

concrete mixes. Although the thermal conductivity of concrete depends on all its constituents, the concrete thermal 

conductivity is largely determined by the amount and type of aggregate used since aggregates normally constitute 60 
188

– 80% the concrete volume. At normal temperatures the thermal conductivity increases with increasing aggregate 

content and moisture content. Crystalline aggregates (e.g., quartz) have higher thermal conductivities at normal 

temperatures with values decreasing as the temperature increases. Amorphous rocks (e.g., basalt and dolerite) have 

low thermal conductivity values at normal temperatures. At normal temperatures moist concretes have relatively 

high thermal conductivity values. Pre-dried siliceous and calcareous aggregate concretes exhibit an approximately 

linear reduction in thermal conductivity with temperature when heated from 20 ̊ to 500 ̊C. In general, at 

temperatures up to 100 ̊C the thermal conductivity seems to increase with temperature; at temperatures greater than 

100 ̊C the conductivity decreases and at 300 ̊ to 400 ̊C there is a further decrease; at temperatures > 600 ̊C the 

thermal conductivity increases slightly: and at temperatures greater than 800 ̊C up to melting there is only a small 

change in thermal conductivity. Lightweight aggregate concretes have even lower thermal conductivity values at 

normal temperatures due to increased porosity. For lightweight aggregate concrete the thermal conductivity is 

relatively constant or increases slightly for temperatures up to 1000 ̊C.

2.2.2.4 Thermal Diffusivity

Information and Data. Thermal diffusivity is a measure of the rate at which heat will diffuse through a material in 

all directions due to a temperature change and is thus an index of the facility with which the material will transfer 

heat due to a temperature change. It is described as follows:

D = / cp (m2/sec) (2.92)

where D = diffusivity, = density (kg/m3), = thermal conductivity (W/mK), and cp = specific heat (J/kgK).

Thermal diffusivity is important to nuclear power plant structures such as prestressed concrete pressure vessels for 

the same reasons cited for thermal conductivity. Thermal diffusivity of concrete is determined by the thermal 

properties of its constituents. At normal temperatures the diffusivity of concrete is mainly governed by the 

diffusivity of the aggregate. Aggregates with increasing values of thermal diffusivity include basalt, rhyolite, 

granite, limestone, dolerite, and quartzite [2.239]. Table 2.14 presents typical values of thermal diffusivity for 

concretes made with different types of coarse aggregate.

Table 2.14 Thermal diffusivity values for concrete with different coarse aggregate

Type coarse aggregate ft2/h m

2

/h

Quartzite

Dolomite

Limestone

Granite

Rhyolite

Basalt

0.058

0.051

0.050

0.043

0.035

0.032

0.0054

0.0047

0.0046

0.0040

0.0033

0.0030

Source: P.K. Mehta and P.J.M. Monteiro, Concrete – Microstructure, Properties, and Materials, McGraw Hill, 

New York, New York, 2006.

Factors that affect thermal conductivity generally have the same influence on thermal diffusivity. Thermal 

diffusivity of limestone and siliceous aggregate concretes is presented in Figures 2.254 and 2.255, respectively 

[2.12]. Figure 2.256 presents the influence of temperature on thermal diffusivity of a siliceous aggregate concrete 

from the French Penly Nuclear Power Plant [2.237] and is noted to be similar to other results presented in the 

literature for siliceous aggregate concretes in the temperature range investigated (30°C to 200°C) [2.240]. 

Figure 2.257 presents the thermal diffusivity of several mortars and concretes as a function of temperature [2.228].

The thermal diffusivity of normal weight concrete falls markedly with increase in temperature as noted in 

Figure 2.258. This is due to the general decrease of thermal conductivity and increase in specific heat at elevated 

temperature. Finally, Figure 2.259 presents thermal diffusivity results for three normal weight concretes and thirteen 

lightweight concretes. As noted in the figure, at about 600 ̊C the thermal diffusivity of all concretes investigated 

becomes approximately 0.0033 cm2/sec.
189

Figure 2.254 Thermal diffusivity of limestone aggregate concrete.

Source: U. Schneider, Behavior of Concrete at High Temperature, HEFT 337, Deutscher Ausschuss fur Stahlbeton, 

Wilhelm Ernst & Sohn, Berlin, Germany, 1982.

Figure 2.255 Thermal diffusivity of siliceous aggregate concrete.

Source: U. Schneider, Behaviour of Concrete at High Temperature, HEFT 337, Deutscher Ausschuss für 

Stahlbeton, Wilhelm Ernst & Sohn, Munich, Germany, 1982.
190

Figure 2.256 Thermal diffusivity of a nuclear power plant concrete.

Source: F. Vodak, R. Cerny, J. Drchalová, S. Hosková, O. Kapicková, O. Michalko, P. Semerák, and J. Toman, 

―Thermophysical Properties of Concrete,‖ Cement and Concrete Research 27(3), pp. 415-426,1997.

Figure 2.257 Thermal diffusivity variation of a Korean nuclear plant concrete as a function of temperature.

Source: K-Y. Shin, S-B. Kim, J-H. Kim, M. Chung, and P-S. Jung, ―Thermo-Physical Properties and Transient 

Heat Transfer of Concrete at Elevated Temperatures,‖ Nuclear Engineering and Design 212, pp. 233–241 

2002.
191

Figure 2.258 Effect of temperature on thermal diffusivity of normal weight concrete excluding latent heat effects.

Source:

Figure 2.259 Thermal diffusivity of normal and lightweight concretes.

Source: T. Z. Harmathy and L.W. Allen, ―Thermal Properties of Selected Masonry Unit Concretes,‖ 

J. American Concrete Institute 70, pp. 132–142, 1973.
192

2.2.2.5 Specific Heat

Information and Data. The thermal or heat capacity of a material, or specific heat, is the amount of heat per unit 

mass required to change the temperature of the material by one degree. 

(2.93)

where H is enthalpy, T is temperature, and p is pressure. The specific heat of ordinary concrete at room temperature 

ranges from 0.5 to 1.13 kJ kg-1 K

-1 while specific heat of hardened cement paste ranges from 0.63 to 1.72 kJ kg-1 K-1

 

[2.5]. At room temperature aggregate type, mix proportions, and age do not have a great effect on specific heat of 

concrete. However, as the moisture content in the concrete increases, the specific heat capacity increases at lower 

temperatures. The specific heat of aggregate can be calculated from its mineralogical content and the specific heat 

of concrete determined based on its relative proportions [2.136].

The specific heat of concrete increases with an increase in temperature. At elevated temperature the specific heat 

value is sensitive to the various transformations that take place in concrete [2.19]. This includes the vaporization of 

free water at about 100 ̊C, the dissociation of Ca(OH)2 at about 400 ̊ to 500 ̊C and the quartz transformation in 

some aggregates. Heating of initially saturated concrete causes a rapid but temporary rise in the specific heat at 

about 90 ̊C due to the rapid release of latent heat of vaporization. 

Figure 2.260 presents the effects of temperature on experimentally determined specific heats of various concretes. 

Figures 2.261 and 2.262 present heat capacities of limestone and siliceous aggregate concretes, respectively. 

Additional information is provided in Figure 2.263, which presents the influence of temperature on the specific heat 

of a siliceous aggregate concrete from the French Penly Nuclear Power Plant . Figure 2.264 presents specific heat as 

a function of temperature for a basaltic aggregate concrete (37.9 MPa, Type V cement) used in construction of 

Yonggwang Nuclear power plant units 3 and 4 in Korea [2.241]. Results show that the specific heat increased to 

500 ̊C, decreased from 700 ̊ to 900 ̊C, and then increased at temperatures above 900 ̊C. Finally, Figure 2.265 

presents specific heat results for three normal weight concretes and thirteen lightweight concretes. It was noted in 

the reference for this data that generally the specific heat of concrete was somewhat insensitive to the aggregate used 

and mix ratios [2.18].

Figure 2.260 Specific heats of various concretes.

Source:
193

Figure 2.261 Specific heat capacity of limestone aggregate concrete.

Source: U. Schneider, Behaviour of Concrete at High Temperature, HEFT 337, Deutscher Ausschuss für Stahlbeton, 

Wilhelm Ernst & Sohn, Munich, Germany, 1982.

Figure 2.262 Specific heat capacity of siliceous aggregate concrete.

Source: U. Schneider, Behaviour of Concrete at High Temperature, HEFT 337, Deutscher Ausschuss für Stahlbeton, 

Wilhelm Ernst & Sohn, Munich, Germany, 1982.
194

Figure 2.263 Specific heat of a nuclear power plant concrete. 

Source: F. Vodak, R. Cerny, J. Drchalová, S. Hosková, O. Kapicková, O. Michalko, P. Semerák, and J. Toman, 

―Thermophysical Properties of Concrete,‖ Cement and Concrete Research 27(3), pp. 415-426,1997.

Figure 2.264 Specific heat of a Korean nuclear power plant concrete.

Source: K-Y. Shin, S-B. Kim, J-H. Kim, M. Chung, and P-S. Jung, ―Thermo-Physical Properties and Transient 

Heat Transfer of Concrete at Elevated Temperatures,‖ Nuclear Engineering and Design 212, pp. 233–241 

2002.
195

Figure 2.265 Specific heat of normal and lightweight concretes.

Source:

Examples of Relationships for Concrete Thermal Capacity. Constitutive relationships for concrete thermal 

capacity at high-temperature have been developed. The American Society of Civil Engineers has proposed the 

following relationships for normal-strength concrete [2.87]:

Siliceous aggregate concrete

20 ̊C < T < 200 ̊C (2.94)

200 ̊C < T < 400 ̊C (2.95)

400 ̊C < T < 500 ̊C (2.96)

500 ̊C < T < 600 ̊C (2.97)

600 ̊C < T (2.98)

Carbonate aggregate concrete

20 ̊C < T < 400 ̊C (2.99)

400 ̊C < T < 410 ̊C (2.100)

410 ̊C < T < 445 ̊C (2.101)

445 ̊C < T < 500 ̊C (2.102)

500 ̊C < T < 635 ̊C (2.103)

635 ̊C < T < 715 ̊C (2.104)

715 ̊C < T < 785 ̊C (2.105)

785 ̊C < T (2.106)
196

This relationship has been modified for high-strength concrete [2.88]: 

Siliceous aggregate concrete

20 ̊C < T < 200 ̊C (2.107)

200 ̊C < T < 400 ̊C (2.108)

400 ̊C < T < 500 ̊C (2.109)

500 ̊C < T < 600 ̊C (2.110)

600 ̊C < T < 635 ̊C (2.111)

Carbonate aggregate concrete

20 ̊C < T < 400 ̊C (2.112)

400 ̊C < T < 475 ̊C (2.113)

475 ̊C < T < 650 ̊C (2.114)

650 ̊C < T < 735 ̊C (2.115)

735 ̊C < T < 800 ̊C (2.116)

800 ̊C < T < 1000 ̊C (2.117)

Finally, Eurocode has developed relationships for both normal- and high-strength concretes [2.46]:

Specific heat (J/kg ̊C)

20 ̊C < T < 100 ̊C (2.118)

100 ̊C < T < 200 ̊C (2.119)

200 ̊C < T < 400 ̊C (2.120)

400 ̊C < T < 1200 ̊C (2.121)

Density change (kg/m3)

= reference density 20 ̊C < T < 115 ̊C (2.122)

115 ̊C < T < 200 ̊C (2.123)

200 ̊C < T < 400 ̊C (2.124)

400 ̊C < T < 1200 ̊C (2.125)

Thermal capacity = x c. (2.126)

Figure 2.266 provides a comparison of specific heat values predicted by ASCE and Eurocode models and test data 

for siliceous and carbonate aggregate normal-strength concretes. The standard deviation of the compiled data 

ranged from 0.56 MJ/m3• ̊C at room temperature to 1.1 MJ/m3• ̊C at 750 ̊C for siliceous aggregate concretes, and 

from 0.6 MJ/m3• ̊C at room temperature to 7.8 MJ/m3• ̊C at 750 ̊C for carbonate aggregate concretes. Results 

indicate the significant influence that aggregate type has on specific heat. High heat capacity of carbonate aggregate 

concrete in temperature range of 600 ̊ to 800 ̊C was due to endothermic reaction that absorbs a significant amount of 

energy. The heat capacity of carbonate aggregate concrete helps to minimize spalling.
197

Figure 2.266 Comparison of experimental results and specific heat 

predicted by different models for normal-strength concretes.

Source: V.K.R. Kodur, M.M.S. Dwaikat, and M.B. Dwaikat, ―High-Temperature Properties of Concrete for Fire 

resistance Modeling of Structures,‖ ACI Materials Journal 105(5), pp. 517-527, American Concrete 

Institute, Farmington Hills, Michigan, September-October 2008.

Summary. Aggregate type, mix proportions, and age do not have a great effect on specific heat at normal 

temperatures. Type of aggregate and its proportion have little influence on heat capacity at temperatures below 

800 ̊C. Mix proportions influence heat capacity in that richer mixes indicate a higher latent heat due to dehydration 

effects. At normal temperatures as the moisture content in the concrete increases the specific heat capacity increases 

(e.g., at temperatures below 200 ̊C wet concretes show an apparent specific heat nearly twice that of oven-dried 

concretes). At elevated temperatures the specific heat value is sensitive to the various transformations that take 

place in concrete (e.g., vaporization of free water at about 100 ̊C, the dissociation of Ca(OH)2 at about 400 ̊ to 500 ̊C 

and the quartz transformation in some aggregates). Specific heat of siliceous and limestone aggregate concretes 

increase as temperature increases up to about 700 ̊C, but the increase for granite aggregate concrete is not as 

significant.

2.2.2.6 Heat of Ablation and Erosion Rates

Information and Data. Under a hypothetical core disruptive accident involving melting of the reactor core in a 

nuclear power plant it is of interest to know how much energy is dissipated by the concrete structures designed to 

contain the core melt. The term heat of ablation is applied in this context and is defined as the heat dissipated per 

unit mass material during a steady-state erosion process resulting in the removal of a unit mass [2.15]. The heat of 

ablation is composed of the sensible heat to be provided and the heats of reaction and transformation. It has been 

calculated that the total amount of heat required to transform a cubic meter of quartzite concrete at normal 

temperature to a molten state is about 2000 x 103kJ and for limestone concrete the amount of heat is greater (2360 x 

103kJ) due to additional heat required to decarbonate limestone aggregate [2.5]. Table 2.15 presents heat of 

ablation values for different concretes. Note that values in the table are dependent on test conditions and the type of 

determination.
198

Table 2.15 Heat of ablation values of different concretes

Type of concrete Heat of ablation (MJ/kg) Source with references identified in [2.15]

Basaltic limestone

Basaltic limestone

Quartzitic

Calcitic

Quartzitic

Calcitic

6±3

4.5±0.5

2.2

1.5

2.4

3.2

Muir (6)

Chu (7)

Hildenbrand et al. (8)

Hildenbrand et al. (8)

Schneider et al. (5)

Schneider et al. (5)

Source:

Under a hypothetical core disruptive accident involving melting of the reactor core in a nuclear power plant it is 

possible that liquid metals such as steel having temperatures up to 3000 ̊C may penetrate and damage concrete 

structures. This is the erosion rate and refers to the velocity at which concrete subjected to high temperatures (i.e., 

>1200 ̊C) is decomposed by mechanical disintegration and by melting process [2.15]. Figure 2.267 presents 

penetration depth versus time results for simulated molten core (thermite) concrete interactions for a basaltic 

aggregate concrete (37.9 MPa, Type V cement) used in construction of Yonggwang Nuclear power plant Units 3 and 

4 in Korea [2.241]. The peak melt temperature measured was 2230 ̊C. Figure 2.268 presents the cross section of a 

concrete specimen after testing. Post-test examination of the cross section showed a 2-3 mm thick crevice between 

the solidified melt and the erosion front with debris generated during the interaction observed in the solidified melt. 

The erosion front was irregular and bow-shaped with a maximum depth of 2 cm. Analysis indicated that the 

maximum concrete erosion rate was 175 cm/h. 

Figure 2.267 Penetration depths vs time for various melt temperature conditions. 

Source: K-Y. Shin, S-B. Kim, J-H. Kim, M. Chung, and P-S. Jung, ―Thermo-Physical Properties and Transient 

Heat Transfer of Concrete at Elevated Temperatures,‖ Nuclear Engineering and Design 212, pp. 233–241, 

2002.

Table 2.16 presents calculated heats of transformation and decomposition reactions associated with physical and 

chemical changes that occur in a cubic meter of concrete made with quartzitic and limestone aggregates and 

Portland blast furnace slag cement [2.242]. Mix proportions by weight for the cement, fine aggregate, coarse 
199

Figure 2.268 Cross section of test specimen after MEK-T1A test.

Source: K-Y. Shin, S-B. Kim, J-H. Kim, M. Chung, and P-S. Jung, ―Thermo-Physical Properties and Transient 

Heat Transfer of Concrete at Elevated Temperatures,‖ Nucl. Eng. and Design 212, pp. 233–241 2002.

Table 2.16 Estimated heats of transformation for quartzitic and 

limestone aggregate concretes exposed to elevated temperatures

Source: U. Schneider and U. Diederichs, ―Physical Properties of Concrete From 20 ̊C Up To Melting,‖ Parts 1 and 

2, Betonwerk & Fertigteiltechnik, Heft 3, pp. 141-150, and Heft 4, pp. 223-230, 1981.
200

aggregate and water were approximately 1:1.9:3.5:0.5, respectively. Calculations indicate that the total amount of 

heat required to transform a cubic meter of quartzitic aggregate concrete at normal temperature to a molten state is 

approximately 2000 x 103kJ. For limestone concrete the amount can be nearly twice as large due to additional heat 

required to decarbonate the aggregate.

Table 2.17 presents average rates of erosion by liquid steel for several concretes. Erosion rates in the table were 

determined either by immersing the specimens in molten steel produced by an induction furnace (1600 ̊C) or by 

metallographic reaction (2600 ̊C) in which thermite was placed into concrete crucibles and ignited. Results indicate 

that basaltic concrete gave lowest erosion rates, followed by limestone concrete and quartzitic concrete. Variations 

in data were the result of different concrete mixes and test conditions. From these results it was concluded that 

concrete could resist very high thermal loadings arising from its interaction with liquid steel, however, very thick 

concrete foundations would be required to resist a hypothetical core disruptive accident involving melting of the 

core [2.243]. 

Table 2.17 Erosion of different concretes by liquid steel

Type of aggregate in 

concrete

Average rate of erosion

(mm/min)

Test method Source with references 

identified in [2.15]

Limestone

Quartzite

Basalt

Limestone

Quartzite

Basaltic limestone

Quartzite

Basalt

Limestone

Basaltic limestone

66

44

10

20

30

20

40

22

35

12

Immersed in molten steel

(about 1600 ̊C)

Metallographic reaction 

(about 2600 ̊C)

Metallographic reaction 

Metallographic reaction 

Molten pool

Arc-heating

Plasma jet

Ehm et al. (11)

Ehm et al. (11)

Sutherland (9)

Perinic et al. (10)

Hildenbrand (8)

Muir (6)

Source:

Summary. Under a hypothetical core disruptive accident involving melting of the reactor core in a nuclear power 

plant it is of interest to know how much energy is dissipated by the concrete structures designed to contain the core 

melt. The term heat of ablation is applied in this context and is defined as the heat dissipated per unit mass material 

during a steady-state erosion process resulting in the removal of a unit mass. For a heat flow between 20 and 

200 W•cm2, the heat of ablation of concrete has been indicated to be 6000 (±3000) kJ•kg-1[2.40]. Others estimates 

of the heat of ablation are 2227 and 2400 kJ•kg-1for concrete made with siliceous aggregates, 1474 kJ•kg-1for 

concrete made from calcareous aggregates, and 3200 kJ•kg-1for concretes made from calcareous aggregates 

[2.242,2.244]. The amount of heat required to transform a given volume of concrete to a molten state is about twice 

as much for limestone aggregate concrete as for quartzitic concrete due to the additional heat required to 

decarbonate the limestone aggregate. The rate at which concrete subjected to high temperatures (e.g., > 1200 ̊C) is 

decomposed by mechanical disintegration and the melting process is the erosion rate. Experimental results for 

erosion rates of limestone, quartzitic, and basalt aggregate concretes indicated erosion rates of 20 to 66, 30 to 44, 

and 10 to 22 mm/min, respectively. Results for a nuclear power plant basaltic concrete mix indicated a rate of about 

30 mm/min. The erosion rate does not appear to be that dependent on the type of aggregate used in the concrete 

mix. Scatter of data presented in the literature is the result of different concrete mixes and test conditions. 
201

2.2.2.7 Moisture Diffusion and Pore Pressure

Information and Data. Movement of moisture in concrete can result by drying from an exposed surface or from 

temperature gradients. Prediction of moisture migration and pore pressure build up in non-uniformly heated 

concrete is important for safe operation of concrete containment vessels in nuclear power plants and for assessing 

the behavior of fire-exposed concrete structures [2.245]. Two main processes are at work when a heat flux is 

applied to a concrete element: thermo-mechanical and thermal-hydral. The thermo-mechanical process is associated 

with the temperature field in the concrete element and induces thermal gradients that result in thermal dilation 

generating tensile stresses perpendicular to the heated face. The thermal-hydral process, illustrated in Figure 2.269, 

is associated with transfer of mass (water in liquid and vapor phases and air) [2.246]. As the temperature increases

water is partly evaporated to generate a pressure in the porous network. The pressure gradient is the major driving 

force for mass transfer. Vapor and air are partially evacuated to the heated surface but they also migrate to the center 

of the element where the vapor can condense forming a quasi-saturated layer. After some time the concrete element 

forms from the heated surface a dry and dehydrated zone, a drying and dehydrated zone, and a quasi-saturated layer 

that can act as an impermeable wall for gases. Under the right conditions (e.g., low permeability) these processes 

can result in spalling of the concrete. Transfer of free moisture in concrete can affect the concrete strength, thermal 

expansion, shrinkage, and creep. Changes in moisture content distribution also can affect the durability and 

radiation shielding capability of concrete. Pressure build up under postulated accident conditions can potentially 

lead to failure of the metallic liner.

Figure 2.269 Illustration of process of pressure build up in 

a concrete element under the effect of a pressure gradient.

Source: P. Kalifa, F-D. Menneteau, and D. Quenard, ―Spalling and Pore Pressure in HPC at High Temperatures, 

Cement and Concrete Research 30, pp. 1915-1927, 2000.

At moderate temperatures transfer of moisture in cement paste as well as uncracked concrete is slow (i.e., diffusivity 

of concrete is approximately 10 mm2day-1) [2.247]. Measurements on a thick-wall section (1.52 m) heated on the 

front face and sealed on all faces except the back face (unheated face) and subjected to a 3.9 ̊C/m temperature 

crossfall indicated that even after one year only the inner 0.3 m had lost free water, but shrinkage was large [2.248]. 

As the temperature increases permeability and diffusivity increase and approximately obey the activation energy 

concept up to about 100 ̊C [2.5]. The permeability of concrete depends on the viscosity of water, and to a smaller 

extent on the density of water (i.e., viscosity drops from 1.00 MPa s at 20 ̊C to 0.35MPa s at 80 ̊C, while the density 
202

is reduced from 1.00 to 0.97 kg/m3) [2.249]. Results obtained by investigating a number of different concretes (44 

to 190 MPa compressive strength) indicate that the concrete permeability increases from 13 to 62% when the 

temperature is raised from 20 ̊ to 50 ̊C and by 3 to 55% by an additional increase to 80 ̊C, with considerable 

variation in results depending on the type of concrete, and that the increase in permeability and diffusivity can be 

predicted from theory of thermodynamics [2.249]. Above 100 ̊C the drying times for concrete become orders of 

magnitude shorter (i.e., the moisture diffusivity increases significantly as 100 ̊C is exceeded as the average pore size 

increases significantly and flow of water probably takes place mainly in a vaporized state as steam) [2.250]. As the 

temperature increases further the moisture diffusivity continues to increase [2.5]. 

Heating concrete can produce very large pore pressures that are functions of temperature, temperature history, and 

size of specimen. The transformation and mass transport of moisture in concrete have a unique influence on 

concrete’s temperature profile and pore pressure development. A general description of the evolution of pore 

pressure build up in heated concrete is provided in results for a study conducted to quantify the effects of factors 

influencing pore pressure build up and the potential for explosive spalling of normal- and high-strength concretes 

[2.251]. A noticeable rise in pore pressure occurred when the concrete reached the temperature range of 105 ̊C to 

160 ̊C that coincided with the vaporization of free water and the transport of water vapor as evidenced by a drop in 

the concrete’s rate of temperature rise. In the temperature range of 160 ̊C to 180 ̊C there was a sharp rise in pore 

pressure that coincided with the release of chemically-bound water. Beyond 180 ̊C the concrete pore pressure 

continued to rise, but with a decreasing rate until peaking occurred in the temperature range from 220 ̊C to 245 ̊C. 

At this point the concrete either spalled explosively or the pore pressure attenuated. General conclusions from this 

study were that specimens with high water-cementitious materials ratios and lower initial moisture contents have 

lower pore pressures than their counterparts and that thermal gradients were an influencing factor in the process of 

pore pressure build up (e.g., higher thermal gradients caused by high heating rates can lead to microcracks that 

permit moisture to escape resulting in lower pore pressures).

A number of experimental studies were conducted several years ago investigating pore pressures and moisture 

distributions under thermal gradients in support of development of gas-cooled reactor technology [2.252]. Pore 

pressure was investigated with regard to concrete spalling and the build up of backpressure behind the liner of a 

nuclear power plant concrete containment due to thermal gradients. The backpressures between the concrete and 

liner result from steam and saturated vapor pressures that develop from the free water in the concrete. Additional 

pressure also develops under elevated temperature from heating of the non-condensable gases that were trapped in 

the concrete pores during the concrete curing process. A consistent finding of the experimental studies was that the 

generated pore pressure under elevated temperature was about equal to the sum of the saturated vapor pressure and 

the partial air pressure for the entrapped pore gases at the temperature of the concrete [2.252]. These results indicate 

that the pore pressure for a sealed specimen can exceed 414 kN/m2(0.414 MPa) at 140 ̊C as noted in Figure 2.270.

In unsealed specimens, diffusion and eventual drying leads to lower pore pressures than those obtained for sealed 

specimens at the same temperature. Additional data are provided in Figure 2.271 which presents temperature and 

pore pressure distributions in a 30.5-cm-thick wall at 4 and 7 hours after heat-up in a test program to investigate 

water-release rates in walls [2.254].

Experiments were performed on thick sections of concrete to determine long-term shrinkage, pore pressures, and 

moisture distributions [2.255]. Pore pressure measurements were made using cylindrical concrete specimens that 

were 0.6-m-long and either sealed completely or had a pressure release to the atmosphere at the low temperature 

end. Low and high temperatures for the vented and sealed specimens were 40 ̊ and 130 ̊C and 40 ̊ and 140 ̊C, 

respectively. Figure 2.272 presents pressure distributions for the vented and sealed specimen tests at different 

exposure times. Results indicate that early in the test the behavior of the two specimen types is similar in that 

differences in sealing become apparent only after 40 days exposure. In both cases the pore pressures dissipate 

slowly in the hot regions as water migrates toward the cooler parts of the specimen. Additional tests of moisture 

distribution under thermal gradients were conducted using cylindrical concrete specimens having lengths 

representing wall thicknesses of 0.6, 0.9, 1.6, or 3.1 m. Each specimen was heated to the same nominal base 

temperature of 125 ̊C with the concrete exposed to atmospheric conditions at the cold end. The cylinders were

 

Results of an analytical study for the Economic Simplified Boiling Water Reactor indicate that a backpressure of 

71.0 kPa can develop at the end of a 72-hour period following a loss-of-coolant accident [2.253].
203

Figure 2.270 Pore pressure build up in sealed concrete specimen.

Source: G.L. England and T.J. Sharp, ―Migration of Moisture and Pore Pressures in Heated Concrete,‖ Proceedings 

of 1st International Conference on Structural Mechanics in Reactor Technology, Elsevier Science 

Publishers, North-Holland, The Netherlands, 1971.

Figure 2.271 Temperature distribution and pore pressure data for 30.5-cm-thick heated concrete wall.

Source: K.H. Chen, E.L. Glueker, E.L. Lam, and V.S. Shippey, ―Comparison of Mechanical Codes for Predicting 

Water Release from Heated Concrete,‖ GEFR-00521, General Electric, April 1980.
204

Figure 2.272 Pressure distribution for vented and sealed pore pressure test at given times.

Source: G.L. England and A.D. Ross, ―Shrinkage, Moisture, and Pore Pressure in Heated Concrete,‖ Paper SP 34-

42 in Concrete for Nuclear Reactors, pp. 883-907, American Concrete Institute, Farmington Hills, 

Michigan, 1972.

sealed along their lengths. Figure 2.273 presents phase diagrams for water in the concrete specimens after 887 days. 

As shown in the figure, drying has occurred at the two ends of the specimen, and depending on the length of the 

moisture path, an excess of water may be present in intermediate regions. In other words, for the longer specimens

Figure 2.273 Phase diagrams for water in concrete at 887 days for concretes

of various lengths and hot face temperature of 125 ̊C.

Source: G.L. England and A.D. Ross, ―Shrinkage, Moisture, and Pore Pressure in Heated Concrete,‖ Paper SP 34-

42 in Concrete for Nuclear Reactors, pp. 883-907, American Concrete Institute, Farmington Hills, 

Michigan, 1972.
205

the temperature gradients have caused more water to migrate into the intermediate regions than to migrate from 

them to produce a higher than normal water content. Drying depths obtained after 887 days heating for 3.05-m-long 

specimens that experienced hot-face temperatures of 150 ̊, 125 ̊, or 100 ̊C were 0.914, 0.49, and 0.305 m, 

respectively. It was concluded from the study that pore pressures are unlikely to be of importance in the bulk of the 

concrete of a concrete reactor vessel but could become of importance relative to stability of the liner adjacent to the 

vessel wall. At temperatures below 100 ̊C results also indicated that drying is not likely to be a very important 

factor in thick sections because even after many years exposure to such a thermal crossfall drying is unlikely to 

penetrate more than 0.5 m from either face. As the hot-face temperature increases the depth to which drying 

penetrates at the cold face remains essentially unchanged while that at the hot face increases. Other research for a 

prestressed concrete pressure vessel having a 35 ̊C maximum temperature (i.e., Wylfa) indicates that no significant 

moisture migration should occur over 30 years except at the outer face [2.136]. 

More recently, thermally-induced pore pressures and corresponding temperatures in normal- and high-strength 

concretes were determined in a experimental study to quantify several factors influencing pore pressure build up and 

the potential for spalling [2.251]. Normal- and high-strength concretes having compressive strengths at 28-days age 

of 40.6 and 75.3 MPa, respectively, were used to fabricate concrete blocks 100 x 100 x 200 mm3. In addition to 

concrete strength, other variables included curing condition (submerged or air-dried), presence of polypropylene 

fibers (either 13 or 38-mm-long, and amounts of either 1.5 or 3 kg/m3), and heating rate (either 5 ̊C/min or 

25 ̊C/min). The concrete blocks contained pore pressure gages located at 25, 50, and 75 mm from the heated face. 

Prior to heating, all surfaces of the specimens, except the heated surface, were insulated with an insulating blanket to 

promote one-directional heat flow. Specimens were heated to 600 ̊C. Results of the study noted the influence of 

transformation and mass transport of moisture in concrete on its temperature profile and pore pressure development. 

There was a noticeable increase in pore pressure when the concrete reached the temperature range of 105 ̊ to 160 ̊C 

which coincides with the vaporization of free water and transport of water vapor as evidenced by a drop in the 

concrete’s rate of temperature rise. A sharp increase in pore pressure occurred in the concrete temperature range of 

160 ̊ and 180 ̊C which coincided with the release of chemically-combined water. At concrete temperatures greater 

than 180 ̊C the pore pressure continued to rise, but with a decreasing rate, until peaking at a concrete temperature of 

220 ̊ to 245 ̊C. After this either the pressure attenuated or explosive spalling occurred. Spalling only occurred for 

the high-strength concrete that had been cured by submerging and heating at 5 ̊C/min. Pore pressures measured 

when spalling occurred averaged 2.05 MPa. Addition of polypropylene fibers significantly reduced pore pressure 

build up with the pressure build up decreasing as fiber content increased. Normal-strength concrete (high watercement ratio and lower initial water content) consistently had lower pore pressures compared with the high-strength 

concretes cured by submerging. Thermal gradient was determined to be an influencing factor in the process of pore 

pressure build up with higher thermal gradients producing lower pore pressures, possibly due to formation of 

microcracks that permitted pore pressure to escape.

In an effort to obtain information regarding the mature of moisture movement and rate of moisture loss in a 

prestressed concrete pressure vessel (PCPV), an experimental study of moisture migration in a pie-shaped specimen 

representing the flow path through a cylindrical wall of the PCPV was conducted [2.256]. The model was 2.74 m in 

length with cross-sectional dimensions of 0.61 x 0.61 m2on one end by 0.81 x 0.81 m2on the other end. It was 

sealed against moisture loss on the small end (interior) and along lateral surfaces and exposed to the atmosphere at 

the other end (exterior). Temperature distributions, shrinkage, and moisture distribution were monitored for 

approximately 17 months prior to application of a 44 ̊C temperature gradient that was maintained for one year. A 

series of heating lamps was used to maintain the required temperature on the interior surface. At the end of the test, 

with the exception of zones nearest the specimen ends, moisture contents were relatively constant as noted in 

Figure 2.274 (i.e., moisture contents at two ends were approximately 15% less than the average moisture content for 

the other measuring locations). Concrete strains corrected for thermal effects were small with only about 1 m (that 

nearest open end) indicating shrinkage strain in excess of 20 millionths, implying that drying shrinkage was 

minimal. It was concluded that moisture migration in thick sections of concrete, such as a thick-walled prestressed 

concrete pressure vessel, is a slow process and is not likely to be a significant factor with temperature differences of 

44 ̊C or less.
206

Figure 2.274 Changes in moisture content in specimen after heating and

moisture content profile at end of test period.

Source: J. E. McDonald, Moisture Migration in Concrete, Technical Report C-75-l, U.S. Army Waterways 

Experiment Station, Vicksburg, Mississippi, May 1975.

A study has been conducted to investigate the effects on temperature distribution, moisture migration, and strain 

variation due to heating of a simulated section of a mass concrete wall [2.257]. Cube specimens 1500 mm in 

dimension, such as shown in Figure 2.275, were tested either with or without venting systems. Five surfaces of each 

specimen were sealed and insulated with glass wool. During a test the bottom surface of the specimen was heated 

to175°C over a 2- to 3-h period, and the temperature was maintained at this level for 91 d. Table 2.18 summarizes 

Figure 2.275 Simulated section of mass concrete wall.

Source: T. Takeda, S. Nakane, and K. Nagao, ―Experimental Studies on Characteristics of Concrete Members 

Subjected to High Temperature,‖ Transactions of 9th International Conference on Structural Mechanics in 

Reactor Technology, Vol. H, pp. 195–200, Lausanne, Switzerland, August 17–21, 1987.
207

the testing conditions for this series of tests. Items measured during a test included temperature, moisture, concrete 

strain, water discharge from the venting system, and compressive strength and modulus of elasticity of concrete after 

heating. Figure 2.276 presents details and measurement positions for a typical specimen. Concrete temperature 

distributions at various times since the start of heating for a vented and nonvented specimen are presented in 

Figure 2.277. Moisture distributions at various times for the vented and unvented specimens are presented in 

Figure 2.278. At 91 d after heating, the moisture distribution showed similar conductivities for normal patterns

Table 2.18 Summary of conditions for simulated mass concrete wall section tests

Source: T. Takeda, S. Nakane, and K. Nagao, ―Experimental Studies on Characteristics of Concrete Members 

Subjected to High Temperature,‖ Transactions of 9th International Conference on Structural Mechanics in 

Reactor Technology, Vol. H, pp. 195–200, Lausanne, Switzerland, August 17–21, 1987.

Figure 2.276 Details of simulated section of mass concrete wall and measurement positions.

Source: T. Takeda, S. Nakane, and K. Nagao, ―Experimental Studies on Characteristics of Concrete Members 

Subjected to High Temperature,‖ Transactions of 9th International Conference on Structural Mechanics in 

Reactor Technology, Vol. H, pp. 195–200, Lausanne, Switzerland, August 17–21, 1987
208

Figure 2.277 Temperature distribution at various times in 

simulated mass concrete wall with and without a venting system.

Source: T. Takeda, S. Nakane, and K. Nagao, ―Experimental Studies on Characteristics of Concrete Members 

Subjected to High Temperature,‖ Transactions of 9th International Conference on Structural Mechanics in 

Reactor Technology, Vol. H, pp. 195–200, Lausanne, Switzerland, August 17–21, 1987

for the two types of specimens, but the high moisture content zone was greater for the nonvented specimen. Water 

discharge from the venting system, shown in Figure 2.279, increased relatively rapidly for the first 7 d of heating 

and then gradually decreased with a total of 150 L (70 L/m2of bottom liner) discharged over the 91-d test duration. 

As the temperature increased, the concrete strains near the bottom liner (heated face) increased, and as heating 

continued the concrete strains at the unheated face increased with time, Figure 2.280. Core samples removed from 

the specimens at conclusion of a test and tested at room temperature were used to determine the effect of heating on 

the concrete’s compressive strength and modulus of elasticity. Test results for strength and modulus of elasticity are 

shown in Figures 2.281 and 2.282, respectively. Reference values for strength and modulus of elasticity obtained 

from water-cured and sealed control specimens are also shown in the appropriate figure. The effect of the

Figure 2.278 Moisture distribution at various times in 

simulated mass concrete wall section with and without a venting system.

Source: T. Takeda, S. Nakane, and K. Nagao, ―Experimental Studies on Characteristics of Concrete Members 

Subjected to High Temperature,‖ Transactions of 9th International Conference on Structural Mechanics in 

Reactor Technology, Vol. H, pp. 195–200, Lausanne, Switzerland, August 17–21, 1987.
209

Figure 2.279 Water discharge from vent pipe of simulated mass concrete wall section.

Source: T. Takeda, S. Nakane, and K. Nagao, ―Experimental Studies on Characteristics of Concrete Members 

Subjected to High Temperature,‖ Transactions of 9th International Conference on Structural Mechanics in 

Reactor Technology, Vol. H, pp. 195–200, Lausanne, Switzerland, August 17–21, 1987.

Figure 2.280 Change in strain distribution with time in 

simulated mass concrete wall section with and without venting.

Source: T. Takeda, S. Nakane, and K. Nagao, ―Experimental Studies on Characteristics of Concrete Members 

Subjected to High Temperature,‖ Transactions of 9th International Conference on Structural Mechanics in 

Reactor Technology, Vol. H, pp. 195–200, Lausanne, Switzerland, August 17–21, 1987.
210

Figure 2.281 Compressive strength test results at selected locations in 

simulated mass concrete wall section with and without venting.

Source: T. Takeda, S. Nakane, and K. Nagao, ―Experimental Studies on Characteristics of Concrete Members 

Subjected to High Temperature,‖ Transactions of 9th International Conference on Structural Mechanics in 

Reactor Technology, Vol. H, pp. 195–200, Lausanne, Switzerland, August 17–21, 1987.

Figure 2.282 Modulus of elasticity test results at selected locations in 

simulated mass concrete wall section with and without venting.

Source: T. Takeda, S. Nakane, and K. Nagao, ―Experimental Studies on Characteristics of Concrete Members 

Subjected to High Temperature,‖ Transactions of 9th International Conference on Structural Mechanics in 

Reactor Technology, Vol. H, pp. 195–200, Lausanne, Switzerland, August 17–21, 1987.
211

elevated temperature was most significant on the concrete modulus of elasticity near the bottom face of the 

specimen, which decreased up to about 40%, relative to sealed control specimens. Compressive strength results at all 

locations in the test specimens exceeded the design strength (240 kg/cm2). 

Although actual data on moisture distribution in nuclear power plant containments is limited, relative humidity 

profiles were obtained at two locations in the Barsebäck BWR containment located in Sweden [2.258]. The 

containment cross-section is somewhat unique in that it consists of an outer 300-mm-thick layer of concrete, a steel 

liner, and an additional inner 800-mm-thick layer of concrete. Relative humidity measurements were made using 

two methods: (1) drilling holes and inserting relative humidity probes, and (2) removal and testing of cores. The 

measurements were made at two locations representing service temperature conditions of 50 ̊C and 25 ̊C. 

Figure 2.283 presents relative humidity results versus distance from the drying surface for the two locations. Since 

the reactor was out of service the in situ measurements were made when the temperature at both locations was 20 ̊ to 

25 ̊C, so relative humidity during service was probably slightly higher than that measured. It was concluded from 

the results that the walls were still drying, even after 30 years, with the drying being extremely slow due to the large

thickness and one-sided drying; drying is different in different parts of the containment due to different temperature 

levels and climatic conditions; and the concrete continued to mature because the moisture conditions in the inner 

parts permitted further cement hydration. Results were used to develop a method for estimating the moisture 

conditions in a nuclear containment wall. The method is based on first quantifying the moisture conditions at the 

outer and inner surfaces, and then describing the moisture fixation and moisture transport within the concrete wall. 

Figure 2.283 Measured relative humidity distributions through Barsebäck containment wall at two locations.

Source: L-O. Nilsson and P. Johansson, ―The Moisture Conditions of Nuclear Reactor Containment Walls – An 

Example for a BWR Reactor,‖ Proceeding of International Conference on Corrosion and Long Term 

Performance of Reinforced Concrete in Nuclear Power Plants and Waste Facilities, NUCPERF 2006, 

Cadarache, France, March 27-30, 2006.

Several researchers have developed models addressing the behavior of concrete materials when subjected to a 

temperature field (i.e., potential for concrete spalling based on temperature distribution and pore pressure build up) 

[2.124,2.259-2.261]. Results of a review of models indicated that although the models are potentially useful for 

modeling internal stresses and moisture transport, none of the models provided validation of the computed pore 

pressures using experimental data [2.262]. Since that review several additional models have been proposed that 

consider the migration of free and bound water, diffusion and migration of dry air and water vapor, and the 

evaporation of free and bound water [e.g. 2.63,2.73,2.263,2.264]. One of the more recent of these is a mathematical 

and computational model developed for chemo-hygro-thermal analysis of concrete when it is subjected to thermal 

load [2.264]. Phase changes of water, changes of material properties caused by temperature and pressure changes, 

as well as coupling among thermal, hydral, and chemical phenomena are taken into consideration. The model is 

implemented through a finite-volume code that was developed. High temperature effects are considered by means 

of temperature and pressure dependence of several parameters. Predicted pore pressure results obtained from the 

model compared to those obtained during step heating of a concrete slab were quantitatively comparable to the 

experimental results obtained from a 30 x 30 x 12 cm3concrete specimen at high temperature [2.246].
212

Summary. Transfer of free moisture in concrete can affect the concrete strength, thermal expansion, shrinkage, and 

creep, as well as potentially affect the concrete durability and radiation shielding capability. Under normal operating 

conditions (e.g., T < 65 ̊C as defined by the ASME Code), moisture migration in mass concrete structures is a very 

slow process and drying is not likely to be an important factor in thick-section concrete structures. Model tests 

involving simulated mass concrete sections heated at one surface and maintained at room temperature at the 

opposite surface indicate that when the surface at ambient conditions is vented, the free moisture is lost over only a 

relatively small length at the heated and unheated surfaces, but shrinkage strains may develop. Above 100 ̊C the 

drying time for concrete becomes much shorter and the moisture diffusivity increases. As the temperature at the 

heated face of a specimen increases, the drying depths increase somewhat. Limited results on moisture distribution 

in a nuclear power plant containment wall indicate that even after 30 years the walls were still drying, with the 

drying being extremely slow due to the large thickness and one-sided drying; drying was different in different parts 

of the containment due to different temperature levels and climatic conditions; and the concrete continued to mature 

because the moisture conditions in the inner parts permitted further cement hydration. Heating of the concrete can 

produce relatively large pore pressures that are a function of temperature, temperature history, and size of specimen. 

The biggest concern relative to pressure build up under postulated accident conditions would be its potential impact 

on the containment metallic liner (e.g., buckling or failure). The pore pressure generated due to heating is about 

equal to the sum of the saturated vapor pressure and partial air pressure for the entrapped pore gases at the 

temperature of the concrete. In unsealed specimens diffusion and eventual drying leads to lower pore pressures than 

that obtained from sealed specimens at the same temperature. Pore pressures dissipate from the hot region of a 

concrete specimen as moisture migrates toward the cooler region. Pore pressures in concrete do not increase 

significantly until the concrete temperature reaches about 105 ̊C. There have been relatively recent modeling 

advancements for chemo-hygro-thermal analysis of concrete when it is subjected to thermal load that consider the 

migration of free and bound water, diffusion and migration of dry air and water vapor, and the evaporation of free 

and bound water. 

2.2.2.8 Simulated Hot Spot Tests

Information and Data. As noted previously, the properties of concrete can be significantly affected by changes in 

temperature. Concrete’s thermal properties are more complex than for most materials because not only is the 

concrete a composite material whose components have different properties, but concrete’s properties depend on 

moisture content and porosity. While the properties of the steel reinforcement are relatively well understood, the 

interaction with concrete is not (i.e., at ambient temperatures the bonding between the reinforcement and concrete is 

considered complete when structural analyses are conducted; however, with an increase in temperature and/or load 

the bond can deteriorate). Prediction of the performance of concrete structural elements at elevated temperatures is 

further complicated due to the presence of cracks that form. At high temperatures, correlation of cracking patterns 

predicted by analytical procedures with experimental results is difficult [2.265]. Due to the problems involved in the 

analytical treatment of concrete structural members, especially under elevated-temperature conditions, model tests 

are often used to develop data under representative conditions. The results of these tests are then used to both 

demonstrate performance and for the validation and refinement of computer codes. 

Several research projects have been conducted to investigate the behavior of reinforced concrete structures at 

elevated temperature; however, the overall level of effort has not been sufficient for establishment of widely 

accepted elevated-temperature concrete design procedures. Available results have been primarily concerned with 

testing of specific structural features in support of development of analytical procedures, or model tests related to 

development of gas-cooled or breeder reactors. A compilation of elevated temperature-related structural features 

tests, model tests in support of prestressed concrete reactor vessel development, and thermal- and moisture 

migration-model tests was completed under a prior study [2.266] and will not be repeated. Provided below is 

information related to tests involving ―hot spots‖ that could develop under conditions such as penetration of steam 

lines through a containment wall or if localized failure of the liner cooling system of a prestressed concrete reactor 

vessel occurs. 

A thermal cylinder experiment was designed both to provide information for evaluating the capability of analytical 

methods to predict the time-dependent stress-strain behavior of a 1:6-scale model of the barrel section of a singlecavity prestressed concrete reactor vessel and to demonstrate the structural behavior under design and simulated 

thermal conditions such as could result from an accident (i.e., hot spot) [2.267]. The model shown in Figure 2.284 

was a thick-walled cylinder having a height of 1.22 m, a thickness of 0.46 m and an outer diameter of 2.06 m. It was 
213

Figure 2.284 Isometric of ORNL thermal cylinder test structure.

Source: J.P. Callahan et al., ―Prestressed Concrete Reactor Vessel Thermal Cylinder Model Study,‖ ORNL/TM5613, Oak Ridge National Laboratory, June 1977.

prestressed both axially and circumferentially and subjected to a 4.83-MPa internal pressure together with a thermal 

crossfall imposed by heating the inner surface to 65.7°C and cooling the outer surface to 24°C. Because the model

was designed to study the behavior of the barrel section of a massive concrete structure, all exposed surfaces were 

sealed to prevent loss of moisture, and the ends of the cylinder were insulated to prevent heat flow in the axial 

direction. The experiment utilized information developed from previous studies of concrete material properties, 

triaxial creep, instrumentation, analysis methods, and structural models. The initial 460 days of testing was divided 

into time periods that simulated prestressing, heatup, reactor operation, and shutdown. At the conclusion of the 

simulated operating period, the model was repressurized and subjected to localized heating at 232°C for 84. This 

temperature excursion was meant to simulate a hot spot condition in which the liner cooling system and/or insulation 

system is assumed to have failed in an operating nuclear reactor. At the end of this period the inner surface was 

cooled to 66 ̊C using the inner water-circulation system, followed by cooling to 24 ̊C. Comparisons of experimental 

data with calculated values obtained using the SAFE-CRACK finite-element computer program showed that the 

program was capable of predicting time-dependent behavior in a vessel subjected to normal operating conditions, 

but that it was unable to accurately predict the behavior during off-design hot-spot heating. Readings made using a 

neutron and gamma-ray backscattering moisture probe showed little, if any, moisture migration in the concrete 

cross-section. Destructive examination (sectioning by saw cutting) indicated that the model maintained its basic 

structural integrity during localized hot spot heating. No significant defects were seen in the cut surfaces, however, 

the concrete immediately surrounding the hot-spot heating elements was darkened as shown in Figure 2.285 

indicating that some change had occurred in the concrete due to the 232 ̊C hot-spot heating. A series of impact 

hammer readings was made in and outside this area and revealed that the relative compressive strength in the hotspot area was lower (~39%) than that obtained for readings outside the heat-affected zone. As the distance from the 

heater increased the reduction in compressive strength decreased. 
214

Figure 2.285 Cross section showing discoloration in the vicinity of the hot-spot heating.

Source: J.P. Callahan et al., ―Prestressed Concrete Reactor Vessel Thermal Cylinder Model Study,‖ ORNL/TM5613, Oak Ridge National Laboratory, June 1977.

During commissioning tests of the prestressed concrete reactor vessel at the Oldbury Nuclear Power Station, there 

were a small number of localized breakdowns of the liner insulation permitting the temperature to reach 180°C in 

the head penetration region and 90°C in the haunch region at the upper boiler instrument penetration [2.268]. 

Since hot spots could induce high thermal stresses in the concrete to result in cracking, a theoretical study was 

conducted that indicated there was no cause for concern about the safety margin against failure of the vessel, but 

concrete cracking could occur close to the liner. To provide input on concrete cracking, a full-scale model (3.66 mdiameter by 1.52 m thick) of the region of the vessel local to the upper boiler instrumentation where the highest liner 

temperatures were recorded was fabricated and tested, Figure 2.286. The test procedure included heating of the 

model over a 24-h period to the steady-state condition achieved in the hot spot region of the prototype, allowing the 

model to attain thermal equilibrium, and maintaining this condition for 3 months with the prestressing force reduced 

as the test progressed, and then permitting the model to cool to ambient. Embedded strain instrumentation, dye 

penetrants, core samples, and ultrasonic testing were used to detect and measure cracking. Results indicated that the 

hot spot was confined to a small area of the model. Gages indicated that some cracking had occurred during reactor 

start-up when the penetration liner was only 24°C hotter than the adjacent concrete and was attributed to thermal 

diffusivity differences between steel and concrete resulting in strain differentials. Cracking was limited to a central 

region of about 0.46 m where the concrete temperature was 100°C. No cracking of the concrete was observed at 

heated locations away from steel parts. A second degree of cracking occurred during shutdown that was caused by 

stress reversals due to residual creep strains. Uncracked regions of the heated concrete showed no loss of strength 

over control samples stored separately. In addition, measurements of vapor pressure behind the liner showed less 

severe pressures than those corresponding to the concrete temperatures imposed. This indicated no pressure build 

up would occur at the hot spot, probably because of pressure relief along the liner-concrete interface. It was 

concluded that the sustained high hot-spot temperature in the Oldbury vessel did not cause serious damage to the 

liner or to the concrete.
215

Figure 2.286 Full-scale Oldbury hot-spot model.

Source: J. Irving, G.D.T. Carmichael, and I.W. Hornby, ―A Full Scale Model Test of Hot Spots in the Prestressed 

Vessels of Oldbury Nuclear Power Station,‖ Paper 7699, Proceedings of Institution of Civil Engineers 57, 

June 1974.

Summary. Due to problems involved in analytical treatment of concrete structural members at elevated temperature 

(e.g., property changes and cracking), structural features tests have often been used to develop data under 

representative conditions. The results of these tests are then used to both demonstrate performance and for the 

validation and the refinement of computer codes. A summary of structural features tests was provided in an earlier 

study [2.266]. One area of interest is related to hot spots that could develop in the vicinity of hot penetrations or if a 

localized failure of a liner cooling system occurs in a prestressed concrete reactor vessel. Although results are 

limited, they tend to indicate that for the conditions investigated ( T = 232 ̊C or 180 ̊C) the effect was localized. 

Some loss of concrete strength and cracking will occur in the region of the hot spot, but is not considered to be a 

threat to structural integrity. Results also seemed to indicate that significant pressure build up behind a liner 

resulting from a localized hot spot may not occur due to pressure relief along the liner-concrete interface.
216

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2.155 H. G. Geymayer, ―Effect of Temperature on Creep of Concrete: A Literature Review,‖

Farmington Hills, Michigan,

2.156 K.W Nasser and A.M. Neville, ―Creep of Concrete at Elevated Temperatures,‖ Proceedings Journal of 

American Concrete Institute 64(12), pp. 1567-1579, Farmington Hills, Michigan, 1965.

2.157 K.W Nasser and A.M. Neville, ―Creep of Old Concrete at Normal and Elevated Temperature,‖ Proceedings 

Journal of American Concrete Institute 64(2), pp. 97-103, Farmington Hills, Michigan, 1967.

2.158 G.L. England and A.D. Ross, ―Reinforced Concrete Under Thermal Gradients,‖ Magazine of Concrete 

Research 14(40), pp. 5-12, March 1962.

2.159 D.J. Hannant, The Strain Behavior of Concrete Under Compressive Stress at Elevated Temperatures, 

Laboratory Note RD/L/N 67/66, Central Electricity Research Laboratories, United Kingdom, June 1966.

2.160 S. Arthanari and C.W. Yu, ―Creep of Concrete Under Uniaxial and Biaxial Stresses at Elevated 

Temperatures,‖ Magazine of Concrete Research 19(60), pp. 149-156, September 1967.

2.161 K.B. Hickey, Creep, Strength, and Elasticity of Concrete at Elevated Temperatures, Report No. C-1257, 

Bureau of Reclamation, Washington, DC, December 1967.

2.162 J.C. Marechal, ‖Causes Physiques et Chimiques du Fluage et du des Béton,‖ RILEM Colloquium on the 

Physical and Chemical Causes of Creep and Shrinkage of Concrete, Materials and Structures 2(8), pp. 852-

855, June 1968.

2.163 J.C. Marechal, ―Evolution des Propriétés Thermiques et Méchaniques des Bétons et Autres Matériaux en 

Fonction de la Température,‖ Annales Travaux Publics 21(246), pp. 852-855, June 1968.
223

2.164 A.E. Theuer, ―Effect of Temperature on the Stress-Deformation of Concrete,‖ Journal of Research 18(2), 

pp. 195-204, National Bureau of Standards, Gaithersburg, Maryland, February 1937.

2.165 A.F. Da Silveira and C.A. Florentino, ―Influence of Temperature on the Creep of Mass Concrete,‖ 

American Concrete Institute Symposium on the Effect of Temperature on Concrete, Memphis, Tennessee, 

November 5-8, 1968.

2.166 C.H. Wang, ―Creep of Concrete at Elevated Temperatures,‖ Paper SP 27-19 in Designing for Effects of 

Creep, Shrinkage, Temperature in Concrete Structures,‖ Farmington Hills, 

Michigan, 1992.

2.167 R.D. Browne and R. Blundell, ―The Influence of Loading Age and Temperature on the Long Term Creep 

Behavior of Concrete in a Sealed, Moisture Stable, State,‖ Materials and Structures 2(8), pp. 133-144, 

March 1969.

2.168 J. Komendant. V. Nicolayeff, M. Polivka, and D. Pirtz, ―Effects of Temperature, Stress Level, and Age at 

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2.169

2.170 T.W. Kennedy, An Evaluation and Summary of a Study on the Long-Term Multiaxial Creep Behavior of 

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2.171 J.E. McDonald, Time-Dependent Deformation of Concrete Under Multiaxial Stress Conditions, Technical 

Report C-75-4, U.S. Army Engineer Waterways Experiment Station, Vicksburg, Mississippi, October 

1975.

2.172 H. Kasami, T. Okuno, and S. Yamane, ―Properties of Concrete Exposed to Sustained Elevated 

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2.174 H. V. Reichel, ―How Fire Affect Steel-to-Concrete Bond,‖ Building Research and Practice 6(3), pp. 176–

186, May/June 1978.

2.175 K. Hertz, ―The Anchorage Capacity of Reinforcing Bars at Normal and High Temperatures,‖ Magazine of 

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2.176 U. Diederichs and U. Schneider, ―Bond Strength at High Temperature,‖ Magazine of Concrete Research

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2.177 C-H. Chiang and C-L. Tsai, ―Time-Temperature Analysis of Bond Strength of a Rebar After Fire 

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2.178 H. Sager and F.S. Rostasy, ―High Temperature Behavior of Reinforcing and Prestressing Steels,‖ 

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2.180 P.D. Morley and R. Royles, ―Response of the Bond in Reinforced Concrete to High Temperatures,‖ 

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2.181 R. Royles and P.D. Morley, ―Further Responses to the Bond in Reinforced Concrete to High 

Temperatures,‖ Magazine of Concrete Research 35(124), pp. 157-163, September 1983.

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April 2000.
224

2.186 A. Menou, G. Mounajed, H. Boussa, A. Pineaud, and H. Carre, ―Residual Fracture Energy of Cement 

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375, American Concrete Institute, Farmington Hills, Michigan, 1972.

2.190 C. DeFigh-Price et al., ―Effects of Long-Term Exposure to Elevated Temperature on Mechanical Properties 

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1985.

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to Long-Term Exposure to High Temperatures, Report No. U95037, Abiko Research Laboratory, Central 

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pp. 614–615, 1967.

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Concrete Shield,‖ Nippon Genshiryoku Gakkaishi 11, pp. 2–8, 1969.

2.200 F. Seboek, ―Shielding Effectiveness of Cracked Concrete,‖ Kerntechnik 12, pp. 4496–4501, November 

1970.

2.201 H.E. Hungerford, A. Hönig, A.E. Desov, F. Du Bois, and H.S. Davis, ―Concretes, Cements, Mortars, and 

Grouts,‖ Engineering Compendium on Radiation Shielding, Vol. II, pg. 117, R.G. Jaeger (Editor), 

Springer-Verlag, Berlin, Germany, 1975.

2.202

2.203 F. Bremer, "On a Triaxial Strength Criterion for Concrete," SP-34-14 in Concrete for Nuclear Reactors, 

pp. 283-294, American Concrete Institute, Farmington Hills, Michigan, 1972. 

2.204 D. Linse, "Strength of Concrete Under Biaxial Sustained Load" SP-34-17 in Concrete for Nuclear 

Reactors, pp. 327-334, American Concrete Institute, Farmington Hills, Michigan, 1972.

2.205 J. Chinn and R.M. Zimmerman, Behavior of Plain Concrete Under Various High Triaxial Compression 

Loading Conditions, WL TR 64-163, Air Force Weapons Laboratory, August 1965. 

2.206 C.D. Goode and M.A. Helmy, "The Strength of Concrete Under Combined Shear and Direct Stress," 

Magazine of Concrete Research 19(59), June 1967. 

2.207 N. J. Gardner, "Triaxial Behavior of Concrete," Journal American Concrete Institute 66(2), Farmington 

Hills, Michigan, February 1969. 
225

2.208 V. Hansson and K. Schimmelpfennig, "Concrete Strength in Multiaxial Stress States," Paper SP 34-15 in 

Concrete for Nuclear Reactors, pp. 295-312, American Concrete Institute, Farmington Hills, Michigan, 

1972.

2.209 G. Schickert and H. Winkler, Results of Test Concerning Strength and Strain of Concrete Subjected to 

Multiaxial Compressive Stresses, Deutscher Auschuss für Stahlbeton, Heft No. 277, Berlin, Germany, 

1977.

2.210 R.H. Atkinson and H.Y. Ko, "A Fluid Cushion, Multiaxial Cell for Testing Cubical Rock Specimens," 

International Journal of Rock Mechanics and Mineral Sciences 10, pp. 35l-361, 1973. 

2.211 E. Andrenes, "Response of Mortar to Biaxial Compression," M.S. Thesis, University of Colorado, 1974. 

2.212 M.A. Taylor and B.K. Patel, "The Influence of Path Dependency and Moisture Conditions on the Biaxial 

Compression Envelope for Normal Weight Concrete," Journal American Concrete Institute 71(12), 

Farmington Hills, Michigan, December 1974. 

2.213 H. B. Kupfer, "Nonlinear Behavior of Concrete in Biaxial Stress," Beton-StahZbetonbau 11, pp. 26-74, 

1973.

2.214 T.N.W. Akroyd, ―Concrete Under Triaxial Stress,‖ Magazine of Concrete Research 13(39), pp. 111-118, 

November 1961.

2.215 J. Isenberg, ―Strength of Concrete Under Combined Stress,‖ Civil Engineering Publics Works Review, 

October 1965.

2.216 R.M. Zimmerman and L.A. Traina, Strength and Deformation Response of Concrete Under Multiaxial 

Loadings – Cooperative Project, Final Report EN676-12097, Engineering Experiment Station, New 

Mexico State University, May 1977.

2.217 F.K. Garas, ―Strength of Concrete Under Different States of Stress,‖ Paper SP 34-18 in Concrete for 

Nuclear Reactors, pp. 335-343, American Concrete Institute, Farmington Hills, Michigan, 1972.

2.218 K.H. Gerstle, D.L. Linse, P. Bertacchi, M.S. Kotosovos, H-Y. Ko, J.B. Newman, P. Rossi, G. Schickert, 

M.A. Taylor, L.A. Traina, R.M. Zimmerman, and R. Bellotti, ―Strength of Concrete Under Multiaxial 

Stress States,‖ Paper SP 55-5 in Douglas McHenry International Symposium on Concrete and Concrete 

Structures, pp. 103-131, American Concrete Institute, Farmington Hills, Michigan, 1978.

2.219 W. Ren, W. Yang, Y. Zhou, and J. Li, ―Behavior of High-Performance Concrete Under Uniaxial and 

Biaxial Loading,‖ Title No. 105-M62, ACI Materials Journal, American Concrete Institute, Farmington 

Hills, Michigan, November-December 2008.

2.220 S-K. Lee, Y-C. Song, and S-H. Han, ―Biaxial Behavior of Plain Concrete of Nuclear Containment 

Building,‖ Nuclear Engineering and Design 227, pp. 143-153, 2004.

2.221 K.-Ch. Thienel and F. S. Rostásy, ―Influences of Concrete Composition on Strength and Deformations 

Under Uniaxial and Biaxial Loading at Elevated Temperature,‖ Paper H04/6 in Transactions of the 12th

International Conference on Structural Mechanics in Reactor Technology, pp. 145–150, 1993.

2.222 B.L. Hu, Y.P. Song, and G.F. Zhao, ―Test on Strength and Deformation of Concrete Under Complex Stress 

at Elevated Temperature,‖ Building Science Research 20(1), Sichuan, pp. 47–50, 1994.

2.223 D.J. Hannant, ―The Effects of Heat on Concrete Strength,‖ Engineering 197(5105), p. 302, February 21, 

1964. 

2.224 M. Petkovski, R.S. Crouch, and P. Waldron, ―Apparatus of Testing Concrete Under Multiaxial 

Compression at Elevated Temperature (max2T),‖ Experimental Mechanics 46, pp. 387-398, 2006.

2.225 M. Petkovski and R.S. Crouch, ―Strains Under Transient Hygro-Thermal States in Concrete Loaded in 

Multiaxial Compression and Heated to 250 ̊C,‖ Cement and Concrete Research 38, pp. 586-596, 2008.

2.226 ACI COMMITTEE 207, ―Mass Concrete for Dams and Other Massive Structures,‖ Proceedings of the 

Journal of American Concrete Institute 67(4), pp. 273–309, April 1970.

2.227 R. Philleo, Some Physical Properties of Concrete At High Temperature, Research Department Bulletin 97, 

Portland Cement Association, Skokie, Illinois, October 1958.

2.228 K.-Y. Shin, J-H. Kim, M. Chung, and P-S. Jung, ―Thermo-Physical Properties and Transient Heat Transfer 

of Concrete at Elevated Temperatures,‖ Nuclear Engineering and Design 212, pp. 233–241, 2002.

2.229 H. Dettling, The Thermal Expansion of Hardened Cement Paste, Aggregates, and Concrete, Bulletin 

No. 164, pp. 1-64, Deutscher Ausschus für Stahlbeton, W. Ernst & Sohn, Berlin, Germany, 1964.

2.230 J.M. Griffiths, Thermal Expansion of Typical American Rocks, Bulletin No. 128, Iowa Engineering 

Experiment Station, Iowa State College, Ames, 1936.

2.231 D.G.R. Bonnell and F.C. Harper, ―The Thermal Expansion of Concrete,‖ National Buildings Studies, 

Technical Paper No. 7, Her Majesty’s Stationary Office,, London, United Kingdom, 1951. 

2.232 Concrete Manual, 7th Ed., U.S. Bureau of Reclamation, Denver, Colorado, 1963.
226

2.233 S.L. Meyers, ―Thermal Coefficient of Expansion of Portland Cements – Long-Time Tests,‖ Industrial and 

Engineering Chemistry 32(8). pp. 1104-1112, 1940.

2.234 C.R. Cruz and M. Gillen, ―Thermal Expansion of Portland Cement Paste, Mortar, and Concrete at High 

Temperature,‖ Fire and Materials 4(2), 1980.

2.235 H. C. Hirth, M. Polivka, and D. Pirtz, Final Report on Thermal Properties of Concrete at Elevated 

Temperature, ORNL/BRP-81/1 (Limited Distribution), Oak Ridge National Laboratory, Oak Ridge, 

Tennessee, July 1981.

2.236 V.K.R. Kodur and M.A. Sultan, ―Effect of Temperature on Thermal Properties of High-Strength Concrete,‖ 

Journal of Materials in Civil Engineering 15(2), pp. 101-107, American Society of Civil Engineers, New 

York, New York, April 2003.

2.237 F. Vodak, R. Cerny, J. Drchalová, S. Hosková, O. Kapicková, O. Michalko, P. Semerák, and J. Toman, 

―Thermophysical Properties of Concrete for Nuclear Safety-Related Structures,‖ Cement and Concrete 

Research 27(3), pp. 415–426, 1997.

2.238 H. Abe, T. Kawahara, T, Ito, and A. Haraguchi, ―Influence Factors of Elevated Temperatures on Thermal 

Properties and Inelastic Behavior of Concrete,‖ Paper SP 34-40 in , pp. 847-870, 

American Concrete Institute, Farmington Hills, Michigan, 1972. 

2.239 S. L. Meyers, ―How Temperature and Moisture Changes May Affect the Durability of Concrete,‖ Rock 

Products, pp. 153–157, August 1951.

2.240 R. Felicetti, P.G. Gambarova, G.P. Rosati, The Mechanical Properties of High Performance Siliceous 

Concrete Exposed to High Temperature; ―Penly‖ Concrete; Uniaxial Compression, Research Project 

1794/20-3-95, Milan University of Technology, Milan, Italy, March 1995.

2.241 K-Y. Shin, S-B. Kim, J-H. Kim, M. Chung, and P-S. Jung, ―Thermo-Physical Properties and Transient 

Heat Transfer of Concrete at Elevated Temperatures,‖ Nuclear Engineering and Design 212, pp. 233–241,

2.242 U. Schneider and U. Diederichs, ―Physical Properties of Concrete From 20 ̊C Up To Melting,‖ Parts 1 and 

2, Betonwerk & Fertigteiltechnik, Heft 3, pp. 141-150, and Heft 4, pp. 223-230, 1981.

2.243

2.244 G. Hildenbrand, M. Peeks, A. Skokan, and M. Reimann, Untersuchung der Wechselvirkung von 

Kernschmelze und Reaktorbeton, Abschlussbericht Förderungverhaben BMFT RS 154, KWU, Erlangen, 

Germany, May 1978.

2.245 Y. Ichikawa and G.L. England, ―Prediction of Moisture Migration and Pore Pressure Build-up in Concrete 

at High Temperatures,‖ Paper #1122, Proceedings of 16thtInternational Conference on Structural 

Mechanics in Reactor Technology, Elsevier Science Publishers, North-Holland, The Netherlands, 2001.

2.247 Z.P. Bazant and L.J. Najjar, ―Nonlinear Water Diffusion in Nonsaturated Concrete,‖ Materials and 

Structures 5, pp. 3-20, 1972.

2.248 A.D. Ross, M. Illston, and G.L. England, ―Short- and Long-Term Deformations of Concrete Influenced by 

Its Physical Structure and State,‖ Proceedings of International Conference on Structural Concrete, 

Paper H1, Imperial College, London, United Kingdom, 1965.

2.249 M. Jooss and H.W. Reinhardt, ―Permeability and Diffusivity of Concrete as a Function of Temperature,‖ 

Cement and Concrete Research 32, pp. 1497-504, 2002,

2.250 Z.P. Bazant and W. Thonguthai, ―Pore Pressure and Drying of Concrete at High Temperature,‖ Journal of 

Engineering Mechanics Division 4(EM5), Paper 14077, pp. 1059-1079, Proceedings of American Society 

of Civil Engineers, New York, New York, 1978.

2.251 L.T. Phan, ―Pore Pressure and Explosive Spalling,‖ Materials and Structures 41(10), pp; 1623-1632, 2008.

2.252 G.L. England and T.J. Sharp, ―Migration of Moisture and Pore Pressure in Heated Concrete,‖ Proceedings 

of 1st International Conference on Structural Mechanics in Reactor Technology, 

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1010071, Electric Power Research Institute, Palo Alto, California, December 2005.

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Water Release from Heated Concrete, GEFR-00521, General Electric, April 1980. 
227

2.255 G.L. England and A.D. Ross, ―Shrinkage, Moisture, and Pore Pressure in Heated Concrete,‖ Paper SP 34-

42 in Concrete for Nuclear Reactors, pp. 883-907, American Concrete Institute, Farmington Hills, 

Michigan, 1972.

2.256 J. E. McDonald, Moisture Migration in Concrete, Technical Report C-75-l, U.S. Army Waterways 

Experiment Station, Vicksburg, Mississippi, May 1975.

2.258 L-O. Nilsson and P. Johansson, ―The Moisture Conditions of Nuclear Reactor Containment Walls – An 

Example for a BWR Reactor,‖ Proceeding of International Conference on Corrosion and Long Term 

Performance of Reinforced Concrete in Nuclear Power Plants and Waste Facilities, NUCPERF 2006, 

Cadarache, France, 27-30 March 2006.

2.259 Z. P. Bazant and W. Thonuthai, ―Pore Pressure in Heated Concrete Walls: Theoretical Prediction,‖ 

Magazine of Concrete Research 31(107), pp. 67–76, 1979.

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Concretes at High Temperatures,‖ Proceedings of the Fourth Weimar Workshop on High Performance 

Concrete: Material Properties and Design, held at Hochschule für Architektur und Bauwesen (HAB), 

Weimar, Germany, pp. 53–68, October 4–5, 1995.

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119 Standard Fire Condition,‖ Journal of Fire Protection Engineering 7(4), pp. 125–132, 1995.

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National Institute of Standards and Technology, Gaithersburg, Maryland, 1996.

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2.266

2.267 J. P. Callahan, D.A. Canonico, M. Richardson, J.M. Corum, W.G. Dodge, G.C. Robinson, and G.D. 

Whitman, Prestressed Concrete Reactor Vessel Thermal Cylinder Model Study, ORNL/TM-5613, Oak 

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1974. 
228
229

3 RADIATION SHIELDING CONCRETES

3.1 Introduction

Portland cement concrete possesses many of the physical qualities of an ideal radiation shield. It is a polyphase 

material consisting of particles of aggregate contained in a matrix of Portland cement paste. Gamma rays are 

absorbed by the high-density aggregate materials and neutrons are attenuated by hydrogen atoms in the cement 

paste. A concrete shield is exposed to two sources of heat: heat transferred from hot parts of the reactor system and 

heat produced internally by the attenuation of neutrons and gamma rays [3.1]. Energy captured from slowed down 

fast neutrons and gamma rays entering the shield from the reactor core is deposited within the shield material and 

liberated as heat. The total amount of heat generated can be considerable. The heat generated may have detrimental 

effects on the physical, mechanical, and nuclear properties of the concrete. Different types of concrete perform 

differently under radiation exposure, although if heated to relatively high temperatures they all will lose waters of 

crystallization and become somewhat weaker and less effective in neutron attenuation [3.2]. Although limited, some 

information is available on the effects of elevated temperature on radiation shielding concretes [3.3]. Provided 

below is a brief summary of the effect of elevated temperature exposure on properties and shielding effectiveness. 

3.2 Heavyweight Concretes

Concrete for radiation shielding typically uses ordinary normal weight aggregates, however, special types of 

aggregate have been used to improve the attenuation properties of concrete or to reduce the thickness of concrete 

shields. Typically this involves incorporation of heavyweight aggregate materials.

Compared to normal weight aggregate concrete having a typical unit weight of 2400 kg/m3, heavyweight concretes 

weigh from 2900 to 6100 kg/m3. Typical heavyweight aggregate materials for use in production of heavyweight 

concrete are listed in Table 3.1. Several specifications or standard practice guidelines have been developed that 

address heavyweight aggregates and heavyweight concretes [3.4-3.12]. More detailed information describing 

aggregate materials for radiation-shielding concretes and composition and mechanical properties of concretes that 

have been used for radiation shielding is available elsewhere [3.3].

Table 3.1 Examples of aggregate materials used to produce heavyweight concretes

Type aggregate Chemical composition of 

principal mineral

Specific gravity of pure 

mineral

Typical bulk density 

(kg/m3)

Witherite BaCO3 4.29 2320

Baryte BaSO4 4.50 2560

Magnetite Fe3O4 5.17 2720

Hematite Fe2O3 4.9-5.3 3040

Lepidocrocite

Geothite

Limonite

Hydrous iron ores 

containing 

8-12% water

3.4-4.0 2240

Ilmenite FeTiO3 4.72 2560

Ferrophosphorus Fe3P, Fe2P, FeP 5.7-6.5 3680

Steel aggregate Fe 7.8 4480

Source: P.K. Mehta and P.J.M. Monteiro, Concrete – Microstructure, Properties, and Materials, McGraw Hill, 

New York, New York, 2006.

Different types of cement have been used in concrete for radiation shielding such as Portland cement, high-alumina 

cement, gypsum-alumina cement, magnesium oxychloride cement, magnesia cement, phosphate cement, and 

oxyacid cement [3.3]. Specialty cements, such as gypsum-alumina, have been used to increase the chemicallybound water content to provide additional hydrogen for neutron shielding. Table 3.2 presents examples of ranges of 

properties that have been reported in the literature for heavyweight aggregate concretes [3.3]. Differences in 
230

properties are due to factors such as proportions and the nature of the materials used in the mixes as well as the test 

conditions. Additional information on mix constituents and test conditions can be obtained by consulting the 

references provided in the table.

Table 3.2 Examples of mechanical and physical properties 

reported in literature for heavyweight aggregate concretes

Property

Type Concrete

Baryte Serpentine Limonite Hematite Ferrophosphous Magnetite Limoniteiron

Density

(g cm-3)

3.5-3.7 2.06-2.2 2.96 3.73-4.2 4.65 3.41-4.38 3.62-5.34

Compressive

Strength 

(MPa)

24.8-42.2 13.1-15.8 40.4 16.2-89.3 30.4 19.2-41.8 38.4-78.0

Modulus 

(GPa)

29.9-35.5 30.7 37.3-69.1 28.2 31.2-62.0 43.4-50.4

Tensile 

strength (MPa)

2.25-4.9 6.5 3.6-4.0

Thermal 

expansion 

(10-6 ̊C-1)

18 ~32 10.35 10.3 6.8-10.0

Thermal

Conductivity

(W m-1 ̊C-1)

0.961-

1.621

0.8 2.914 2.7-3.5

Specific

Heat

(cal g-1 ̊C-1)

0.123-

0.157

0.17 0.17-0.21

Diffusivity

(m2h

-1

)

0.002 0.003 0.32

Data Source 3.13-3.17 3.18 3.19,3.20 3.21,3.22 3.23 3.19.3.20,

3.15,3.24

3.20,3.25

Primary source: M.F. Kaplan, Concrete Radiation Shielding, Concrete Design and Construction Series, Longman 

Scientific & Technical, New York, New York, 1989.

3.3 Effect of Elevated Temperature on Properties

Concrete radiation shields may be exposed to external sources of heat as well as heat produced within the shield as a 

result of attenuation or absorption of gamma and neutron radiation. Generation of heat can produce thermal stresses 

and affect the mechanical and physical properties of the concrete. Although limited, some data have been assembled 

on the effects of elevated temperature on the mechanical and physical properties of shielding concretes [3.3]. 

Information, primarily from this reference, for several of the concretes listed in Table 3.2 will be summarized below.

3.3.1 Serpentine Aggregate-Based Concrete

The effect of elevated temperature on the residual compressive strength of a serpentine and an ophicalcite concrete 

has been investigated [3.26]. The serpentine concrete mix had a dry density of 2085 kg/m3and a 28-day 

compressive strength of 19.3 MPa. The cylindrical test specimens were 10-cm-diameter by 20-cm-long. The 

specimens were cured for seven weeks in water at 20 ̊C, permitted to air dry for several weeks (e.g., 83 to 

92 weeks), and then heated to temperatures of 100 ̊, 200 ̊, or 600 ̊C where they remained for 30 minutes. Part of the 

specimens were immediately removed from the oven and tested and part were permitted to cool to room temperature 

prior to testing. Figure 3.1 presents the effect of elevated temperature on the strength (% of reference room 

temperature value) of the serpentine concrete tested hot and the ophicalcite concrete (91-day compressive strength of 
231

30.7 MP) with specimens tested while hot or permitted to cool to room temperature prior to testing. Also shown in 

the figure is the effect of temperature on the density of the serpentine concrete. 

Figure 3.1 Effect of temperature on the compressive strength of a serpentine (tested hot) and ophicalcite (residual 

and tested hot) aggregate concrete, and on density of a serpentine concrete.

Source: S. Ohgishi, S. Miyasaka, and J. Chida, ―On Properties of Magnetite and Serpentine Concrete at Elevated 

Temperatures for Nuclear Reactor,‖ SP 34-57 in Concrete for Nuclear Reactors, pp. 1243-1253, Special 

Publication 34, American Concrete Institute, 1972.

The effects of elevated temperature on the properties of iron shot, limonite, magnetite, and ferrophosphorous 

concretes as well as various combinations of these aggregate materials was investigated in conjunction with the 

N-Production Reactor located at Hanford, Washington [3.23]. Concrete placement was by conventional methods as 

well as by the preplaced aggregate method. Table 3.3 presents the effect of elevated temperature on selected 

properties of a serpentine-iron concrete in which the aggregates were preplaced. The effects of elevated temperature 

on mechanical and physical properties of a magnetite-serpentine aggregate concrete are presented in Table 3.4. 

Magnetite was used as the coarse aggregate and serpentine as the fine aggregate. Approximately 46% of the total 

aggregate weight was serpentine. Type II Portland cement was used in the mix and placement was by conventional 

means. Application of 20 thermal cycles from 40 ̊ to 200 ̊ to 40 ̊C reduced the compressive strength by 11%, and if 

the peak cycle temperature was increased to 350 ̊C the reduction was 21%. Modulus of rupture results for the two 

cycling scenarios were reduced 30% and 77%, respectively. The static modulus of elasticity was reduced by 36% 

after 20 thermal cycles from 40 ̊ to 200 ̊ to 40 ̊C. Residual compressive strengths of cube specimens soaked in 

water for two weeks prior to heating to temperatures up to 300 ̊C were less than results obtained for specimens that 

had not been soaked prior to heating with the dry specimens exhibiting residual strengths at temperatures up to 

350 ̊C of 98%, or more. 

3.3.2 Limonite Aggregate-Based Concrete

Compressive strength ratios for limonite concrete 100-mm cubes tested at elevated temperature have been reported 

as 94, 88, and 85% the reference room temperature strength after exposure from 3 to 7 hours at 100 ̊, 150 ̊, and 

200 ̊C, respectively, [3.27]. Average coefficients of thermal expansion of Portland cement concretes containing 
232

Table 3.3 Effect of temperature on properties of a serpentine-iron concrete

Source: H.S. Davis, ―N-Reactor Shielding,‖ Paper SP-34-52 in Concrete for Nuclear Reactors, pp. 1109-1161, 

American Concrete Institute, Farmington Hills, Michigan, 1972.

Table 3.4 Effect of temperature on properties of a magnetite-serpentine concrete

Source: H.S. Davis, ―N-Reactor Shielding,‖ Paper SP-34-52 in Concrete for Nuclear Reactors, pp. 1109-1161, 

American Concrete Institute, Farmington Hills, Michigan, 1972.
233

limonite aggregate have been reported as 4.86 x 10-6  ̊C-1for temperatures up to 300 ̊C, 4.5 x 10-6 ̊C-1between 300 ̊ 

and 600 ̊C, and 4.2 x 10-6 ̊C-1between 600 ̊ and 800 ̊C [3.28].

Table 3.5 presents results for a hydrous-iron aggregate concrete whose aggregate was a mixture of limonite, 

magnetite, and haemitite [3.29]. The concrete was made from Type II Portland cement and had a water-cement ratio 

of 0.43 (by weight). Specimens were moist cured for 28 days at 23 ̊C, stored in laboratory air at about 50% relative 

humidity for an additional 62 days, heated for a period of two weeks to temperatures of 85 ̊, 200 ̊, or 350 ̊C, and 

then permitted to cool to room temperature prior to testing.

Table 3.5 Effect of elevated temperature on the physical and 

mechanical properties of hydrous-iron aggregate concrete

Source: H.S. Davis and O.E. Borge, ―High-Density Concrete Made With Hydrous-Iron Aggregates,‖ Journal 

American Concrete Institute 30(10), pp. 1141-1147, Farmington Hills, Michigan, 1959.

3.3.3 Magnetite Aggregate-Based Concrete

Tests have been conducted on 10-cm diameter by 20-cm long cylindrical test specimens of magnetite concrete that 

were initially cured in water at 20 ̊C for twelve weeks and then stored in moist air for 92 weeks prior to heating 

[3.26]. The specimens were heated to temperatures from 100 ̊ to 600 ̊C with the final temperature maintained for 30 

minutes. Some of the specimens were tested at temperature and others permitted to cool to room temperature prior 

to testing. The density of the magnetite concrete was found to decrease approximately linearly as the temperature 

increased with the decrease being about 4% at 600 ̊C. Figure 3.2 presents the effect of elevated temperature on the 

compressive strength of the magnetite concrete for specimens tested at temperature and for specimens that returned 

to room temperature prior to testing. Also shown in the figure are results obtained for a normal weight concrete.

Table 3.6 presents the effect of elevated temperature on the properties of a magnetite concrete used for the radiation 

shield of the EL4 reactor in France [3.30]. The concrete utilized Portland cement, magnetite fine and coarse 

aggregate, a water-cement ratio of 0.63 (by weight), and had a 28-day compressive strength of 28.3 MPa and density 

of 3550 kg/m3. The specimens were cured in water at 20 ̊C or in air at 20 ̊C and 50% relative humidity and then 

heated in an oven to 110 ̊C.
234

Figure 3.2 Effect of elevated temperature on the compressive strength of 

a magnetite concrete tested at temperature and after cooling to room temperature.

Source: S. Ohgishi, S. Miyasaka, and J. Chida, ―On Properties of Magnetite and Serpentine Concrete at Elevated 

Temperatures for Nuclear Reactor,‖ SP 34-57 in Concrete for Nuclear Reactors, pp. 1243-1253, American 

Concrete Institute, Farmington Hills, Michigan, 1972.

Table 3.6 Effect of elevated temperature on the physical and 

mechanical properties of a magnetite aggregate concrete

Source: F. Du Bois, ―Beton Lourds abase de Mineral de Magnetite de Dielette,‖ Centre d’Etudes Nucleaires de 

Saclay, Commissariat a l’Energie Atomique, France, June 1964.

3.3.4 Haematite Aggregate-Based Concrete

Mechanical and physical properties of haemitite aggregate concretes are very similar to concrete made with 

magnetite [3.3]. Coefficients of thermal expansion for Portland cement concretes with haemitite aggregate at 

temperatures up to 300 ̊C, between 300 ̊C and 600 ̊C, and between 500 ̊ and 800 ̊C have been reported as 

5.94 x 10-6 ̊C-1, 11.5 x 10-6 ̊C-1, and 16.2 x 10-6 ̊C-1, respectively [3.28]. Residual compressive strength ratios for this 

concrete were 90, 85, and 150% at temperatures of 100 ̊, 150 ̊, and 300 ̊C, respectively, but when alumina cement 

was used the ratios were 80, 35, and 55%, respectively.
235

Investigations have been conducted related to use of haemitite concrete in the thermal shield of a liquid-sodiumcooled fast reactor in Italy [3.31]. The concrete mixes were made using pozzolanic cement or a proprietary cement 

consisting of a mixture of materials (e.g., Portland cement, plasticizer, expansive cement, and anti-bleeding agent). 

The pozzolanic concrete with 15-mm-maximum size haematite aggregate had a water-cement ratio of 0.49 (by 

weight), density of 3850 kg/m3, 28-day static modulus of elasticity of 37.3 GPa, and 28-day compressive strength of 

50.5 MPa. In order to investigate the effect of temperature on compressive strength, 15-cm concrete cube specimens 

were heated to either 200 ̊ or 300 ̊C and tested at temperature or permitted to cool to room temperature prior to 

testing. The pozzolanic concrete specimens heated to 200 ̊C and tested at temperature had a compressive strength 

that was 127% the reference room temperature strength, 138% the reference room temperature strength when tested 

after cooling to room temperature, and 117% the reference room temperature strength when heated to 300 ̊C and 

tested at temperature. Results obtained for the proprietary cement-based mixes ranged from 91 to 141%, 123 to 

145%, and 106 to 139% the reference room temperature compressive strengths for the three test conditions. For 

each of the concrete mixes investigated, residual compressive strengths from specimens permitted to cool to room 

temperature prior to testing were greater than results obtained from specimens tested at temperature. 

3.3.5 Ferrophosphorus Aggregate-Based Concrete

The effect of elevated temperature on mechanical and physical properties of conventionally-placed concrete made 

with ferrophosphorus aggregates was investigated as part of the study related to the N-Production Reactor located at 

Hanford, Washington [3.23]. The 4650 kg/m3concrete mix utilized Type II Portland cement, 19-mm maximum size 

ferrophosphorus aggregate, and had a water-cement ratio of 0.53 (by weight). Thermal expansion, thermal 

diffusivity, thermal conductivity, and specific heat values for the reference concrete were 10.35 x 10-6  ̊C-1, 

0.003 m2h

-1

, 2.914 W m-1 ̊C-1, and 0.17 cal g-1 ̊C-1, respectively. The effect of elevated temperature on mechanical 

and physical properties of the ferrophosphorus aggregate concrete is presented in Table 3.7. Results obtained from 

additional specimens that were soaked in water after being heated to 85 ̊, 200 ̊, or 350 ̊C indicated that the residual 

compressive strengths were 87, 74, and 76%, respectively. It was concluded that residual ratios for compressive 

strength, flexural strength, bond strength, and modulus of elasticity were comparable to those obtained from a 

companion ordinary concrete. 

Table 3.7 Effect of elevated temperature on the physical and 

mechanical properties of a ferrophosphorus aggregate concrete

Source: H.S. Davis, ―N-Reactor Shielding,‖ Paper SP-34-52 in Concrete for Nuclear Reactors, pp. 1109-1161, 

American Concrete Institute, Farmington Hills, Michigan, 1972.
236

3.3.6 Baryte Aggregate-Based Concrete

A baryte aggregate concrete has been studied under a EURATOM Program to investigate its potential application to 

fabrication of pressure vessels for nuclear reactors [3.32]. Results were compared to a normal weight limestone 

aggregate concrete. Reference compressive strength for the two concretes was about 65 MPa. Specimens were 

heated for a period of 28 days after being cured in water for a period of 28 to 56 days. Coefficients of thermal 

expansion for the baryte concrete were 13.4 x 10-6 ̊C-1, 13.5 x 10-6 ̊C-1, and 13.5 x 10-6 ̊C-1, for temperatures between 

50 ̊ and 150 ̊C, 150 ̊ and 250 ̊C, and 250 ̊ and 350 ̊C, respectively, which were greater than coefficients obtained for 

the limestone aggregate concrete. Thermal conductivity of the baryte concrete increased from 2.2 W m-1 ̊C-1to 2.3 

W m-1 ̊C-1 between 65 ̊ and 85 ̊C and then decreased to 1.4 W m-1 ̊C-1 at 180 ̊C. The residual compressive strength 

of the baryte concrete heated to 500 ̊C and then cooled to room temperature prior to testing was 83%, and after six 

heating cycles the ratio decreased to 50%. Residual values obtained for these conditions for the limestone aggregate 

concrete were 65 and 36%, respectively. The residual modulus of elasticity of the baryte concrete after heating to 

500 ̊C was 25% which was about the same as that for the limestone aggregate concrete after heating to 350 ̊C. 

Residual modulus of elasticity after six cycles for the baryte aggregate concrete was about the same as that obtained 

after one cycle. It was concluded that the baryte aggregate concrete retained properties after elevated temperature 

exposure better than the limestone aggregate concrete because the aggregate thermal expansion coefficient was more 

compatible with that of the mortar matrix.

The effect of elevated temperature on the physical and mechanical properties of baryte aggregate and quartz 

aggregate concretes has been investigated [3.33]. Mixes were prepared using either Type I Portland cement or blast 

furnace slag cement with a water-cement ratio of 0.6 (by weight). Reference room temperature compressive 

strengths for the limestone aggregate and baryte aggregate concretes were 49 and 52 MPa, respectively. Results 

were obtained for exposure periods of three hours at temperatures up to 750 ̊C. Figure 3.3 presents the effect of 

temperature on the residual ratios of compressive strength and weight loss for two cementitious and two aggregate 

materials. Reduction in compressive strength and weight loss of baryte aggregate concrete was less than that for

Figure 3.3 Weight loss and residual compressive strength after thermal treatment

for baryte and quartz aggregate concretes.

Source: H. Weigler and R. Fischer, ―Influence of High Temperatures on Strength and Deformations of Concrete,‖ 

Paper SP-34-26 in Concrete for Nuclear Reactors, pp. 481-493, American Concrete Institute, Farmington 

Hills, Michigan, 1972.
237

the quartz aggregate concrete. Compressive strength of specimens tested hot was generally less than that of 

specimens tested cold. Application of load to about 1/3 the reference compressive strength while heating resulted in 

no strength loss at elevated temperature for the baryte aggregate concrete while for limestone aggregate concrete the 

strength was slightly reduced. Thermal cycling to 600 ̊C did not result in a strength loss for the baryte aggregate 

concrete but produced a significant strength loss for the quartz aggregate concrete. 

The effect of different durations (1, 2, and 3 hours) and temperatures (250 ̊, 500 ̊, 750 ̊, and 950 ̊C) on the physical 

and mechanical properties of a baryte aggregate concrete was investigated [3.34]. The baryte aggregate concrete 

used Portland cement, had a 20-mm maximum aggregate size, water-cement ratio of 0.40 (by weight), and a density 

of 3250 kg/m3. Reference compressive strength, tensile strength, flexural strength, and modulus of elasticity values 

for the baryte aggregate concrete were 47 MPa, 2.85 MPa, 4.42 MPa, and 29.03 GPa, respectively. After two hours 

exposure to temperatures of 250 ̊, 500 ̊, 750 ̊, or 950 ̊C, the compressive strength at temperature was reduced 6.5, 

39.9, 68.1, and 85.6%, respectively, relative to the 47 MPa reference room-temperature strength. These reductions 

in compressive strength were less than that obtained for a normal weight gravel concrete. Radiation attenuation 

coefficient results for a Cobalt 60 source (1330keV) obtained for the baryte aggregate concrete indicate that the 

attenuation coefficients were reduced 9.5, 13.0, 17.0, and 21.0% after heating for two hours at 250 ̊, 500 ̊, 750 ̊, and 

950 ̊C, respectively. Similar reductions were obtained for a Cesium 127 (660 keV) source. The reductions were 

attributed to the effect of elevated temperature on concrete properties, especially the density due to loss of water.

The effect of elevated temperature on the compressive and flexural strength and static modulus of elasticity of a 

baryte aggregate concrete was investigated at the Saclay Nuclear Centre in France [3.35]. Specimens were heated 

for 28, 90, and 180 days at 150 ̊ or 250 ̊C. Table 3.8 summarizes the results and shows that the residual ratios 

decreased with increasing temperature and exposure period, probably due to cracking resulting from a mismatch in 

coefficients of thermal expansion between the matrix and aggregate.

Table 3.8 Effect of elevated temperature on mechanical properties of a baryte concrete

Source: Engineering Compendium on Radiation Shielding, Vol. II, S.9.1.12, p. 156, R.G. Jaeger (Editor), SpringerVerlag, Berlin, Germany, 1975.

3.3.7 Ilmenite Aggregate-Based Concrete

The effect of different durations (1, 2, and 3 hours) and temperatures (250 ̊, 500 ̊, 750 ̊, and 950 ̊C) on the physical 

and mechanical properties of an ilmenite aggregate concrete was investigated [3.34]. The ilmenite aggregate 

concrete used Portland cement, had a 20-mm maximum aggregate size, water-cement ratio of 0.40 (by weight), and 

a density of 3450 kg/m3. Reference compressive strength, tensile strength, flexural strength, and modulus of 

elasticity values for the baryte aggregate concrete were 51 MPa, 3.53 MPa, 5.4 MPa, and 34.43 GPa, respectively. 

After two hours exposure to temperatures of 250 ̊, 500 ̊, 750 ̊, and 950 ̊C, the compressive strength at temperature 

was reduced 6.7, 15.3, 41.3, and 73.7%, respectively, relative to the 51 MPa reference room-temperature strength. 

These reductions in compressive strength were less than that obtained from a normal weight gravel concrete and the 

baryte concrete reported earlier. Radiation attenuation coefficient results for a Cobalt 60 source (1330keV) obtained 

for the ilmenite aggregate concrete indicate that the attenuation coefficients were reduced 11.9, 14.0, 16.0, and 

18.0% after heating for two hours at 250 ̊, 500 ̊, 750 ̊, and 950 ̊C, respectively. Similar reductions were obtained for 
238

a Cesium 127 (660 keV) source. The reductions were attributed to the effect of elevated temperature on concrete 

properties, especially the density due to loss of water.

3.3.8 Iron/Steel Aggregate-Based Concrete

Average coefficients of thermal expansion for a concrete made with Portland cement, a water-cement ratio of 0.4 

(by weight), and steel-shot aggregate have been reported as 4.20 x 10-6 ̊C-1up to 300 ̊C, 8.5 x 10-6 ̊C-1between 

300 ̊ and 600 ̊C, and 16.20 x 10-6 ̊C-1between 600 ̊ and 800 ̊C [3.36]. Results for concrete using iron and steel 

scrap were 5.1 x 10-6 ̊C-1up to 300 ̊C, 7.2 x 10-6 ̊C-1between 300 ̊ and 600 ̊C, and 8.6 x 10-6 ̊C-1between 600 ̊ and 

800 ̊C. Residual ratios of compressive strength for iron-steel aggregate concrete cubes were 83, 113, 133, and 113% 

at temperatures of 100 ̊, 150 ̊, 200 ̊, and 300 ̊C, respectively.

3.4 Shielding Effectiveness

The effectiveness of concrete as a shield may be reduced under service conditions (elevated temperature) as drying 

reduces the hydrogen content or cracking occurs. Figure 3.4 presents results of elevated-temperature exposure on 

distribution of thermal and fast neutrons in ordinary concrete having a density of 2328 kg/m3[3.37]. These results 

demonstrate how an increase in temperature producing a decrease in concrete water content results in an increase in 

neutron flux density transmitted through a concrete of given thickness. 

Figure 3.4 Effect of elevated temperature on distribution of thermal and fast neutrons in ordinary concrete.

Source: E. G. Peterson, ―Shielding Properties of Ordinary Concretes as a Function of Temperature,‖ HW-65572, 

Hanford Atomic Products Operation, Richland, Washington, August 2, 1960.

The increase in neutron flux due to a known loss of water can be estimated for different concrete thicknesses [3.38]:

, where (3.1)

= increase in neutron flux,

= flux after water loss in neutrons cm-2s

-1

,

= flux before water loss in neutrons cm-2s

-1

, and

x = thickness of shield.
239

For the case when the water content is not known, the increase in neutron flux through an ordinary concrete shield 

can be estimated if the densities of concrete before and after water loss are known [3.38]:

, where (3.2)

= density of concrete in shield before water loss in lb ft-3,

= density of concrete in shield after water loss in lb ft-3, and

x = thickness of shield.

Estimates for heavyweight-aggregate concretes can be obtained by replacing the constant 0.0195 in the equation by 

0.0284, 0.024, 0.0208, 0.0227 for ferrophosphorus concrete, magnetite concrete, limonite concrete, and baryte 

concrete, respectively.

Gamma dose rates for a ferrophosphorus concrete (4821 kg/m3) as a function of shield thickness and temperature 

have been investigated [3.39]. Figure 3.5 presents gamma dose rate versus concrete thickness for this concrete 

tested in the unheated state and at 320 ̊C. Results indicate that for a shield thickness of 1.22 m the leakage of 

gamma rays at a concrete temperature of 320 ̊C is about 50 times greater than that for unheated concrete. Additional 

results indicated that the gamma leakage increased by factors of: 1.7 and 20 at temperatures of 100 ̊ and 320 ̊C, 

respectively, for 4231 kg/m3 limonite-steel aggregate concrete [3.40]; 1.2 and 1.8 at temperatures of 100 ̊ and 320 ̊C, 

respectively, for 4299 kg/m3iron-serpentine concrete [3.41]; and 1.3 at a temperature of 320 ̊C for 3524 kg/m3ironserpentine concrete [3.41].

Figure 3.5 Effect of elevated temperature on gamma dose rate distribution in a ferrophosphorus concrete.

Source: E. G. Peterson, ―Shielding Properties of Ferrophosphorus Concrete as a Function of Temperature,‖ 

HW-64774, Hanford Atomic Products Operation, Richland, Washington, 1960.

The effect of different durations (1, 2, and 3 h) of high temperature (250°, 500°, 750°, and 950°C) on the physical, 

mechanical, and radiation properties of heavyweight concrete has been studied [3.34]. Results showed that ilmenite 

concrete had the highest density and modulus of elasticity and lowest percent absorption, and it also had higher 

values of compressive, tensile, bending, and bond strength than that obtained from either the baryte or gravel 

concretes. Ilmenite also showed the highest attenuation of transmitted gamma rays and was most resistant to 

elevated temperature. As the magnitude of thermal exposure increased, the attenuation coefficient decreased.
240

In concrete structures cracks develop due to effects such as thermal stress and heat of hydration and as a result the 

shielding effectiveness of the concrete can be reduced, especially if through-cracks develop. The effect of straight 

and crooked cracks through a concrete shield has been investigated [3.42]. In the immediate vicinity of the concrete 

surface, leakage of gamma rays through a slit contributed significantly to -dose rate, but diminished rapidly with 

distance from the surface as a result of shield thickness and scattering effects. The shielding effectiveness of cracked 

concrete has been investigated and formulas were developed to define the resulting effects [3.43]. Guidelines 

developed to compensate for cracking note that it might be economically advantageous to allow a concrete shield to 

crack and then shield the resulting radiation by other means.

The effect of collinear crack width on -ray shielding of concrete has been investigated [3.44]. The concrete 

specimens were fabricated from the concrete mix used at the Wolsung nuclear power station (i.e., Type V cement, 

maximum aggregate size of 20 mm, and 30 MPa compressive strength). The test specimen is illustrated in 

Figure 3.6. Also shown in the figure is the experimental setup used for increasing the crack width from 0 to 2 mm in 

increments. The basic principal for evaluating the effect of crack width on -ray attenuation for concrete samples 

Figure 3.6 Test specimen schematic and experimental setup.

Source: C-M. Lee, Y.H. Lee, and K.J. Lee, ―Cracking Effect on Gamma-Ray Shielding Performance in Concrete 

Structure,‖ Progress in Nuclear Energy 49, pp. 303-312, 2007

that were either 10-, 15-, or 20-cm thick is shown in Figure 3.7. Results are also presented in the figure showing 

that the relative intensity, as determined by the ratio of the intensity of radiation transmitted to the intensity of 

radiation before attenuation, increased as the crack width increased and decreased as the specimen thickness 

increased. Results were utilized to develop a correlation between crack width and intensity for a collinear crack in 

concrete:

, where (3.3)

I = intensity of radiation,

I0 = intensity of radiation before attenuation,

= linear attenuation coefficient for medium, 

t = thickness of the medium,

x = crack width, and 

b = constant (i.e., 140-150).

The surface dose was found to increase logarithmically with an increase in crack width. It was concluded from the 

study that if the shielding thickness of the concrete structure exceeds 20 cm and the crack width is 0.4 mm, the effect 

on shielding effectiveness will be less than 10%. It should be noted that these results were for collinear cracks and 

in actual concrete structures crack surfaces are irregular.
241

Figure 3.7 Schematic diagram of experimental principle and correlation between crack width and intensity.

Source: C-M. Lee, Y.H. Lee, and K.J. Lee, ―Cracking Effect on Gamma-Ray Shielding Performance in Concrete 

Structure,‖ Progress in Nuclear Energy 49, pp. 303-312, 2007.

3.5 Summary

Results reported in the literature for properties of shielding concretes display wide differences because of factors 

such as mix proportions, the nature of the materials utilized, and the test conditions. The impact of elevated 

temperature on the water content of a shielding concrete has a significant effect on its nuclear properties, in 

particular its neutron radiation attenuation. Mechanical properties of shielding concretes tend to decrease with 

increasing temperature and exposure period. Thermal cycling can reduce the compressive strength, modulus of 

rupture, and static modulus of elasticity, with the reduction increasing as the peak cycling temperature increases. 

Specimens soaked in water prior to testing have lower residual compressive strengths than dry specimens. 

Specimens loaded while heated retain their residual compressive strength better than specimens that were not loaded 

during thermal treatment. Concrete density decreases with increasing temperature as evaporable water is removed, 

with the value becoming relatively stable after evaporable water is lost. As temperature increases (and water content 

decreases) the neutron flux density transmitted through concrete increases for a given section thickness. The 

intensity of radiation transmitted through cracked (collinear) concrete increases as the crack width increases and 

specimen thickness decreases.

3.6 References

3.1 H. Etherington, Nuclear Engineering Handbook, McGraw-Hill, New York, 109–111, December 1958. 

3.2 S. Glasstone and A. Sesonske, Nuclear Reactor Engineering, D. Van Nostrand, Princeton, New Jersey, 

pp. 614–615, December 1967.

3.3 M.F. Kaplan, Concrete Radiation Shielding, Concrete Design and Construction Series, Longman Scientific 

& Technical. New York, New York, 1989.

3.4 Standard Specification for Aggregates for Radiation-Shielding Concrete, ASTM C637-98a, American 

Society for Testing and Materials, West Conshockon, Pennsylvania, 2003.

3.5 Standard Descriptive Nomenclature of Constituents of Aggregates for Radiation-Shielding Concrete,

ASTM C638-82, American Society for Testing and Materials, West Conshockon, Pennsylvania, 2002.

3.6 Guide for Use of Normal Weight and Heavyweight Aggregate in Concrete, ACI 221R-96, American 

Concrete Institute, Farmington Hills, Michigan, 2001.
242

3.7 Standard Practice for Selecting Proportions for Normal, Heavyweight, and Mass Concrete, ACI 211.1-91, 

American Concrete Institute, Farmington Hills, Michigan, 2002.

3.8 Heavyweight Concrete: Measuring, Mixing, Transporting, and Placing, ACI 304.3R-96, American 

Concrete Institute, Farmington Hills, Michigan, 2004.

3.9 Nuclear Analysis and Design of Concrete Radiation Shielding for Nuclear Power Plants, ANSI/ANS-6.4-

2006, American Nuclear Society, La Grange Park, Illinois, 2006.

3.10 Specification for Radiation Shielding Materials, ANSI/ANS-6.4.2-2006, American Nuclear Society, La 

Grange Park, Illinois, 2006.

3.11 Specification for Aggregate from Natural Sources for Concrete,‖ BS 882, British Standards Institute, 

London, United Kingdom, 1992.

3.12 F. Du Bois, Specifications Techniques: Betons Tres Lourdes a Base de Riblons, de Fonte, et d’Un Mineral, 

Section de Physique de Materiaux Nucleaires, Commissarat a L’Energie Atomique, Saclay, France 

September 1965.

3.13 R. Walker and M. Grotenhuis, A Summary of Shielding Constants for Concrete, USAEC Report ANL6443, Argonne National Laboratory, November 1961.

3.14 A. Missenard, ―Theoretical and Experimental Research on the Heat Conductivity of Concrete,‖ Annnales 

de l’Institute Technique du Batiment et des Travaux Publics 18, pp. 211-212, July/August 1965.

3.15 M. Tourasse, ‖The Heavy Concretes – Physical Properties and Mechanical Tests,‖ Proceedings of Second 

United Nations Conference on the Peaceful Use of Atomic Energy, Paper 15/P/1152, 1958.

3.16 R.B. Gallaher and A.S. Kitzes, Summary Report on Portland Cement Concretes for Shielding, USAEC 

Report ORNL-1414, Oak Ridge National Laboratory, Oak Ridge, Tennessee, March 1953.

3.17 H.E. Hungerford, A. Hönig, A.E. Desov, F. Du Bois, and H.S. Davis, ―Concretes, Cements, Mortars, and 

Grouts,‖ Engineering Compendium on Radiation Shielding, Vol. II, S.9.1.12.15, p. 153, R.G. Jaeger 

(Editor), Springer-Verlag, Berlin, Germany, 1975.

3.18 Toledo Testing Laboratory, Inc., Reports 308907 and 308987 (November 1958) and Report 309587, 

February 1959.

3.19 H.E. Hungerford, A. Hönig, A.E. Desov, F. Du Bois, and H.S. Davis, ―Concretes, Cements, Mortars, and 

Grouts,‖ Engineering Compendium on Radiation Shielding, Vol. II, S.9.1.12, R.G. Jaeger (Editor), 

Springer-Verlag, Berlin, Germany, 1975.

3.20 H.S. Davis, F.L. Brown, and H.C, Winter, ―Properties of High-Density Concrete Made With Iron 

Aggregate,‖ Journal American Concrete Institute 27(7), Farmington Hills, Michigan, March 1956.

3.21 A.N. Komarovskii, Design of Nuclear Plants, Chapter 5, Translated from Russian by Israel Progress for 

Scientific Translation, Jerusalem, 1968.

3.22 G. Massa, R. De Stefano, M. Collepardi, S. Chatterji, and V. Maniscalco, ―Shielding Concretes for Liquid 

Sodium Cooled Nuclear Reactors,‖ Paper H3/8 in 7

th International Conference on Structural Mechanics in 

Reactor Technology, Chicago, Illinois, North-Holland Publishers, 1983.

3.23 H.S. Davis, ―N-Reactor Shielding,‖ Paper SP-34-52 in Concrete for Nuclear Reactors, pp. 1109-1161,

3.24 J.M. Raphael, ―The Structural Properties of Magnetite Concrete,‖ Journal of the Structural Division 84, 

Proceedings of the American Society of Civil Engineers, No. ST1, Paper 1511, New York, New York, 

January 1958.

3.25 H.E. Hungerford, A. Hönig, A.E. Desov, F. Du Bois, and H.S. Davis, ―Concretes, Cements, Mortars, and 

Grouts,‖ Engineering Compendium on Radiation Shielding, Vol. II, S.9.1.12.18, p. 162, R.G. Jaeger 

(Editor), Springer-Verlag, Berlin, Germany, 1975.

3.26 S. Ohgishi, S. Miyasaka, and J. Chida, ―On Properties of Magnetite and Serpentine Concrete at Elevated 

Temperatures for Nuclear Reactor,‖ SP 34-57 in Concrete for Nuclear Reactors, pp. 1243-1253,

3.27

Paper SP-34-23 in Concrete for Nuclear Reactors, pp. 423-434,

3.30 F. Du Bois, Beton Lourds abase de Mineral de Magnetite de Dielette, Centre d’Etudes Nucleaires de 

Saclay, Commissariat a l’Energie Atomique, France, June 1964.
243

3.31 G. Massa, R. De Stefano, M. Collepardi, S. Chatterji, and V. Maniscalco, ―Shielding Concretes for Liquid 

Sodium Cooled Nuclear Reactors,‖ Paper H3/8 in Transactions of the 7th International Conference on 

Structural Mechanics in Reactor Technology, pp. 151-158, North-Holland Publishers, 1983.

3.32 E. Crispino, ―Studies on the Technology of Concretes Under Thermal Conditions,‖ Paper SP-34-25 in 

Concrete for Nuclear Reactors, pp. 443-479,

3.33 H. Weigler and R. Fischer, ―Influence of High Temperatures on Strength and Deformations of Concrete,‖ 

Paper SP-34-26 in Concrete for Nuclear Reactors, pp. 481-493, American Concrete Institute, Farmington 

Hills, Michigan, 1972. 

3.34 K. Sakr and E. El-Hakim, ―Effect of High Temperature on Heavy Weight Concrete Properties,‖ Cement 

and Concrete Research 35, pp. 590-596, 2005.

3.35 Engineering Compendium on Radiation Shielding, Vol. II, S.9.1.12, p. 156, R.G. Jaeger (Editor), SpringerVerlag, Berlin, Germany, 1975.

3.36 A.N. Komarovskii, Design of Nuclear Plants, Chapter 7, Translated from Russian by Israel Progress for 

Scientific Translation, Jerusalem, 1968.

3.37 E. G. Peterson, Shielding Properties of Ordinary Concretes as a Function of Temperature, HW 65572, 

Hanford Atomic Products Operation, Richland, Washington, August 2, 1960.

3.38 H.E. Hungerford, ―Shielding,‖ in Fast Reactor Technology Plant Design, Chapter 8, J.G. Yevick Editor, 

MIT Press Cambridge, Massachusetts, 1966.

3.39 E. G. Peterson, Shielding Properties of Ferrophosphorus Concrete as a Function of Temperature, 

HW 64774, Hanford Atomic Products Operation, Richland, Washington, 1960.

3.40 W.L. Bunch, Attenuation Properties of High Density Portland Cement Concretes as a Function of 

Temperature, HW 54656, Hanford Atomic Products Operation, Richland, Washington, 1958.

3.41 E. G. Peterson, Shielding Properties of Ordinary Concretes as a Function of Temperature, HW 73255, 

Hanford Atomic Products Operation, Richland, Washington, 1962.

3.42 S. Miyasaka, Y. Kanemori, Y. Fukushima, T, Yamada, ―Gamma-Ray Leakage Through Slit in Concrete 

Shield,‖ Nippon Genshiryaku Gakkaishi 11, pp. 2–8, 1969.

3.43 F. Seboek, ―Shielding Effectiveness of Cracked Concrete,‖ Kerntechnik 12, pp. 4496–4501, November 

1970.

3.44 C-M. Lee, Y.H. Lee, and K.J. Lee, ―Cracking Effect on Gamma-Ray Shielding Performance in Concrete 

Structure,‖ Progress in Nuclear Energy 49, pp. 303-312, 2007.
244
245

4 EXAMPLES OF CODES AND STANDARDS THAT ADDRESS 

CONCRETE UNDER ELEVATED TEMPERATURE CONDITIONS

4.1 Current Practice

The compressive strength influences the load-carrying capacity of concrete components, and the stiffness (modulus 

of elasticity) impacts the structural deformations and loads that develop at restraints. Current requirements for the 

design of concrete containments are presented in American Society of Mechanical Engineer’s Section III—

Division 2, ―Code for Concrete Containments – Rules for Construction of Nuclear Facility Components.‖

Supplemental load combination criteria are presented in Sect. 3.8.1 of the NRC Standard Review Plan [4.1].

Table 4.1 presents ASME Code temperature limits for various conditions or load categories in a prestressed concrete 

reactor vessel (PCRV) for the appropriate conditions (normal operation and abnormal environment). As noted in 

this table, the temperature in the concrete should not exceed 65°C at the liner-concrete interface and in the bulk 

concrete. Between cooling tubes (near the liner), 93°C is given as the maximum allowable. Temperature limits for 

Table 4.1 Condition categories and temperature limits for concrete and prestressing systems for PCRVs

Load category Area Temperature limits,  ̊C

Construction Bulk concrete 54

Normal Liner

 Effective at liner-concrete interface

 Between cooling tubes

Bulk concrete

Bulk concrete with nuclear heating

Local hot spots

Distribution asymmetry

At prestressing tendons

Liner interface transients (twice daily) range

66

93

66

71

121

10

66a

38-66

Abnormal and severe environmental Liner

 Effective at liner-concrete interface

 Between cooling tubes

Bulk concrete

Local hot spots

Distribution asymmetry

At prestressing tendons

Liner interface transients range

93

132

93

191

38

79

38-83

Extreme environmental Liner

 Effective at liner-concrete interface

 Between cooling tubes

Bulk concrete

Local hot spots

Distribution asymmetry

At prestressing tendons

Liner interface transients range

149

204

154

260

38

149

38-93

Failure Bulk concrete

 Unpressurized condition

 Pressurized condition

204

316

aHigher temperatures may be permitted as long as the effects on material behavior (e.g., relaxation) are accounted 

for in design.

Source: ―Code for Concrete Reactor Vessels and Containments,‖ Nuclear Power Plant Components, ASME Boiler 

and Pressure Vessel Code, Section III, Division 2, American Society of Mechanical Engineers, New York, 

New York, July 2003.

 

ACI 359-01 is endorsed by U.S. Nuclear Regulatory Guide 1.136 (Rev. 3), ―Material for Concrete Containments,‖

March 2007.
246

concrete safety-related structures other than concrete reactor vessels and concrete containments are contained in 

Appendix E, ―Thermal Considerations,‖ of ACI 349, Code Requirements for Nuclear Safety-Related Structures and

Commentary.‖ The temperature limitations for: (1) normal operation or any other long-term period are that the 

concrete surface temperatures shall not exceed 65.6°C except for local areas, such as around penetrations, that are 

allowed to have increased temperatures not to exceed 93.3°C; (2) for accident or any other short-term period the 

temperatures shall not exceed 176.7°C for the surface, except, local areas are allowed to reach 343.3°C from steam 

or water jets in the event of a pipe failure, however, after exposure to these temperatures, the serviceability of the 

structure needs to be assessed before resuming the operation of the plant with the extent of the assessment 

determined by the engineer and it may be limited to visual inspection only; and (3) higher temperatures than those 

given in (1) and (2) above may be allowed if tests are provided to evaluate the reduction in concrete strength and this 

reduction is applied to design allowables and evidence is provided verifying that the increased temperatures do not 

cause deterioration of the concrete either with or without load.

The French specification for PCRVs [4.2] limits temperatures in active parts of the concrete to 90°C; the British 

specification [4.3] states that if the normal operating temperature of any section of the vessel structure is such that 

the failure strength of the concrete at that temperature is significantly less than at ambient temperature, this will be 

taken into account. The British specification further notes that most concrete mixes subjected to temperatures above 

100°C will suffer a reduction in compressive strength, and concrete with certain aggregates, particularly limestone, 

may suffer significant losses below that temperature. Figure 4.1 presents the BS 8110 design curves for strength 

reduction with temperature of unsealed dense concrete and lightweight aggregate concrete. Permissible 

temperatures for the concrete in PCRVs for gas-cooled reactors has generally been limited to the range of 45 to 

80°C [4.4]. 

Figure 4.1 BS 8110 design curves for strength variation with the temperature 

of dense concrete and lightweight concrete.

Source: Structural Use of Concrete. Code of Practice for Special Circumstances, Section Four, ―Fire Resistance,‖ 

BS 8110: Part 2, British Standards Institution, London, United Kingdom, 1985.

4.2 Other Code Provisions for Addressing Concrete at Elevated Temperature

The American Concrete Institute in ACI/TMS 216, ―Code Requirements for Determining the Fire Resistance of 

Concrete and Masonry Construction Assemblies,‖ provides guidance on reduction in concrete compressive strength 

in the zone of flexural behavior that is based on the elevated temperature and coarse aggregate type [4.5]. Figure 4.2 

presents information on the compressive strength of different aggregate concretes at elevated temperature and after 

cooling. Code guidance on the effect of elevated temperature on the strength of flexural reinforcement steel bar and 

strand is presented in Figure 4.3.
247

(a) Siliceous aggregate concrete.

(b) Carbonate aggregate

(c) Semi-lightweight aggregate concrete.

Figure 4.2 Compressive strength of different aggregate 

concretes at elevated temperature and after cooling

Source: Code Requirements for Determining the Fire Resistance of Concrete and Masonry Construction 

Assemblies, ACI 216.1M-07, American Concrete Institute, Farmington Hills, Michigan, 2007.
248

Figure 4.3 Strength of flexural steel bar and strand at elevated temperature.

Source: Code Requirements for Determining the Fire Resistance of Concrete and Masonry Construction 

Assemblies, ACI 216.1M-07, American Concrete Institute, Farmington Hills, Michigan, 2007.

National and international standards are available that provide guidance for computing concrete strength at elevated 

temperature in their fire design provisions: Comité Europeen de Normalisation (Eurocode 2—Part 1-2, Structural 

Fire Design, Eurocode 4—Part 1-2, ―General Rules for Structural Fire Design‖) [4.6], Comites Euro-International 

du Beton (CEB model code) Bulletin D’Information No. 208, Fire Design of Concrete Structures [4.7], and 

National Building Code of Finland’s RakMK B4 [4.8].

Eurocode 2 (CEN) specifies rules for the strength and deformation properties of uniaxially stressed concrete at 

elevated temperature. Figure 4.4 presents strength reduction factors for concrete with respect to elevated 

temperature. The Code makes no distinction between normal-strength and high-strength concretes in its fire design 

provisions. As the Code does not explicitly indicate whether the results are for concrete in service, (i.e., under load) 

it has been assumed that this is the case since the Code is for design of structural concrete [4.9,4.10].

Figure 4.4 CEN compressive strength reduction factor for concrete subjected to elevated temperature. 

Source: L.T. Phan, Fire Performance of High-Strength Concrete: A Report of the State-of-the-Art, NISTIR 5934, 

National Institute of Standards and Technology, Gaithersburg, Maryland, December 1966.
249

Comites Euro-International du Beton (CEB) has provided information on the effect of elevated temperature on 

concrete compressive strength, modulus of elasticity, and tensile strength. Figure 4.5 presents the recommended 

design curve for compressive strength of siliceous normal weight concrete at high temperature. Also shown in the 

figure is a comparison to the design curve of results obtained by several investigators for concretes having a 

compressive strength of < 50MPa. The design curve for compressive strength of lightweight concrete at high 

temperature is presented in Figure 4.6. Design curves for the effect of elevated temperature on concrete modulus of 

elasticity and tensile strength are presented in Figures 4.7 and 4.8, respectively. The CEB Code also makes no 

distinction between normal-strength and high-strength concretes in its fire design provisions and does not indicate 

whether or not the results are for concrete in service (i.e., under load), but it has been assumed that this is the case 

[4.9,4.10].

Figure 4.5 CEB design curve for compressive strength of 

siliceous normal weight concrete subjected to elevated temperature. 

Source: L.T. Phan, Fire Performance of High-Strength Concrete: A Report of the State-of-the-Art, NISTIR 5934, 

National Institute of Standards and Technology, Gaithersburg, Maryland, December 1966.

Figure 4.6 CEB design curve for compressive strength of 

lightweight concrete subjected to elevated temperature. 

Source: L.T. Phan, Fire Performance of High-Strength Concrete: A Report of the State-of-the-Art, NISTIR 5934, 

National Institute of Standards and Technology, Gaithersburg, Maryland, December 1966.
250

Figure 4.7 CEB design curve for effect of elevated temperature on 

modulus of elasticity of lightweight and normal-strength concretes. 

Source: L.T. Phan, Fire Performance of High-Strength Concrete: A Report of the State-of-the-Art, NISTIR 5934, 

National Institute of Standards and Technology, Gaithersburg, Maryland, December 1966.

Figure 4.8 CEB design curve for effect of elevated temperature on tensile strength. 

Source: L.T. Phan, Fire Performance of High-Strength Concrete: A Report of the State-of-the-Art, NISTIR 5934, 

National Institute of Standards and Technology, Gaithersburg, Maryland, December 1966.

Finland’s RakMK B4 prescribes different design rules for high-strength (70 MPa < fc’ < 100 MPa, 150-mm cube 

strength), and normal-strength (10 MPa < fc’ < 70 MPa, 150-mm cube strength) concretes. The RakMK B4 also 

prescribes different design rules for concrete in service (stressed, 0.3f23 ̊C) and for concrete that is not (unstressed). 

[4.9,4.10]. Figure 4.9 presents the RakMk B4 design code for effect of elevated temperature on concrete relative 

compressive strength. The RakMK B4’s strength prediction for unstressed normal-strength concrete prescribes a 
251

range of no strength loss between room temperature and 220 oC. Above 220 oC, RakMK B4 prescribes a strength 

loss that has a similar rate of strength reduction with temperature as in the case for high-strength concrete. A 

summary comparison of the design codes for compressive strength is presented in Figure 4.10.

Figure 4.9 RakMK B4 design curves for effect of elevated temperature on concrete compressive strength. 

Source: L.T. Phan, Fire Performance of High-Strength Concrete: A Report of the State-of-the-Art, NISTIR 5934, 

National Institute of Standards and Technology, Gaithersburg, Maryland, December 1966.

Figure 4.10 Comparison of design curves for effect of elevated temperature on concrete compressive strength. 

 

Reference [4.11] provides the source of information on the French DTU design code for high-strength concrete.
252

4.3 References

4.1 U.S. Atomic Energy Commission, ―Concrete Containment,‖ Sect. 3.8.1 in Regulatory Standard Review 

Plan, NUREG-0800, Washington, D.C., July 1981. 

4.2 ―Nuclear Reactor Vessels of Prestressed Concrete with Metal Reinforcements,‖ J. Offic. Republ. Fr., 

pp. 6119 – 6128, June 1970.

4.3 Specification for Prestressed Concrete Pressure Vessels for Nuclear Reactors, BS 4975, British Standards 

Institution, London, United Kingdom, July 1973

4.4 Structural Use of Concrete. Code of Practice for Special Circumstances, Section Four, ―Fire Resistance,‖ 

BS 8110: Part 2, British Standards Institution, London, United Kingdom, 1985.

4.5 Code Requirements for Determining the Fire Resistance of Concrete and Masonry Construction 

Assemblies, ACI 216.1M-07, American Concrete Institute, Farmington Hills, Michigan, 2007.

4.6 Comité Européen de Normalisation, ―prENV 1992-1-2: Eurocode 2: Design of Concrete Structures. Part 1 -

2: Structural Fire Design,‖ CEN/TC 250/SC 2, 1993.

4.7 Comites Euro-International Du Beton, ―Fire Design of Concrete Structures - In Accordance with CEB/FIP 

Model Code 90,‖ CEB Bulletin D'Information No. 208, Lausanne, Switzerland, 1991.

4.8 Concrete Association of Finland, High Strength Concrete Supplementary Rules and Fire Design, 

RakMK B4, 1991.

4.9 L.T. Phan, Fire Performance of High-Strength Concrete: A Report of the State-of-the-Art, NISTIR 5934, 

National Institute of Standards and Technology, Gaithersburg, Maryland, December 1966.

4.10 L.T. Phan, ―High-Strength Concrete at High Temperature – An Overview,‖ Proceedings of 6th 

International Symposium on Utilization of High Strength/High Performance Concrete, Volume 1, pp. 501-

518, Leipzig, Germany, June 2002.

4.11 B. Persson, Self-Compacting Concrete at Fire Temperatures, TVBM-03/3110, Division of Building 

Materials, Lund Institute of Technology, Lund University, Sweden, September 2003.
253

5 POTENTIAL METHODS FOR ASSESSMENT OF

CONCRETE EXPOSED TO HIGH TEMPERATURES

5.1 Introduction

Thermal loading of reinforced concrete can result in damage ranging from cosmetic blemishes to more serious 

effects (e.g., misalignment and distortions). Heating of concrete may result in a variety of structural changes such as 

cracking, spalling, debonding of aggregate, expansion, and mineralogical or chemical changes such as discoloration, 

dehydration, and dissociation. With respect to the cement paste, evaporation and dissolution, dehydration and 

dissociation of ettringite, gypsum, calcium hydroxide, calcium carbonate and other phases such as calcium silicate 

hydrates may occur [5.1]. 

Evaluation of the heating history of a structure is important and has three primary purposes: (1) assess the thermal 

(i.e., fire) resistance of a particular element of a major structure, (2) assist in forensic research of the cause of a 

thermal excursion by determining the duration of exposure of concrete to elevated temperature, and (3) determine 

whether the structure is still structurally sound and to assess items such as whether the steel reinforcement has been 

exposed to high enough temperature to effect its properties [5.2]. One approach for performing an assessment of a 

reinforced concrete structure following a thermal event is presented in Figure 5.1. A listing of possible approaches 

for nondestructive assessment of fire-damaged concrete structures is provided in Table 5.1. Information on selected 

visual assessment approaches, and field and laboratory testing methods is provided below. 

Figure 5.1 Example of evaluation process for structure subjected to thermal excursion such as resulting from a fire. 

Source: N.K. Gosain, R.F. Drexler, and P. Choudhuri, ―Evaluation and Repair of Fire-Damaged Buildings,‖ 

Structure, 14 pages, September 2008.

 
254

Table 5.1 Listing of possible nondestructive approaches for use in assessment of fire-damaged concrete

Average response of cover 

concrete

Point-by-point response of small samples Special interpretation techniques

• Hammer tapping

• Schmidt rebound hammer

• Windsor probe

• Capo test

• BRE internal fracture

• Ultrasonic pulse velocity

• Small-scale mechanical testing

• Differential thermal analysis

• Thermogravimetric analysis (TGA)

• Dilatometry (TMA)

• Thermoluminescense

• Porosimetry

• Microcrack density analysis

• Colorimetry

• Petrographic analysis

• Chemical analysis

• UPV method

• Impact echo

• Sonic tomography

• Modal analysis of surface waves

• Ground-penetrating radar

• Electrical resistivity

Source: M. Colombo and R. Felicetti, ―New NDT Techniques for the Assessment of Fire-Damaged Concrete 

Structures,‖ Fire Safety Journal 42, pp. 461-472, 2007.

5.2 Visual Assessment

Information related to the maximum temperature experienced and an assessment of the condition of a concrete 

structure following a thermal excursion generally starts with an on-site condition assessment. The first phase of the 

on-site assessment involves a general visual inspection of the structures for signs of distress (e.g., cracking, spalling, 

deflections, distortions, misalignment, and exposed steel reinforcement). Table 5.2 provides simplified general 

guidance for visual concrete fire damage classification [5.5]. Table 5.3 provides more detailed guidance that has 

been developed to assist in assessing and categorizing damage of individual concrete members [5.6]. Steel 

reinforcement in flexure members that has not experienced severe distortion usually indicates that the steel has not 

suffered a significant reduction in yield strength due to thermal exposure, and if concrete spalling does not extend to 

the depth of steel reinforcement the structural strength is relatively unaffected [5.7]. Columns that contain numerous 

ties or spiral confinement, however, can reach temperatures high enough to reduce the yield strength without 

showing signs of severe distortion or buckling. Shear failure of normal weight concrete beams exposed to fire 

conditions are rare and flexural cracks can form in the negative moment regions over supports to redistribute 

moments as a result of fire exposure [5.7]. If cracks are not present in the negative moment region, the fire was not

Table 5.2 Simplified visual concrete fire damage classification

Class of damage

Features observed

Color Crazing Spalling Reinforcement Cracks Deflection

0

(Decoration req’d)

Normal None None None exposed None None

1

(Superficial repair req’d)

Normal Slight Minor None exposed None None

2

(General repair req’d)

Pink Moderate Localized

Up to 25% 

exposed

None None

3

(Principal repair req’d)

Whitish 

gray

Extensive Considerable

Up to 50% 

exposed

Minor None

4

(Major repair req’d)

Buff Surface lost Almost total

Up to 50% 

exposed

Major Distorted

Source: Assessment and Repair of Fire-Damaged Concrete Structures, Technical Report No. 33, Concrete Society, 

Camberley, United Kingdom, 1990.

 

General guidance on conduct of a condition assessment is available [5.3,5.4].
255

Table 5.3 Initial assessment of damage and probable treatment required

Damage Class 1 Damage Class 2 Damage Class 3 Damage Class 4

COLUMNS

Soot and smoke

deposits Present - - -

Color change - Pink to buff surface Buff surface -

Spalling - Only minor Local Extensive

Steel exposure - - Steel showing Considerable areas

Surface separation Peeling Substantial

Surface mostly

gone. Remainder

sounds hollow

when struck

-

Number of main

bars buckled - - Not more than one One or more

Microcracking - Extensive - -

Distortion - - Possible -

SLABS

Reinforced concrete 

solid slabs

 Soot and smoke

 deposits

Present General coverage Completely covered, 

or color changed -

 Color change - - Pink -

 Spalling Minor Present Present -

 Steel exposure - 10% or less Over 10% -

 Adherence of

 steel to concrete - Adhering Adhering Fallen clear

Plank Some broken Substantial damage - -

Ribs Intact - - -

 Soot and smoke

 deposits Present - - -

 Spalling None Present Extensive -

 Steel exposure - Small areas - -

 Adherence of

 steel to concrete - - Generally adhering Fallen clear

Suspended ceiling Extensive damage - - -

Deflection - - Not severe Substantial

BEAMS

Soot and smoke

deposits Present Completely covered or color changed - -

Color change - Pink Buff Buff to gray

Spalling Minor Substantial but at 

edges only Substantial on soffit sides Extensive on soffit sides

Steel exposure Little or none Outer edges of corner 

bars

Main bars each about

50%

Almost al lower 

main bars

Surface separation -

Cover concrete of 

soffit sounds hollow

when struck

- -

Number of main

bars buckled - - Not more than one Possibly several

Microcracking - Surface - -

Cracking - -

Several cracks 

~6.4 mm -

Deflection or 

Fracture - - Not severe

Substantial 

deflection or

fracture or both

COLUMNS, SLABS, 

BEAMS

Probable treatment

required Cosmetic only Some replacement

Examination in 

greater detail. 

Considerable 

replacement or 

reclassification as 

Class 2 or 4

Removal and 

replacement or 

strengthening 

extensively with 

additional 

concrete and 

reinforcement

Source: ―Concrete Structures After Fire,‖ Concrete Construction 17(3), Addison, Illinois, 1972.
256

significant enough for the concrete’s cracking moment strength to be reached and therefore the strength of the steel 

reinforcement can be considered to be intact. Fire exposure of prestressed concrete members requires a more 

detailed evaluation as the elevated temperature can result in strength loss, permanent relaxation losses, and a 

reduction in prestressing force [5.8]. Increased deflection of prestressed concrete members would be an indication 

of exposure to high temperatures. Critical temperatures for cold-worked steel reinforcement and hot-rolled steel 

reinforcement at which considerable residual strength is lost have been cited as > 450 ̊C and > 600 ̊C, respectively, 

and prestressing steel exhibits considerable strength loss at 200 ̊ to 400 ̊C [5.9] Masonry can exhibit distress 

similar to that of concrete (e.g., cracks, pitting of aggregates, shallow spalling and minor surface damage). Masonry 

under fire exposure can exhibit some softening of the mortar joints, but the damage is typically confined to within 

about 19 mm of the fire-exposed wall and excellent structural fire endurance characteristics are maintained 

[5.7,5.10].

In addition to spalling, loss of chemically combined water, and loss of strength, Portland cement concretes can 

exhibit changes in color as a result of a sudden thermal excursion such as resulting from a fire. Crazing, cracking, 

popouts caused by quartz or chert aggregate particles, spalling, and dehydration (crumbling and powder of cement 

paste) provide general indications of temperatures to which the concrete was exposed. Approximate corresponding 

temperatures for these occurrences are 300 ̊, 550 ̊ (deep cracking), 575 ̊, 800 ̊, and 900 ̊C, respectively [5.12]. 

Heating can change the color (hue) of concrete that in turn can sometimes provide an indication of the temperature 

attained and an equivalent fire duration [5.9]. An example of this is presented in Figure 5.2 which shows the 

presence of cracks in a region of pink discoloration and a partial change of aggregate color from yellow gold to red 

in upper right of figure. 

Figure 5.2 Discoloration and cracking in concrete resulting from elevated temperature exposure.

Source: Assessing the Condition and Repair Alternatives of Fire-Exposed Concrete and Masonry Members, Report 

SR332.01B, National Codes and Standards Council of the Concrete and Masonry Industries, Skokie, 

Illinois, August 1994.

Concrete made with sedimentary or metamorphic aggregates can show permanent color change on heating [5.13]. 

Color is normal to 230 ̊C; goes from pink to red from 290 ̊ to 590 ̊C; and from 590 ̊ to 900 ̊C color changes to gray 

and then buff. For temperatures up to about 500 ̊C temperature distribution is little affected by using carbonate 

rather than siliceous aggregate. Actual concrete colors observed depend on aggregate type present and are most 

pronounced for siliceous aggregate and less so for limestone, granite, and Lytag [5.9]. Flint (chert) exhibits the most 

striking colors as illustrated in Figure 5.3. Colorimeters have been applied directly to the surface of concrete 

samples as part of the assessment of the Mont Blanc Tunnel fire and the measurements were found to correlate well 

with the maximum temperature experienced [5.14].

 

Appendix B presents information on effects of elevated temperature on properties of steel reinforcement and prestressing steels, 

Chapter 2 discusses effects on concrete’s physical and mechanical properties, and spalling has been addressed elsewhere [5.11].

Additional information on color change and temperature is presented in Section 5.4.
257

Figure 5.3 Appearance of flint-aggregate concrete after heating.

Source: J.P. Ingham, ―Assessment of Fire-Damaged Concrete and Masonry Structures: Application of 

Petrography,‖ 11th Euroseminar on Microscopy Applied to Building Materials, 16 pages, Porto, Portugal, 

5-9 June 2007.

An overview of the changes in coloration of concrete resulting from temperature exposure has been compiled [5.15]. 

Presented below is material extracted from this compilation and references cited can be obtained from the main 

reference. 

―At room temperature, concrete is reported as yellowish [1]. A faint pink hue is observed beginning at either 

232° C [90; 96; 98] or at 250° C [95]. This has been attributed to changes in the limestone [98] and/or the 

iron oxide [95]. Between 250° C and 300° C, a distinct pink hue is reported [1; 69; 74; 90; 95; 96; 98-100].19 

The distinct pink is attributed to iron compounds, sand, and sandstone [1]; to ferric oxide [95]20; to siliaceous 

aggregates [74]19; and to ferrous salts [100]. The pink color at 300° C changes to a brick red at about 600° C. 

This has been observed by numerous researchers [74; 90; 95; 98; 102].19 Again, the color change is mostly 

attributed to limestone [69; 98] and to ferric oxide [95; 102]. The brick red color changes to gray at higher 

temperatures. This gray color has been reported in the range between 600° C and 900° C [98] and at 900° C 

[74]19. Smith [90] found an additional color change, from brick red to black to gray, in the 600° C to 900° C 

color range. The final color change is to buff. This color has been reported at approximately 1000° C [74]19

and at about 900° C [90; 98]. These color changes are still largely attributed to the iron oxide [95],[102].‖ 

These color changes and their rough temperature ranges are loosely summarized in Table 5.4. Additional 

information is presented in Table 5.5 covering change in appearance of concrete exposed to very high temperatures 

(T > 800 ̊C) [5.16]. The samples in this study were obtained from two structures in Taiwan and subjected to the 

temperatures noted in the table for either 10 minutes or one hour and then cooled to room temperature and 

examined. Results of thermogravimetric analysis, x-ray diffraction, and scanning electron microscopy examinations 

are also available in the reference. It was noted that 1100 ̊ to 1200 ̊C was considered to be critical in that exposure 

to temperatures below 1000 ̊C resulted in light color, loose and friable structure, and cracks when cooled to ambient, 

while exposure to temperatures greater than 1000 ̊C tended to result in a dark color with glossy appearance, hard 

structure, and the disappearance of the cracks. Concrete apparently liquefied at temperatures of 1300 ̊ or 1400 ̊C.
258

Table 5.4 Indication of concrete color change on heating

Temp.,  ̊C 20 ̊ 232 ̊ - 250 ̊ 300 ̊ 300 ̊ - 600 ̊ 600 ̊ 600 ̊ - 900 ̊ 900 ̊ 900 ̊ 1000 ̊

concrete 

color yellow faint pink distinct pink

pink changing 

to brick red

brick 

red black gray gray buff buff

Source: R.K. Schroeder, Post-Fire Analysis of Construction Materials, PhD Thesis, University of California, 

Berkeley, 1999.

Table 5.5 Appearance of concrete after being subjecting to very high temperatures

Thermal exposure 

condition

Appearance

Sample A Sample B Sample B-1

800 ̊C

10 min Pale red, loose, has small 

cracks

Gray, loose, has small 

cracks

Dark brown, glossy, 

hard (remained 

unchanged)

1 hour Pale red, loose, has small 

cracks

Gray, loose, has small 

cracks -

900 ̊C 10 min Pale red, loose, has small 

cracks

Gray, loose, cracks slightly 

enlarged Remained unchanged

1 hour Pale red, loose, cracks 

slightly enlarged

Gray, loose, cracks slightly 

enlarged -

1000 ̊C 10 min Light red, loose, cracks 

slightly enlarged

Gray, loose, cracks 

enlarged Remained unchanged

1 hour

Light red, loose, expanded 

volume with large cracks

Gray, loose, expanded 

volume with large cracks -

1100 ̊C

10 min Gray, loose, expanded 

volume with large cracks

Gray, loose, fragile, 

cracked

Corners become 

smooth

1 hour Gray, loose, expanded 

volume with large cracks

Brownish yellow, loose, 

fragile, cracked -

1200 ̊C

10 min Grayish yellow, hard, 

glossy, has small cracks

Dark brown, hard, glossy, 

flat

Dark brown, has 

smooth surface, hard, 

flat

1 hour

Partially grayish yellow and 

partially brown, hard, 

glossy, no cracks

Dark brown, hard, glossy, 

flat, partly molten while 

hot

-

1300 ̊C 10 min Brown, hard, glossy, flat, 

molten while hot

Dark brown, hard, glossy, 

flat, molten while hot

Dark brown, glossy, 

hard, liquid-like while 

hot

1 hour Brown, hard, glossy, flat, 

molten while hot

Dark brown, hard, glossy, 

flat, liquid-like while hot -

1400 ̊C

10 min Dark brown, glossy, hard, 

liquid-like while hot -

Dark brown, glossy, 

hard, liquid-like while 

hot

1 hour Dark brown, glossy, hard, 

liquid-like while hot - -

Source: W-T. Chang, Y-S. Giang, C-T. Wang, and C-W. Huang, ―Concrete at High Temperatures Above 1000 ̊C,‖ 

1993 International Carnahan Conference on Security Technology, Institute of Electrical and Electronics 

Engineers, Ottawa, Ontario, Canada, October 1993.
259

The transition of siliceous aggregate from normal to pink or red at 300 ̊to 600 ̊C is considered to be useful 

in damage assessments in that this is the temperature range at which significant loss of concrete strength 

occurs [5.17]. The red color change is due to presence of oxidizable iron. The color change from pink to 

red is less prominent with calcareous and igneous aggregate. Not all aggregates will undergo color changes 

on heating and consideration should be given to the possibility that the pink/red color may be a feature of 

the aggregate rather than heat-induced [5.9]. Also surface temperatures usually are quite different from 

temperatures at different depths in the concrete and thus estimates of strength based on surface color 

indications of temperature are speculative [5.12].

Visual methods are often complemented by tapping tests with a metallic object such as a hammer or chisel 

to detect hollow-sounding, delaminated material. Additional information that can be revealed from the 

tapping tests includes hardness, integrity, depth of damage, seriousness of cracking, and condition of steel 

[5.12].

5.3 Field-Testing Techniques

A number of non-destructive techniques can be used to provide a more detailed assessment of the material in-situ 

(e.g., see Table 5.1). Techniques typically utilized include rebound hammer, break-off method, impact echo, and 

ultrasonics. ACI 364.1R-07, Guide for Evaluation of Concrete Structures Before Rehabilitation, in Tables 5.1 and 

5.2 provides a listing of test methods to evaluate hardened concrete in existing structures and test methods to 

determine structural properties and determine the condition of reinforcing steel, respectively. Descriptions and an 

overview of methods for detection of degradation in reinforced concrete structural members are also available 

[5.18,5.19]. Discussed below are two relatively recent field-testing techniques that have been proposed for 

assessment of fire-damaged concrete.

Indirect ultrasonic pulse velocity measurements based on the refraction of longitudinal ultrasonic waves has been 

investigated [5.20]. The measurement of pulse arrival time is performed by placing the emitting and receiving 

probes on the same face of the structure. As the maximum depth of the material investigated is a function of the 

distance between the transmitting and receiving probes, by increasing the separation the structure can be 

investigated to deeper depths. A plot of distance versus pulse arrival time is developed that is interpreted using 

numerical methods. Figure 5.4 presents an illustration of application of the method to assessment of damage 

thickness in a wall. A 20% decrease in the velocity was selected as the threshold for damage in this study. It was 

noted in the reference that application of this method can be tedious, and a flat surface is required. 

Figure 5.4 Illustration of application of indirect UPV method to identify damage depth (intercept at X = 0 in middle 

figure is related to thickness of the sizably damaged concrete as shown in right-hand side of figure)

Source: M. Colombo and R. Felicetti, ―New NDT Techniques for the Assessment of Fire-Damaged Concrete 

Structures,‖ Fire Safety Journal 42, pp. 461-472, 2007.

Measurement of drilling resistance has been proposed as a method to ascertain the thermal damage in concrete 

structures resulting from a fire [5.21]. The approach is based on measurement of either the thrust to be exerted to 
260

drill the material at constant feed rate or the work to drill a unit deep hole (J/mm). A common battery hammer drill 

that has been modified to monitor power consumption, the bit rotation, and hole depth is utilized. Electrical signals 

acquired are fed to a laptop computer and processed by dedicated software so the results can be displayed in real 

time. Figure 5.5 presents an example of results that were obtained from a column that had been severely damaged 

by fire. The most severely damaged areas were initially selected based on results from rebound hammer tests. Since 

the rebound hammer results could not indicate the depth of damage, drilling resistance tests were performed to 

establish a damage depth of slightly over 20 mm. A significant advantage noted for this method is that specific 

calibration curves do not have to be generated for comparison because the results are compared to the material 

response of the inner undamaged layer of the member under investigation.

Figure 5.5 Illustration of application of drilling resistance method to fire-damaged concrete column.

Source: M. Colombo and R. Felicetti, ―New NDT Techniques for the Assessment of Fire-Damaged Concrete 

Structures,‖ Fire Safety Journal 42, pp. 461-472, 2007.

5.4 Laboratory Techniques

Removal and testing of concrete cores samples in the laboratory provides the most effective approach for obtaining 

the comprehensive and detailed information on the mechanical properties and internal condition of concrete 

following a thermal excursion. Guidance on sampling and material testing is available [5.3,5.4].

Ultrasonic pulse velocity or resonant frequency testing have been noted as a good tool for indicating damage to 

concrete resulting from elevated temperature exposure, but are not reliable indicators of the concrete compressive 

strength [5.22]. Examples of other potential lab test methods include thermoluminescence, scanning electron 

microscopy, x-ray diffraction, differential thermal analysis, thermogravimetric analysis, and derivative 

thermogravimetric analysis. Figure 5.6 presents an example of use of scanning electron microscopy to examine the 

change in morphologies of cement hydrates resulting from temperature exposure. It has been noted that these 

potential lab test methods tend to be somewhat qualitative in nature or have primarily been used in academic 

research and are not routinely used to investigate thermally-damaged structures [5.9,5.23]. However, concrete 

petrography provides an approach for examining concrete to characterize and determine any damage caused as a 

result of a thermal exposure. 

Petrographic analysis of extracted concrete core samples can provide information on bond loss between the cement 

matrix and steel reinforcement, crack orientation and its relationship to the aggregate, microcracking, extent of 

cement hydration, chemical compositional changes of cementitious materials and aggregates, and temperature 

distribution within a given concrete depth [5.7]. Petrography is the primary tool for assessing the internal condition 

of concrete and can be used to determine the depth of damage for reinforced concrete as well as providing a means
261

Figure 5.6 Effect of elevated temperature exposure on morphologies of a

siliceous aggregate Type I Portland cement concrete.

Source: W-M. Lin, T.D. Lin, and L.J. Powere-Couche, ―Microstructures of Fire-Damaged Concrete,‖ ACI 

Materials Journal 93(3), American Concrete Institute, Farmington Hills, Michigan, September-October 

1996. 

to determine whether the damage was caused by thermal exposure or some other mechanism. Detailed information 

on practices for petrographic examination of hardened concrete is available [5.13,5.24]. Two areas related to 

petrography that can be utilized in assessing the significance of damage resulting from thermal exposure are 

examinations of phase changes and microcracking occurrence. 

Changes in both the cement paste and the aggregate result from elevated temperature exposure and can result in 

conversion of certain phases into new phases that may alter the concrete color (discussed previously) and the 

original mineralogical composition of the cement paste so that lines of equal physico-chemical conditions (isograds) 

can be mapped [5.2]. Since the isograds occur as a result of temperature, they can be used to provide an indication 

of temperature variations with concrete depth. Table 5.6 provides an indication of results that might be derived by 

examining the concrete with a hand lens or stereomicroscope. More detailed information (e.g., disappearance or 

Table 5.6 Visual indications of elevated temperature effects on concrete

Temperature,  ̊C Indicators

< 300 ̊ Normal, no apparent macroscopic changes in concrete; color remains gray

300 ̊ - 350 ̊

Oxidation of iron hydroxides like FeOOH in aggregate and cement paste to hematite, -Fe2O3, 

causing a permanent change of color of the concrete from gray to pinkish brown. This isograd is of 

importance as concrete and steel properties will undergo loss of properties at higher temperatures.

573 ̊

Transition of -quartz to -quartz, accompanied by an instantaneous increase in volume of quartz of 

about 5%, resulting in a radial cracking pattern around the quartz grains in the aggregate; this phase 

transition itself is reversible, but the radial cracking provides a diagnostic feature that remains after 

cooling

> 800 ̊ Complete disintegration of calcareous constituents of the aggregate and cement paste due to both 

dissociation and extreme thermal stresses causing a whitish gray coloration of the concrete

Source: T.G. Nijland and J.A. Larbi, ―Unraveling the Temperature Distribution in Fire-Damaged Concrete by 

Means of PFM Microscopy: Outline of Approach and Review of Useful Readings,‖ HERON 46(4), pp. 

253-264, 2001.
262

appearance of cement phases, and secondary phases) can be provided by polarizing and fluorescent microscopy 

which examines fluorescent thin sections with the aid of a combined polarizing and fluorescent microscope [5.2]. 

Careful examination of microscopically-observed features allows thermal contours to be plotted through the depth of 

individual concrete members [i.e., under favorable conditions contours can be plotted for 105 ̊C (increased porosity 

of cement matrix), 300 ̊C (red discoloration of aggregate), 500 ̊C (cement matrix becomes completely isotropic), 

600 ̊C ( to quartz transition), 800 ̊C (calcination of limestone), and 1200 ̊C (first signs of melting)] [5.9]. By 

conducting a point by point examination of concrete material constituents and outlines of color profiles using the 

combination of an optical microscope and a workstation to provide optical magnification and image resolution, it 

has been shown that locations significantly affected by high temperature can be identified through changes in hue 

[5.17].

It has been shown that the crack density, residual concrete compressive strength, and temperature can be correlated 

(e.g., crack density/color change can be used to indicate depth to which compressive strength is likely to have been 

significantly affected) [5.23]. In this study specimens were fabricated using ordinary Portland cements (OPCs) and 

blended cements (OPC and 30% fly ash, OPC and 50% blast furnace slag) in conjunction with quartz sand and 

siliceous gravel. Additional specimens were fabricated using OPC, sand, and either crushed limestone, granite, or 

Lytag coarse aggregates. After moist curing for 28 days and 32 days air drying, the specimens were heated at 105 ̊C 

for 48 hours to provide well-cured specimens and prevent spalling. The beam, cube, and cylindrical test specimens 

were then heated at temperatures ranging from 175 ̊ to 700 ̊C for one hour and then permitted to cool to room 

temperature. After elevated temperature exposure, 10-mm-thick sections were cut from the center of the cube 

specimens, vacuum impregnated with low viscosity resin, glued to a slide, polished, and examined using thin-section 

petrography with the results photographed. Crack densities were then determined in terms of mm of crack length 

per cm2of concrete. Figure 5.7 presents the effect of temperature on crack density for specimens made with 

siliceous aggregates and OPC and can be used to indicate the temperature at which thermally-induced cracking 

begins (350 ̊C). Correlation between this temperature and temperature at which significant reductions in 

compressive strengths occurred (325 ̊C) were found to be good. Correlations were also found to be good for the 

limestone aggregate and granite aggregate with OPC mixes. It was concluded that measurements of crack density 

after transient heating can be used to identify the depth to which the compressive strength of concrete is likely to 

have been significantly affected. Table 5.7 presents an example of criteria that have been established to aid in 

classifying the extent of microcracking in concrete [5.2].

Figure 5.7 Correlation of crack density and residual compressive strength vs temperature.

Source: N. Short and J. Purkiss, ―Petrographic Analysis of Fire-Damaged Concrete,‖ Proceedings of the Workshop 

Fire Design of Concrete Structures: What now? What next?, pp. 221-230, Milan University of 

Technology, Italy, December 2-3, 2004.
263

Table 5.7 Criteria used to classify the extent of microcracking in concrete

Classification of microcracking Description of microcracking

Very low < 20% of the cement paste contains more than 5 microcracks/surface unit1

Low 20-40% of the cement paste contains more than 5 microcracks/surface unit

Moderate 40-60% of the cement paste contains more than 5 microcracks/surface unit

High 60-80% of the cement paste contains more than 5 microcracks/surface unit

Very high > 80% of the cement paste contains more than 5 microcracks/surface unit

1This applies to a specific surface unit area, for example mm2

or cm2, as long as the same unit is used throughout the 

investigation. 

Source: T.G. Nijland and J.A. Larbi, ―Unraveling the Temperature Distribution in Fire-Damaged Concrete by 

Means of PFM Microscopy: Outline of Approach and Review of Useful Readings,‖ HERON 46(4), pp. 

253-264, 2001.

Without removal of concrete or presence of spalling, assessments of color change in the field can be ineffective in 

assessing the depth of concrete potentially affected as a result of a thermal excursion where temperature gradients 

can be large. Utilizing samples removed from the structure in question, modern color measurement systems can be 

utilized in the laboratory to provide a more detailed and objective inspection of concrete color changes resulting 

from exposure to elevated temperature [5.17,5.23,5.25]. 

A simplified approach to colorimetry has been proposed that is based on analysis of a digital camera photograph of a 

section through the concrete [5.14]. Although it is noted that digital cameras are not that accurate from a 

colorimetric viewpoint, they are capable of exhibiting color variations from point to point in the same sample. Also 

as a result of the extensive amount of data available in a single digital image, the mortar and aggregate can be 

analyzed separately. Color measurement is based on three color-matching functions with each defining a 3-D space 

(e.g., Commission Internationale de l’Eclairage red-green-blue system). Figure 5.8 presents an example related to 

digital camera colorimetry evaluation of concrete after exposure to elevated temperature. In Figure 5.8(a) the digital 

image of a concrete specimen heated to 600 ̊C is presented. Figure 5.8(b) illustrates the change in chromaticity on 

heating and for the Mont Blanc tunnel results there is a variation toward the red and yellow directions denoting an 

increasing temperature. Figure 5.8(c) presents color profiles obtained from four cores removed from a 80-mm-thick 

panel that had been subjected to a thermal gradient (i.e., > 5 ̊C/mm). The breakpoint corresponds to the change in 

color variation profile and when combined with temperature and strength profile results (when available) can 

indicate the maximum temperature and provide an estimate of residual strength.

Figure 5.8 Digital camera colorimetry: (a) uniformly heated concrete core, 

(b) effect of high temperature on chromaticity of ordinary concrete, and 

(c) color variation profiles with panel depth after exposure to constant thermal gradient.

Source: M. Colombo and R. Felicetti, ―New NDT Techniques for the Assessment of Fire-Damaged Concrete 

Structures,‖ Fire Safety Journal 42, pp. 461-472, 2007.
264

The fire behavior test (FB Test) has been developed for use in determining the depth of deteriorated concrete in 

structural elements subjected to fire [5.26]. After removal of concrete cores form the structure, 1.5-cm-thick disks 

are sliced along the complete length of the core starting from the face that was subjected to the highest temperature 

(e.g., fire). The disks are then placed into an oven and dried for 24 hours. After drying the disks are permitted to 

cool to room temperature and weighed to determine the dry weight. The disks are then submerged in water for 48 

hours to saturate them, after which the disk’s surfaces are dried and the disks reweighed to obtain a saturated weight 

so that water absorption can be calculated. The disks are then loaded in diametrical compression to obtain a tensile 

failure stress. Results are used to plot the variation in water absorption and tensile failure stress as a function of 

depth. Water absorption tends to decrease with depth (or distance from the heated face) until it reaches a constant 

value. This results because greater porosity and cracking occur at the heated face. Tensile stress also tends to 

increase with depth until it stabilizes. The lower tensile stress also is the result of the presence of increased porosity 

and cracking near the heated surface. The interface between deteriorated and sound concrete occurs where the water 

absorption and tensile stress start to stabilize.

5.5 Summary

Damage to reinforced concrete structural members resulting from thermal loadings can range from cosmetic 

blemishes to more serious damage (e.g., misalignment and distortions). Heating of concrete can result in a number 

of changes to its structure that can produce cracking, spalling, debonding of aggregate, expansion, and mineralogical 

and chemical changes. In the cement paste, evaporation and dissolution, dehydration and dissociation of ettringite, 

gypsum, calcium hydroxide, calcium carbonate and other phases such as calcium silicate hydrates can occur. 

Evaluation of a reinforced concrete structure following a thermal excursion is important in order to determine 

whether the structure is sound structurally or to assess items such as whether the steel reinforcement has been 

exposed to temperatures high enough to affect its properties. An on-site condition assessment involving a visual 

inspection is typically the first step in providing information related to the maximum temperature experienced and 

the suitability of the structure for continued service. Guidance has been developed for conduct of a condition 

assessment as well as visual classification of concrete fire damage. For concretes containing certain aggregate types 

(e.g., siliceous), color change can provide important information of the potential structural significance of the fire in 

that a color change from normal to pink or red will occur in the temperature range of 300 ̊ to 600 ̊C which 

corresponds to the temperature range at which significant loss of concrete strength occurs and steel reinforcement 

residual strength can decrease. A number of field-testing methods are available for performing a more detailed 

assessment of the material in-situ (e.g., rebound hammer, impact echo, and ultrasonics). Removal and testing of 

concrete cores is probably the most effective approach for obtaining detailed information on the mechanical 

properties and internal condition of the concrete. Petrography is the primary tool for assessing the internal condition 

of concrete and can be used to determine the depth of damage for reinforced concrete as well as a means to 

determine whether the damage was caused by thermal exposure or some other mechanism. Petrographic analysis of 

extracted concrete core samples also can provide information on bond loss between the cement matrix and steel 

reinforcement, crack orientation and it relationship to the aggregate, microcracking, extent of cement hydration, 

chemical compositional changes of cementitious material and aggregate, and temperature distribution within a 

concrete depth. Examples of other potential methods for examining and assessing concrete materials after being 

subjected to a thermal excursion include crack density determinations, colorimetry using digital cameras, and the 

fire behavior test. 

5.6 References

5.1 D.A. St John, A.W. Poole, and I. Sims, Concrete Petrography. A Handbook of Investigative Techniques, 

478 pages, Arnold Publishing, London, United Kingdom, 1998.

5.2 T.G. Nijland and J.A. Larbi, ―Unraveling the Temperature Distribution in Fire-Damaged Concrete by 

Means of PFM Microscopy: Outline of Approach and Review of Useful Readings,‖ HERON 46(4), 

pp. 253-264, 2001.

5.3 Guide for Evaluation of Concrete Structures before Rehabilitation, ACI 364.1R-07, American Concrete 

Institute, Farmington Hills, Michigan, 2007.

5.4 Guideline for Structural Condition Assessment of Existing Buildings, SEI/ASCE 11-99, 148 pages, 

American Society of Civil Engineers, Reston, Virginia, 1999. 

5.5 Assessment and Repair of Fire-Damaged Concrete Structures, Technical Report No. 33, Concrete Society, 

Camberley, United Kingdom, 1990.
265

5.6 ―Concrete Structures After Fire,‖ Concrete Construction 17(3), Addison, Illinois, 1972.

5.7 Assessing the Condition and Repair Alternatives of Fire-Exposed Concrete and Masonry Members, Report 

SR332.01B, National Codes and Standards Council of the Concrete and Masonry Industries, Skokie, 

Illinois, August 1994.

5.8 J. Gales, L.A. Bisby, C. MacDougall, and K. MacLean, ―Transient High-Temperature Stress Relaxation of 

Prestressing Tendons in Unbonded Construction,‖ Fire Safety Journal 44, pp. 570-579, 2009.

5.9 J.P. Ingham, ―Assessment of Fire-Damaged Concrete and Masonry Structures: Application of 

Petrography,‖ 11th Euroseminar on Microscopy Applied to Building Materials, 16 pages, Porto, Portugal,

5-9 June 2007.

5.10 C.A. Menzel, Tests of Fire Resistance and Strength of Walls of Concrete Masonry Units, Portland Cement 

Association, Skokie, Illinois, 1934.

5.11 K. Willam, Y. Xi, K. Lee, and B. Kim, ―Thermal Response Behavior of Reinforced Concrete Structures in 

Nuclear Power Plants,‖ Department of Civil, Environmental, and Architectural Engineering, University of 

Colorado, Boulder, 2009.

5.12 B. Erlin, W.G. Hime, and W.H. Kuenning, ―Evaluating Fire Damage to Concrete Structures,‖ Concrete 

Construction, Publication #C720154, Aberdeen Group, Addison, Illinois, April 1972.

5.13 Standard Practice for Petrographic Examination of Hardened Concrete, ASTM C 856-04, American 

Society for Testing and Materials, West Conshohocken, Pennsylvania, 2004.

5.14 R. Felicetti, ―Digital Camera Colorimetry for the Assessment of Fire-Damaged Concrete,‖ Proceedings of 

the Workshop Fire Design of Concrete Structures: What now? What next?, pp. 211-220, Milan University 

of Technology, Italy, December 2-3, 2004.

5.15 R.K. Schroeder, Post-Fire Analysis of Construction Materials, PhD Thesis, University of California, 

Berkeley, 1999.

5.16 W-T. Chang, Y-S. Giang, C-T. Wang, and C-W. Huang, ―Concrete at High Temperatures Above 1000 ̊C,‖ 

1993 International Carnahan Conference on Security Technology, Institute of Electrical and Electronics 

Engineers, Ottawa, Ontario, Canada, October 1993.

5.17 N.R. Short, J.A. Purkiss and S.E. Guise, ―Assessment of Fire Damaged Concrete Using Colour Image 

Analysis,‖ Construction and Building Materials 15(1), pp. 9-15, February 2001.

5.18 Handbook of Nondestructive Testing of Concrete, V.M. Malhotra and N. J. Carino Editors, CRC Press, 

Boca Raton, Florida, 1991

5.19 D.J. Naus, Inspection of Nuclear Power Plant Structures – Overview of Methods and Related Applications, 

ORNL/TM-2007/191, Oak Ridge National Laboratory, Oak Ridge, Tennessee, May 2009.

5.20 M. Colombo and R. Felicetti, ―New NDT Techniques for the Assessment of Fire-Damaged Concrete 

Structures,‖ Fire Safety Journal 42, pp. 461-472, 2007.

5.21 R. Felicetti, ―The Drilling-Resistance Test for Assessment of Thermal Damage in Concrete,‖ Proceedings 

of the Workshop Fire Design of Concrete Structures: What now? What next?, pp. 241-248, Milan 

University of Technology, Italy, December 2-3, 2004.

5.22 A. Di Maio, G. Giaccio, and R. Zerbino, ―Non-Destructive Tests for the Evaluation of Concrete Exposed to 

High Temperature,‖ Cement, Concrete, and Aggregates 24(2), pp.1-10, 2002.

5.23 N. Short and J. Purkiss, ―Petrographic Analysis of Fire-Damaged Concrete,‖ Proceedings of the Workshop 

Fire Design of Concrete Structures: What now? What next?, pp. 221-230, Milan University of 

Technology, Italy, December 2-3, 2004.

5.24 H.N. Walker, D.S. Lane, and P.E. Stutzman, Petrographic Methods of Examining Hardened Concrete: A 

Petrographic Manual, FHWA-HRT-04-150, Federal Highway Administration, U.S. Department of 

Transportation, McLean, Virginia, July 2006.

5.25 N. Yüzer, F. Aköz, L.D. Öztürk, ―Compressive Strength-Color Change Relation in Mortars at High 

Temperature,‖ Cement and Concrete Research 34, pp. 1803-1807, 2004. 

5.26 J.R. dos Santos, F.A. Branco, and J. de Brito, ―Assessment of Concrete Structures Subjected to Fire – The 

FB Test,‖ Magazine of Concrete Research 54(3), pp. 203-208, June 2002.
266
267

6 SUMMARY AND CONCLUSIONS

6.1 Summary

Under normal conditions most concrete structures are subjected to a range of temperature no more severe than that 

imposed by ambient environmental conditions. However, there are important cases where these structures may be 

exposed to much higher temperatures (e.g., jet aircraft engine blasts, building fires, chemical and metallurgical 

industrial applications in which the concrete is in close proximity to furnaces, and some nuclear power-related 

postulated accident conditions). Of primary interest in the present study is the behavior of reinforced concrete 

elements in designs of new generation reactor concepts in which the concrete may be exposed to long-term steadystate temperatures in excess of the present ASME Code limit of 65°C. Secondary interests include performance of 

concrete associated with radioactive waste storage and disposal facilities, and postulated design-basis accident 

conditions involving unscheduled thermal excursions. Under such applications the effect of elevated temperature on 

certain mechanical and physical properties may determine whether the concrete will maintain its structural integrity. 

Concrete’s properties are more complex than for most materials because not only is the concrete a composite 

material whose constituents have different properties, but its properties also depend on moisture and porosity. 

Exposure of concrete to elevated temperature affects its mechanical and physical properties. Elements could distort 

and displace, and, under certain conditions, the concrete surfaces could spall due to the build up of steam pressure. 

Because thermally induced dimensional changes, loss of structural integrity, and release of moisture and gases 

resulting from the migration of free water could adversely affect plant operations and safety, a complete 

understanding of the behavior of concrete under long-term elevated-temperature exposure as well as both during and 

after a thermal excursion resulting from a postulated design-basis accident condition is essential for reliable design 

evaluations and assessments. Because the properties of concrete change with respect to time and the environment to 

which it is exposed, an assessment of the effects of concrete aging is also important in performing safety 

evaluations.

The objective of this study was to provide a compilation of data and information on the effects of elevated 

temperature on concrete materials. The effect of elevated temperature exposure on the general behavior of Portland 

cement pastes including supplementary cementitious materials, aggregate materials, and the bond between cement 

paste and aggregate is discussed. Data and information is provided on the influence of elevated temperature on the 

mechanical and physical properties of concrete materials. Mechanical characteristics addressed include: stress and 

strain characteristics, Poisson’s ratio, modulus of elasticity, compressive strength, thermal cycling, tensile strength, 

shrinkage and creep, concrete-steel reinforcement bond strength, fracture energy and fracture toughness, long-term 

exposure, radiation shielding effectiveness, and multiaxial conditions. Physical characteristics discussed include: 

porosity and density, coefficient of thermal expansion, thermal conductivity, thermal diffusivity, specific heat, heat 

ablation and erosion rates, moisture diffusion, and simulated hot spots. Heavyweight concretes are described and 

results on the effect of elevated temperature on their properties and shielding effectiveness presented. Examples of 

design codes and standards that address concrete under elevated temperature applications are identified and 

described. Methods that can be utilized in assessments of concrete that has been exposed to high temperature are 

discussed.

6.2 Conclusions

A substantial body of knowledge on the material properties of ordinary Portland cement concretes at elevated 

temperature is available. The use of these data for a quantitative interpretation of the response of reinforced concrete 

structural elements in nuclear power plants to long-term moderate elevated-temperature exposure (≥65°C) or design 

basis and hypothetical severe accident conditions needs to be carefully evaluated. In many of these elevatedtemperature tests, neither representative materials nor representative environmental conditions were modeled: 

(1) samples were tested hot or cold, (2) moisture migration was either free or totally restricted, (3) concrete was 

either loaded or unloaded while heated, (4) concrete constituents and proportions varied from mix to mix, (5) test 

specimen size was not consistent, (6) specimens were tested at different degrees of hydration and moisture contents, 

and (7) heat up rates and thermal stabilization periods varied. 
268

Based on information presented in the report, several observations can be made relative to the behavior of Portland 

cement concrete materials at elevated temperature:

General Behavior

• Deterioration of concrete’s mechanical properties can be attributed to three material factors: 

(1) physicochemical changes in the cement paste, (2) physicochemical changes in the aggregate, and 

(3) thermal incompatibility between the aggregate and the cement paste. Concrete properties are influenced 

by environmental factors such as temperature level, heating rate, applied loading, and external sealing 

influencing moisture loss. 

• Key material features of Portland cement paste that influence properties of concrete at high temperature are 

its moisture state (i.e., sealed or unsealed), chemical structure (i.e., loss of chemically-combined water from 

C-S-H in unsealed condition, CaO/SiO2 ratio of hydrate in sealed condition, and amount Ca(OH)2 crystals 

in sealed and unsealed condition), and physical structure (i.e., total pore volume, average pore size, and 

amorphous/crystalline structure of solid).

• Microcracking is a major cause of deterioration when concretes are exposed to high temperatures and is 

reported to initiate around calcium hydroxide crystals and then around unhydrated cement particles.

• The aggregate-cement paste bond region has been shown to be the weakest link because it is normally 

weaker that the cement paste which is normally weaker than the aggregate. If the aggregate-cement paste 

bond fails on heating, chemically, or as a result of thermal incompatibility between the aggregate and 

cement paste, the concrete will exhibit a significant reduction in strength, even if both the aggregate and 

surrounding mortar matrix remain intact.

• Aggregate characteristics of importance to behavior of concrete at elevated temperature include physical 

properties (e.g., thermal conductivity and thermal expansion), chemical properties (e.g., chemical stability 

at temperature), and thermal stability/integrity.

• Normally aggregates are more stable at elevated temperature than the hardened cement paste when exposed 

to high temperatures, although decomposition of some less thermally stable aggregates may occur when 

exposure temperatures are extremely high. Lightweight aggregates exhibit little thermal expansion and 

therefore little damage to concrete while having small thermal conductivity. It has been noted that the 

thermal stability of aggregates increases in order of gravel, limestone, basalt, and lightweight.

• In general, for structural applications involving service temperatures in the range of ambient to 300°C or 

400°C, provided many temperature cycles of large magnitude are not present, Portland cement concretes 

are the best materials if heat-resistant aggregates (basalt, limestone, or serpentine) are used. 400 ̊C appears 

to be a critical temperature for Portland cement concretes above which concretes would disintegrate on 

subsequent post-cooling to ambient conditions. Cracking of heated concrete during post-cooling can be 

ascribed to the rehydration of dissociated calcium hydroxide resulting in a 44% volume increase. Reduced 

calcium hydroxide in cement paste due to presence of pulverized fly ash that can consume the calcium 

hydroxide can reduce cracking. At higher temperatures or for prolonged exposure to temperatures around 

600°C, special procedures would have to be considered such as removal of the evaporable water by 

moderate heating.

Mechanical Properties

Stress-Strain Behavior

• Relative to temperature effects on concrete’s stress-strain curve, several general observations can be made. 

The ascending branch of the stress-strain curve consists of three components: (1) an elastic recoverable 

strain that is temperature dependent and is strongly influenced by the load level during initial heating to the 

test temperature; (2) an irrecoverable plastic strain component; and (3) a time-dependent creep component 

that is normally small at room temperature but can be significant at high temperatures, particularly above 

550 ̊C.

• Under steady-state conditions, the original concrete strength, water-cement ratio, heating rate, and type of 

cement have minor influence on the stress-strain behavior. Aggregate-cement ratio, aggregate type, and 

presence of a sustained load during heating affect the shape of the stress-strain curve. Harder aggregates 
269

(siliceous, basaltic) exhibit a steeper decrease of initial slope due to elevated temperatures than softer 

aggregates (lightweight). Other parameters that affect the stress-strain relationship at elevated temperature 

are the moisture condition and the number of thermal cycles.

• Aggregate type has an effect on ultimate strain under high temperature conditions (i.e., carbonate aggregate 

concrete strain at peak strength was up to 40% greater than that for siliceous aggregate concrete). Siliceous 

aggregates expand the most and give the largest damage, and are further damaged on cooling.

• As the temperature increases the ultimate strain increases and the stiffness decreases. Specimens tested at 

temperature are stiffer than specimens permitted to cool to room temperature prior to testing; type of 

cement and duration of thermal treatment have a minor effect on slope of stress-strain curve. Compressive 

peak strain increases almost linearly with temperature. Peak strain after exposure (residual) is higher than 

peak strain measured at temperature and is greater for concrete cooled rapidly with water than cooled 

slowly by air.

• Unsealed specimens are stiffer than sealed specimens but strains at ultimate are reduced.

• Mineral additions can improve the elevated temperature stress-strain performance of concrete. 

• The mechanical strains in biaxial compression loading were found to be dependent on the stress level, the 

stress ratio, and the test temperature. When concrete is under a biaxial tension-compression state of stress 

the capability of concrete to resist cracking is diminished, both the tensile and compressive strengths are 

found to decrease rapidly with increasing exposure temperature, and the concrete fails abruptly due to 

tensile stress.

Poisson’s Ratio

• Poisson’s ratio data at elevated temperature are limited and tend to be inconsistent.

• Poisson’s ratio values tend to increase with increasing moisture content at high temperatures and Poisson’s 

ratio values after drying are less than before drying.

• When the specimen is permitted to cool after a given change in state due to heating (T<400 ̊C), the 

variation of Poisson’s ratio with temperature is slight and negligible when evaporable water has been 

removed.

• The first few thermal cycles tend to decrease the Poisson’s ratio the most with the number of thermal cycles 

to reach a minimum value decreasing as the concrete strength increases. Poisson’s ratio tends to return to 

the room temperature value with an increase in number of thermal cycles. 

Modulus of Elasticity

• Elastic modulus decreases with elevated temperature exposure due to breakage of bonds in the cement 

paste microstructure with the reduction in modulus increasing as the heating rate increases. Variation of 

modulus values at temperatures up to 80 ̊C is considerable, primarily as a result of use of different 

aggregate materials, and above 100 ̊C the decrease is linear with increasing temperature up to a critical 

temperature at which concrete experiences deterioration.

• Type of aggregate has strong influence on modulus with lightweight concrete exhibiting the lowest decrease 

with temperature and siliceous aggregate concrete the highest. Aggregates that are more compatible with the 

cement paste and chemically stable provide lower loss of modulus. Concrete containing aggregates with low 

thermal expansion experience a greater reduction in modulus than those with a higher thermal expansion.

• Modulus after high temperature exposure (residual) is lower than that obtained at temperature.

• Normal strength concretes (fc’ < 60 MPa) retain their modulus better at temperature than high strength 

concretes.

• Sealed (mass concrete) specimen modulus values tend to be more sensitive to elevated temperature than 

specimens that are unsealed.

• The presence of a preload improves the modulus retention under high temperature exposure.

• The method of cooling after elevated temperature exposure affects the residual modulus of elasticity with 

rapid cooling (water quenching) producing the lowest values of residual modulus.

Compressive Strength

• Considerable data scatter exists for relative compressive strength (unsealed, hot testing), particularly at the 

lower exposure temperatures, due to different materials and testing conditions utilized (e.g., constituents 

and mixture proportions, specimen size and shape, specimens tested at different degrees of hydration and 

initial moisture contents, and heating rates and thermal stabilization periods). Considerable data scatter 
270

also exists for residual compressive strength (unsealed, cold testing), particularly at lower temperatures due 

to differences in material and environmental factors.

• Original concrete strength of normal strength concrete, type of cement, aggregate size, heating rate, and 

water-cement ratio appear to have a minor effect on the relative concrete strength at elevated temperature.

• Type of aggregate appears to be one of main factors influencing concrete compressive strength at high 

temperature with siliceous aggregate concrete having lower strengths (by percentage) at high temperature 

than calcareous and lightweight aggregate concretes.

• Concrete in the temperature range of 20°C to 200°C can show a small strength loss. Between 22 ̊ and 

120°C any strength loss that occurs is attributed to the thermal swelling of the physically bound water, that 

causes disjoint pressures. A regain of strength is often observed between 120°C and 300°C and is 

attributed to greater van der Waals forces as a result of the cement gel layers moving closer to each other 

during heating. Between 200°C and 250°C the residual compressive strength is nearly constant. Beyond 

350°C there can be a rapid decrease in strength. 

• Strength losses for unstressed concretes up to about 300 ̊C are generally < 20% irrespective of type of 

cement or aggregate. At temperatures > 300 ̊C strength losses can become relatively great and at 

temperatures > 450 ̊C concrete compressive strength drops significantly due to differences in thermal 

expansion coefficients between aggregate and cement paste (i.e., loss of bond) and decomposition of 

calcium hydroxide.

• The presence of preload, within reasonable limits, improves retention of compressive strength at elevated 

temperature for both Portland cement concretes and Portland cement concretes containing supplementary 

cementitious materials. Although data are limited, a similar conclusion can apparently be derived with 

respect to residual compressive strength retention. Improved performance has been attributed to 

densification of the cement paste resulting in a large transient creep component and possibly a reduction in 

porosity relative to the unloaded state, and precompression can reduce the tensile stresses in the concrete, 

particularly during cooling.

• Moisture content at time of testing has a significant effect on the strength of concrete at elevated 

temperature with strength of unsealed concrete being higher than strength of sealed concrete.

• Factors that contribute to the general trend for concrete compressive strength to decrease with increasing 

temperature are: aggregate damage; weakening of the cement paste-aggregate bond; and weakening of the 

cement paste due to an increase in porosity on dehydration, partial breakdown of the C-S-H, chemical 

transformation on hydrothermal reactions, and development of cracking.

• Concrete containing fly ash can exhibit an increase in strength in the temperature range of 121 ̊ to 149 ̊C 

due to the formation of tobermorite which has been reported to be two to three times stronger than C-S-H 

gel. Although limited, results for concretes containing supplementary cementitious materials indicate that 

the residual compressive strength retention is highest for concretes containing slag and fly ash, followed by 

concrete containing metakaoline, with silica fume exhibiting the lowest residual compressive strength at the 

higher temperatures.

• The higher the cement content the greater the loss in strength. Strength loss in saturated normal strength 

concrete is greater than dry concrete and higher moisture contents contribute to increased spalling of 

concrete during significant thermal gradients such as could occur under fire conditions.

• Concrete that is rapidly cooled form elevated temperature (e.g., water quenching) exhibits a lower residual 

compressive strength than specimens that are gradually cooled from elevated temperature prior to testing. 

The effect of rapid cooling seems to decrease somewhat at high temperatures, however, the residual 

compressive strength already has been severely reduced by exposure to these temperatures.

• Concrete containing lightweight/thermally stable aggregates exhibit greater retention of compressive 

strength at temperature (relative) than those permitted to cool to room temperature (residual) prior to 

testing.

• Residual and relative compressive strengths of fibrous concretes at temperatures above 200 ̊C both exhibit 

a trend for a linear decrease in strength with increasing temperature.

Thermal Cycling

• Thermal cycling, even at relatively low temperatures (e.g., 65 ̊C) can have a degrading effect on concrete’s 

mechanical properties. Compressive strength, tensile strength, and bond strength to steel reinforcement 

decrease under thermal cycling.

• The largest percentage decrease in properties for thermal cycling at higher temperatures (e.g., 200 ̊ to 

300 ̊C) occurs during the first thermal cycle with the extent of damage dependent on the aggregate type and 
271

is associated with loss of bond between the aggregate and cement paste matrix.

Tensile Strength

• Aggregate type and mixture proportions have a significant effect on the tensile strength vs temperature 

relationship of concrete.

• The effect of elevated temperature on the tensile strength of concrete shows a similar trend to its effect on 

compressive strength, but tensile strength is more sensitive to deterioration at elevated temperature.

• The decrease in tensile strength of calcareous aggregate concrete with temperature is greater than that of 

siliceous aggregate concrete at 500 ̊C, being about twice as much at this temperature.

• Concretes with lower cement contents have lower reduction in tensile strength than those with higher 

cement contents.

• At relative low temperatures (T < 175 ̊C) sealed specimens seem to exhibit improved retention of splittingtensile strength relative to results obtained from unsealed specimens.

• The rate of heating has minimal effect on the tensile strength at high temperature.

• Residual tensile strength (cold testing) is somewhat lower than the tensile strength determined at 

temperature (hot testing).

• Quenching specimens in water after exposure to high temperatures produces a significant decrease in 

flexural and splitting-tensile strengths (as well as compressive strength), with the decrease increasing as the 

maximum exposure temperature increased. 

Shrinkage and Creep

• Shrinkage occurs as a result of drying or autogeneous volume change, with drying shrinkage being 

dominant, resulting from the loss of absorbed water, and increasing with increasing temperature. 

Autogeneous shrinkage results from continued cement hydration and is more prevalent in mass concrete 

structures where the total moisture content remains relatively constant.

• Shrinkage at temperatures < 100 ̊C is not considered to be a significant factor in mass concrete structures 

over their nominal 30- to 40-year design life.

• Creep of unsealed specimens is less than creep of sealed specimens.

• Creep decreases with increasing concrete maturity and increases with increasing moisture content of the 

specimen at loading.

• Creep increases as the temperature increases, probably for temperatures at least up to 150 ̊C. Below 100 ̊C 

concrete creep at moderate stress levels originates in the cement paste as a result of the mutual approach of 

adjacent laminar particles of the cement gel. 

• For unsealed specimens the influence of drying becomes significant in the temperature range of 70 ̊ to 

100 ̊C. At moderate temperature levels hydration accelerates, but as the temperature increases the reverse 

of this effect takes place as dehydration accelerates creep. Above 100 ̊C drying of concrete is very rapid 

with an associated increase in creep rate until a stable moisture condition is reached. After moisture is lost 

and a stable moisture content at a given temperature is reached, the creep rate becomes less than that before 

loss of moisture.

• Creep recovery is less than associated creep strain and the degree of creep recovery seems to be 

independent of temperature but dependent on stress.

Concrete-Steel Reinforcement Bond Strength

• The bond strength between concrete and steel decreases as the temperature increases with the relative 

reduction in bond strength at elevated temperature being greater than that for the concrete compressive 

strength. 

• Time at temperature affects the bond strength and the bond strength at temperature is higher than residual 

bond strength.

• Ribbed steel reinforcement bars retain residual bond strength at elevated temperature better than round steel 

bars. Diameter of ribbed steel reinforcement does not appear to have a significant affect on bond strength.

• The type of aggregate is one of the main factors affecting the elevated temperature bond strength, with the 

lower the thermal strain of concrete the higher the bond strength.

• Sealed specimens retain bond strength at elevated temperature better than unsealed specimens and 

prolonged heating at moderate elevated temperatures (T < 175 ̊C) provides a positive effect on bond 

strength of sealed specimens.

• Load cycling produces a decrease in maximum bond stress.
272

• Reduction in bond stress appears to be greater for concretes in which >10% (by weight) of the Portland 

cement has been replaced by a natural pozzolanic material.

• Incorporation of random discontinuous fiber reinforcement into the concrete mix can improve residual 

bond strength at elevated temperature by increasing confinement of the steel reinforcement (e.g., reduce 

concrete cracking and spalling tendencies).

Fracture Energy and Fracture Toughness

• Residual fracture energy of normal- and high-strength concretes can exhibit an increase at temperatures up 

to about 300 ̊C or 400 ̊C due to cement hydration before it decreases with increasing temperature as a result 

of microcracking and cement dehydration

• Residual fracture toughness decreases with increasing temperature with the decrease being relatively 

significant at temperatures > 50 ̊C.

• Heating cycles produce a decrease in fracture toughness with the amount of decrease becoming more 

significant at temperatures greater than 300 ̊C.

Long-Term Exposure (Aging)

• Results for exposure of concrete to elevated temperature for relatively long periods of time are limited.

• For moderate exposure periods (e.g., < 180 days) and temperatures (e.g., T < 232 ̊C), the concrete 

compressive strength and modulus of elasticity of sealed specimens tended to decrease with increasing 

exposure time and temperature, while specimens that were not sealed exhibited an improvement of or very 

slight reduction in compressive strength and improved retention of the modulus of elasticity relative to 

sealed specimens. The decreased performance of the sealed specimens was attributed to deterioration in 

the structural properties of the cement gel resulting from the saturated steam pressure at high temperatures. 

Modulus of elasticity was affected more than compressive strength as a result of extended exposure at 

elevated temperature.

• In general, most of the loss of concrete’s mechanical properties under elevated temperature exposure for 

extended exposure periods will occur during the first two or three months exposure for unsealed conditions 

and at longer periods for sealed conditions as the moisture condition stabilizes (e.g., one to two years).

• There is an indication that partial replacement of Portland cement with fly ash or blast furnace slag imparts 

improved performance at elevated temperature.

Radiation Shielding Effectiveness

• Drying resulting from elevated temperature exposure reduces the shielding (attenuation) effectiveness of 

concrete as a result of cracking or reduction in hydrogen content.

• For neutron shields maximum recommended internal and ambient concrete temperatures are 88 ̊ and 71 ̊C, 

respectively, and for gamma shields the corresponding recommended maximum concrete temperatures are 

177 ̊ and 149 ̊C, respectively.

Multiaxial Conditions

• The platen-interface constraint influences the results for multiaxial loadings (i.e., strength increases as 

constraint increases).

• Under biaxial compression loading the strength of concrete decreases with increasing temperature with the 

maximum aggregate size having a significant effect on behavior while the aggregate content and watercement ratio have a lesser effect.

• Differences between uniaxial and biaxial concrete strength increased as the temperature increased.

Physical Properties and Thermal Effects

Porosity and Density

• As the temperature increases the porosity and average pore size of the concrete increases thus reducing the 

cement bulk mass density.

• Aggregate type plays an important role in the change in concrete density with increasing temperature (e.g., 

thermal dilation and dissociation).
273

Coefficient of Thermal Expansion

• Concrete thermal expansion is a non-linear function of temperature with aggregate type and coarse 

aggregate volume fraction being the primary factors influencing its value. The coefficient of thermal 

expansion of concrete increases with temperature even though the cement paste contracts because the 

aggregate dominates and increases as the temperature increases.

• Moisture content, water-cement ratio, and type of cement only affect the coefficient of thermal expansion at 

relatively low temperatures (e.g., T < 200 ̊C).

• Considerable differences exist relative to expansion of aggregate materials with aggregates having the 

highest percentage of silica by weight exhibiting the largest expansion. Lightweight aggregate materials 

have lower thermal expansion than naturally-occurring aggregate materials and can show residual 

contraction after heating.

• Thermal expansion on subsequent cooling is not fully reversible due to irreversible chemical changes and 

changes in microstructure.

• The presence of load reduces the thermal expansion of concrete with increasing temperature.

Thermal Conductivity 

• Primary factors that affect concrete thermal conductivity are moisture content, hardened cement paste, pore 

volume and distribution, and amount and type of aggregate material. The amount and type of coarse 

aggregate used have the most influence, however, because aggregates normally occupy 60 to 80% the 

concrete volume.

• Thermal conductivities of highly crystalline aggregate concretes decrease with increasing temperature up to 

1000 ̊C, while amorphous aggregate concretes have thermal conductivity values that are relatively constant.

• In general, at temperatures up to 100 ̊C the thermal conductivity seems to increase with temperature; at 

temperatures greater than 100 ̊C the conductivity decreases; at 300 ̊ to 400 ̊C there is a further decrease as 

increased cracking occurs; at temperatures greater than 600 ̊C the thermal conductivity increases slightly; 

and at temperatures greater than 800 ̊C up to melting there is only a small change in thermal conductivity.

• Lightweight aggregate concrete’s have lower thermal conductivity values because of their porosity and the 

thermal conductivity is relatively constant or increases slightly for temperatures up to 1000 ̊C.

Thermal Diffusivity 

• Thermal diffusivity of concrete is determined by the thermal properties of its constituents and at normal 

temperatures is mainly governed by the aggregate diffusivity.

• Concrete thermal diffusivity decreases fairly rapidly with an increase in temperature (T < ~200 ̊C) and 

exhibits a trend to stabilize at high temperatures (T > 600 ̊C).

Specific Heat

• At room temperature the aggregate type, mix proportions, and age do not have a great effect on the specific 

heat of concrete, but the moisture content in the concrete at lower temperatures does have an effect.

• Specific heat of concrete increases with increasing temperature and at high temperatures is sensitive to 

transformations that take place (e.g., water vaporization, dissociation of Ca(OH)2, quartz transformation in 

aggregates).

Heat of Ablation and Erosion Rates

• Heat of ablation is of interest for hypothetical core disruptive accidents involving a core melt relative to 

how much energy is dissipated by the concrete structures that are designed to contain the core melt. 

• Erosion rate refers to the velocity at which concrete subjected to high temperatures is decomposed by 

mechanical disintegration and by the melting process.

• For a heat flow between 20 and 200 W•cm2the heat of ablation of concrete has been estimated to be in the 

range of 1474 to 6000 kJ•kg-1.

• The amount of heat required to transform a given volume of concrete to a molten state is about twice as 

much for limestone aggregate as for quartzitic aggregate concrete due to the additional heat required to 

decarbonate the limestone aggregate.

• Erosion rates determined experimentally for limestone, quartzitic, and basalt aggregate concretes are about 

20 to 66, 30 to 44, and 10 to 22 mm/min, respectively.
274

Moisture Diffusion and Pore Pressure

• Heating of concrete can produce large pore pressures that are functions of temperature, temperature history, 

and size of specimen.

• Transfer of free moisture in concrete can affect the concrete strength, thermal expansion, shrinkage, and 

creep, as well as potentially affect the concrete durability and radiation shielding capability.

• Under normal operating conditions in a nuclear plant (i.e., T < 65 ̊C) moisture migration in mass concrete 

structures is a very slow process and drying is not likely to be an important factor in thick-section concrete 

structures.

• Limited data on moisture distribution in a nuclear power plant containment wall indicates that even after 

30 years the walls were still drying, with the drying being extremely slow due to the large thickness and 

one-sided drying; drying was different in different parts of the containment due to different temperature 

levels and climatic conditions, and the concrete continued to mature because moisture conditions in the 

inner parts of the wall permitted further cement hydration.

• In unsealed specimens diffusion and eventual drying leads to lower pore pressures than obtained for sealed 

specimens at the same temperature level.

• Pore pressures in concrete do not increase significantly until the concrete temperature reaches about 105 ̊C.

Simulated Hot Spot Tests

• Results of tests designed to simulate hot spots that could develop in the vicinity of hot penetrations of a 

localized failure of a liner cooling system indicate that the effect is localized, some loss of concrete strength 

would occur, but not significant enough to affect structural integrity, and pressure build up behind a liner 

may not be significant due to pressure relief along the liner-concrete interface.

Radiation Shielding Concretes

• Thermal cycling can reduce the compressive strength, modulus of rupture, and static modulus of elasticity 

of radiation shielding concretes.

• Specimens loaded while heated retain their residual compressive strength better than specimens that were 

not loaded during thermal treatment.

• Concrete density decreases with increasing temperature as evaporable water is removed, with the value 

becoming relatively stable after evaporable water is lost.

• The impact of elevated temperature on the water content of a shielding concrete has a significant effect on 

its nuclear properties, in particular its neutron radiation attenuation. As temperature increases the neutron 

flux density transmitted through concrete increases for a given section thickness.

• The intensity of radiation transmitted through cracked (collinear) concrete increases as the crack width 

increases and specimen thickness decreases.

Codes and Standards

• Codes and standards for concrete technology recognize that concrete strength tends to decrease with 

increasing temperature. Consequently, current design procedures specify concrete temperature limits to 

ensure predictable concrete behavior. 

• Analytical models for accurately predicting the response of a structure to thermal loadings for practical 

design considerations, where thermal environments exceed the limits contained in the code, are very 

complex. As a result, most existing methods utilize various types and degrees of simplification that affect

the accuracy of results. Current designs for nuclear structures cover these shortcomings by appropriate 

conservatism in designs. 

• When design conditions exceed established temperature limits, experimental investigations for 

characteristic mechanical and physical properties data and for design verification may be required to avoid 

undue and impractical conservatism in design. 

Methods for Assessment of Concrete Exposed to High Temperatures

• Damage to concrete structural members resulting from thermal loadings can range from cosmetic blemishes 

to more serious damage (e.g., misalignments and distortions).
275

• Evaluation of a reinforced concrete structure following a thermal excursion is important in order to 

determine whether the structure is structurally sound or to assess items such as whether the steel 

reinforcement has been exposed to temperatures high enough to affect its properties.

• For concretes containing certain aggregate types (e.g., siliceous), color change can provide important 

information on the potential structural significance of a thermal excursion (i.e., peak temperature reached).

• A number of field testing methods (e.g., rebound hammer, impact echo, and ultrasonics) are available for 

performing assessments of reinforced concrete structures after being subjected to thermal excursions.

• Removal and testing (e.g., compressive strength and petrography) of concrete cores is probably the most 

effective approach for obtaining detailed information on the mechanical properties and internal condition of 

the concrete (e.g., depth affected and cause of damage).

• Other methods having potential for assessing concrete materials after being subjected to a thermal 

excursion include crack density determinations, colorimetry using digital cameras, fire behavior test, and 

drilling resistance test.
276
A-1

APPENDIX A – REFERENCES NOTED IN DATA PLOTS AND CITED IN FIGURES OR TABLES

A.1 References Noted in Data Plots

 Data 

Sheet # ________________________________Reference____________________________________________

1. T. Suzuki, M. Tabuchi, and K.Nagao, ―Study on the Degradation of Concrete Characteristics in the High 

Temperature Environment,‖ Concrete Under Severe Conditions: Environment and Loading, Vol. 2, 

pp. 1119–1128, E & FN Spon Publishers, 1995.

2. D. Campbell-Allen and P. M. Desai, ―The Influence of Aggregate on the Behavior of Concrete at 

Elevated Temperature,‖ Nuclear Engineering and Design 6(1), pp. 65–77, August 1967.

3. J. C. Saeman and G. W. Washa, ―Variation of Mortar and Concrete Properties with Temperature,‖ J. 

American Concrete Institute 29(5), pp. 385–395, 1957.

4. C.R. Cruz, ―Elastic Properties of Concrete at High Temperatures,‖ Department Bulletin 191, Journal of 

the Portland Cement Association Research and Development Laboratories 8(1), pp. 37-45, January 

1966.

5. J. Xiao and G. König, ―Study of Concrete at High Temperature in China—An Overview,‖ Fire Safety 

Journal 39, pp. 89–103, 2004.

6. G.G. Carette, K.E. Painter, and V.M. Malhotra, ―Sustained High Temperature Effects on Concretes Made 

with Normal Portland Cement, Normal Portland Cement and Slag, or Normal Portland Cement and Fly 

Ash,‖ Concrete International 4(7), 41–51 (July 1982).

7. J. Guo and P. Waldron, ―Deterioration of PCPV Concrete,‖ Nuclear Engineering and Design 198, 

pp. 211-226, 2000.

8. T. Yagashita, K. Shirai, C. Ito, and H. Shiojiri, ―Effect of Strain Rate on Concrete Strength Under High 

Temperature,‖ International Conference on Structures Under Shock & Impact VI, pp. 539-548, 

Cambridge, United Kingdom, WIT Press, July 2000.

9. I. Hager and P. Pimienta, ―Mechanical Properties of HPC at High Temperature,‖ Proceedings of the 

Workshop - Fire Design of Concrete Structures: What now? What Next?, pp. 95-100, 

Milan, Italy, December 2-3, 2004.

10. P.J.E. Sullivan, ―Chapter 1, The Effects of Temperature on Concrete,‖ in Developments in Concrete 

Technology -1, Applied Science Publishers, Ltd., London, 1979.

11. H. Tanyildizi and A. Coskun, ―Performance of Lightweight Concrete with Silica Fume After High 

Temperature,‖ Construction and Building Materials 22, pp. 2124-2129, 2008.

12. E. Tolentino, F.S. Lameiras, A.M. Gomes, C.A.R. de Silva, and W.L. Vasconcelos, ―Effects of High 

Temperature on the Residual Performance of Portland Cement Concretes,‖ Materials Research 5(3), 

pp. 301-307, July/September 2002.

13. E. Sancak, Y.D. Sari, and O. Simsek, ―Effects of Elevated Temperature on Compressive Strength and 

Weight Loss of the Light-Weight Concrete with Silica Fume and Superplasticizer,‖ Cement and 

Concrete Composites 30(8), pp. 715-721, 2008. 

14. C-S. Poon, S. Azhar, M. Anson, Y-L. Wong, ―Performance of Metakaolin Concrete at Elevated 

Temperature,‖ Cement and Concrete Composites 25, pp. 83-89, 2003.

15. S.K. Handoo, S. Agarwal, and S.K. Agarwall, ―Physiochemical, Mineralogical, and Morphological 

Characteristics of Concrete Exposed to Elevated Temperatures,‖ Cement and Concrete Research 32, 

pp. 1009-1018, 2002.

16. G.T.G. Mohamedbhai, ―Effect of Exposure Time and Rates of Heating and Cooling on Residual Strength 

of Heated Concrete,‖ Magazine of Concrete Research 38(136), pp. 151-158, September 1986.

17. A.N. Noumowe and C. Galle, ―Study of High Strength Concretes at Raised Temperatures up to 200 ̊C: 

Thermal Gradient and Mechanical Behaviour,‖ Transactions of 16th International Conference on 

Structural Mechanics in Reactor Technology, Paper #1580, Washington, DC, August 2001.

18. A.N. Noumowe, P. Clastres, G. Debocki, and J.-L. Costaz, ―Transient Heating Effect on High Strength 

Concrete, ― Nuclear Engineering and Design 166, pp. 99-108, 1996.

19. G.C. Hoff, A. Bilodeau, and V.M. Malhotra, ―Elevated Temperature Effects on HSC Residual Strength,‖ 

Concrete International 22(4), pp. 41-47, April 2000.

20. M. Li, C.X. Quian, and W. Sun, ‖Mechanical Properties of High-Strength Concrete After Fire,‖ Cement 

and Concrete Research 34, pp. 1001-1005, 2004.
A-2

21. Y.N. Chan, G.F. Peng, and M. Anson, ―Residual Strength and Pore Structure of High-Strength Concrete 

and Normal Strength Concrete After Exposure to High Temperatures,‖ Cement and Concrete Research

21, pp. 23-27, 1999.

22. R.S. Ravindrarajah, R. Lopez, and H. Reslan, ―Effect of Elevated Temperature on the Properties of HighStrength Concrete Containing Cement Supplementary Materials,‖ 9

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A-3

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A-4

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American Concrete Institute, Farmington Hills, Michigan, 1972.
A-5

82. D. A. Campbell-Allen, E.W.E. Roper, and H. Roper, ―An Investigation of the Effect of Elevated 

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A.2 References Cited by Others in Figures

Figure

Number Reference(s)

———— —————————————————————————————————————————

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2.54 F.J. Roux, Concrete at Elevated Temperatures, Doctoral Thesis, University of Capetown, South Africa, 

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T. Harada, T. Takeda, S, Yamane, and F, Furumura, ―Strength, Elasticity and Thermal Properties of 

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D. T. Lankard, D.L. Birkimer, F.F. Fondriest, and M.J. Snyder, ―Effects of Moisture Content on the 

Structural Properties of Portland Cement Concrete Exposed to Temperatures Up to 500°F,‖ SP-25 

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J.C. Marechal, ―Variations in the Modulus of Elasticity and Poisson’s Ratio with Temperature,‖ Paper SP 

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Michigan, 1972.

S. Ohgishi, S. Miyasaka, and J. Chida, ―On Properties of Magnetite and Serpentine Concrete at Elevated 

Temperatures for Nuclear Reactor,‖ SP 34-57 in Concrete for Nuclear Reactors, pp. 1243-1253, American 

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2.80 K.B. Hickey, Creep, Strength and Elasticity of Concrete at Elevated Temperatures, Report No. C-1257, 

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2.81 K.W. Nasser and R.P. Lothia, ―Mass Concrete Properties at High Temperatures,‖ Journal American 

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All references cited by an author may not be listed.
A-6

S. Ohgishi, S. Miyasaka, and J. Chida, ―On Properties of Magnetite and Serpentine Concrete at Elevated 

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H.L. Malhotra, ―The Effect of Temperature on the Compressive Strength of Concrete,‖ Magazine of 

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2.82 S. Ohgishi, S. Miyasaka, and J. Chida, ―On Properties of Magnetite and Serpentine Concrete at Elevated 

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Concrete Institute, Farmington Hills, Michigan, 1972.

F.J. Roux, Concrete at Elevated Temperatures, Doctoral Thesis, University of Capetown, South Africa, 

August 1974.

T. Harada, T. Takeda, S, Yamane, and F, Furumura, ―Strength, Elasticity and Thermal Properties of 

Concrete Subjected to Elevated Temperatures,‖ Paper SP-34-21 in Concrete for Nuclear Reactors, pp. 377-

406, American Concrete Institute, Farmington Hills, Michigan, 1972.

D. T. Lankard, D.L. Birkimer, F.F. Fondriest, and M.J. Snyder, ―Effects of Moisture Content on the 

Structural Properties of Portland Cement Concrete Exposed to Temperatures Up to 500°F,‖ SP-25 

Temperature and Concrete, pp. 59–102, American Concrete Institute, Farmington Hills, Michigan, 1971.

H.L. Malhotra, ―The Effect of Temperature on the Compressive Strength of Concrete,‖ Magazine of 

Concrete Research 8(2), pp.382-388, 1965.

M. S. Abrams, ―Compressive Strength of Concrete at Temperatures to 1600°F,‖ SP-25, Temperature and 

Concrete, American Concrete Institute, pp. 33–58, 1971.

2.1155 G.L. England and A.D. Ross, ―Reinforced Concrete Under Thermal Gradients,‖ Magazine of Concrete 

Research 14(40), pp. 5-12, March 1962.

K.B. Hickey, Creep, Strength, and Elasticity of Concrete at Elevated Temperature, Report No. C-1257, 

Bureau of Reclamation, Washington, DC, December 1967.

J.C. Marechal, ―Causes Physiques et Chimiques du Fluage et du Retrait du Béton,‖ RILEM Colloquim on 

the Physical and Chemical Causes of Creep and Shrinkage of Concrete, Munich, Germany, April 1968.

2.1156 K.B. Hickey, Creep, Strength, and Elasticity of Concrete at Elevated Temperature, Report No. C-1257, 

Bureau of Reclamation, Washington, DC, December 1967.

J.C. Marechal, ―Causes Physiques et Chimiques du Fluage et du Retrait du Béton,‖ RILEM Colloquim on 

the Physical and Chemical Causes of Creep and Shrinkage of Concrete, Munich, Germany, April 1968.

G.L. England and A.D. Ross, ―Reinforced Concrete Under Thermal Gradients,‖ Magazine of Concrete 

Research 14(40), pp. 5-12, March 1962.

K.W. Nasser and A.M. Neville, ―Creep of Concrete at Elevated Temperatures,‖ Journal of the American 

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S. Arthanari and C.W. Yu, ―Creep of Concrete Under Uniaxial and Biaxial Stresses at Elevated 

Temperatures,‖ Magazine of Concrete Research 19(60), pp. 149-156, September 1967.

2.240 T. Harada, T. Takeda, S, Yamane, and F, Furumura, ―Strength, Elasticity and Thermal Properties of 

Concrete Subjected to Elevated Temperatures,‖ Paper SP-34-21 in Concrete for Nuclear Reactors, 

pp. 377-406, American Concrete Institute, Farmington Hills, Michigan, 1972.

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A-7

M.S. Crowley, ―Initial Thermal Expansion Characteristics of Insulating Refractory Concretes,‖ Bulletin of 

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2.244 T. Z Harmathy. and L.W. Allen, ―Thermal Properties of Selected Masonry Unit Concretes,‖ Journal 

of American Concrete Institute 70, pp. 132–142, 1973.

T. Harada, The Thermal Expansion of Concrete at High Temperatures, Part 1.2, Dissertation, Tokyo 

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2.245 T.Z Harmathy. and L.W. Allen, ―Thermal Properties of Selected Masonry Unit Concretes,‖ Journal 

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T. Harada, The Thermal Expansion of Concrete at High Temperatures, Part 1.2, Dissertation, Tokyo 

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2.249 T. Harada, T. Takeda, S, Yamane, and F, Furumura, ―Strength, Elasticity and Thermal Properties of 

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A-8

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2.255 G. Hillenbrand et al., Untersuchung der Wechselwirkung von Kernschmelze und Reaktorbeton, 

Abschlussbericht – Förderungsvorhaben BMFT RS 154, KWU, Erlangen, Germany, May 1978.

T.Z Harmathy. and L.W. Allen, ―Thermal Properties of Selected Masonry Unit Concretes,‖ Journal 

of American Concrete Institute 70, pp. 132–142, 1973.

T. Harada, T. Takeda, S, Yamane, and F, Furumura, ―Strength, Elasticity and Thermal Properties of 

Concrete Subjected to Elevated Temperatures,‖ Paper SP-34-21 in Concrete for Nuclear Reactors, 

pp. 377-406, American Concrete Institute, Farmington Hills, Michigan, 1972.

J.A. Pogorzelski, Thermal Properties of Some Building Materials, Report to RILEM Committee 44-PHT, 

Warsaw, Poland, 1979.

2.257 T. Harada, T. Takeda, S, Yamane, and F, Furumura, ―Strength, Elasticity and Thermal Properties of 

Concrete Subjected to Elevated Temperatures,‖ Paper SP-34-21 in Concrete for Nuclear Reactors, pp. 377-

406, American Concrete Institute, Farmington Hills, Michigan, 1972.

2.261 G. Hillenbrand et al., Untersuchung der Wechselwirkung von Kernschmelze und Reaktorbeton, 

Abschlussbericht – Förderungsvorhaben BMFT RS 154, KWU, Erlangen, Germany, May 1978.

T.Z Harmathy and L.W. Allen, ―Thermal Properties of Selected Masonry Unit Concretes,‖ Journal 

of American Concrete Institute 70, pp. 132–142, 1973.

T.Z Chu, Radiant Heat Evolution of Concrete – A Study of the Erosion of Concrete Due to Surface 

Heating, SAND 77-0922, Sandia National Laboratory, Albuquerque, New Mexico, 1978.

Y. Collet and E. Tavernier, Etude des Propiétés du Béton Soumis a des Témperatures Élevés, Group de 

Travail, Comportement du Matériau Béton en Fonction de la Température, Brussels, Belgium, November 

1976.

2.262 G. Hillenbrand et al., Untersuchung der Wechselwirkung von Kernschmelze und Reaktorbeton, 

Abschlussbericht – Förderungsvorhaben BMFT RS 154, KWU, Erlangen, Germany, May 1978.

T.Z Harmathy and L.W. Allen, ―Thermal Properties of Selected Masonry Unit Concretes,‖ Journal 

of American Concrete Institute 70, pp. 132–142, 1973.

Y. Collet and E. Tavernier, Etude des Propiétés du Béton Soumis a des Témperatures Élevés, Group de 

Travail, Comportement du Matériau Béton en Fonction de la Température, Brussels, Belgium, November 

1976.

J.A. Pogorzelski, Thermal Properties of Some Building Materials, Report to RILEM Committee 44-PHT, 

Warsaw, Poland, 1979.

K. Ödeen and A. Nordström, ―Termiska Egenskaper hos Beton vid Höga Temperaturer, ‖Cement och Beton

1, Stockholm, Sweden, 1972.

A.3 References Cited by Others in Tables

 Table

Number Reference(s)

——— —————————————————————————————————————————

2.15

G. Hillenbrand et al., Untersuchung der Wechselwirkung von Kernschmelze und Reaktorbeton, 

Abschlussbericht – Förderungsvorhaben BMFT RS 154, KWU, Erlangen, Germany, May 1978.

T.Z Chu, Radiant Heat Evolution of Concrete – A Study of the Erosion of Concrete Due to Surface Heating, 

SAND 77-0922, Sandia National Laboratory, Albuquerque, New Mexico, 1978

U. Schneider and U. Diederichs, ―Physical Properties of Concrete From 20 ̊C Up To Melting,‖ Parts 1 and 

2, Betonwerk & Fertigteiltechnik, Heft 3, pp. 141-150, and Heft 4, pp. 223-230, 1981.

2.17 H.J. Sutherland, ―Acoustic Measurement of the Penetration of a Molten Metallic Pool into Concrete,‖ 

Nuclear Technology 46, ISO-355, December 1979.
A-9

D. Perinic et al., Betontiegelversuche mit Thermitschmelzen, Kernforschungszentrum Karlsruhe, KfK 2572, 

Germany, July 1979.

M. Peehs , A. Skokan and M. Reimann, ―Investigations in Germany of the Barrier Effect of Reactor 

Concrete Against Propagating Molten Corium in the Case of a Hypothetical Core Meltdown Accident of a 

LWR,‖ ENS/ANS International Topical Meeting on Nuclear Power Reactor Safety 1, Brussels, Belgium 9-

16 October 1978.
B-1

APPENDIX B - TEMPERATURE-DEPENDENT PROPERTIES OF MILD STEEL AND 

PRESTRESSING MATERIALS FOR USE WITH PORTLAND CEMENT CONCRETES

Bonded reinforcement (i.e., deformed bars) is provided to control the extent and width of cracks at operating 

temperatures, resist tensile stresses and computed compressive stresses for elastic design, and provide structural 

reinforcement where required by limit condition design procedures. Bonded reinforcement in nuclear power plant 

structures is often used in conjunction with prestressed steel. The prestressed steel provides the structural rigidity 

and the major part of the strength while the bonded reinforcement distributes cracks, increases ultimate strength and 

reinforces those areas not adequately strengthened by the prestressed steel, and provides additional safety for 

unexpected conditions of loading.

The elevated temperature response of a reinforced concrete member depends on the mechanical and thermal 

properties of the materials of which it is composed. Structural elements fabricated from reinforced concrete, 

because of their typical size, have a high thermal inertia that results in relatively slow rates of temperature increase 

through the cross section. As a result, the steel reinforcement temperatures generally are kept sufficiently low to 

avoid significant softening. In addition, due to the monolithic nature of construction, the existence of alternate load 

paths, and compartmentalazation of fires (i.e., conventional civil engineering construction), reinforced concrete 

structures generally perform well under elevated-temperature conditions such as fires. However, under certain 

scenarios (e.g., rapid heat build up) spalling of the concrete could occur to expose the steel reinforcement to the 

effects of elevated temperature, or the thermal event could last for a prolonged period of time (e.g., hot spot). As a 

result the strength of steel can decrease and deformations and other property changes may occur under prolonged 

exposure. Primary mechanical properties of steel that affect the response of structural members to thermal events 

are the strength, modulus of elasticity, coefficient of thermal expansion, and creep. Primary thermal properties are 

those that influence the temperature rise and its distribution in the member (i.e., thermal conductivity, specific heat, 

and density). Although this report primarily addresses cementitious materials, for completeness limited information 

is provided below on effects of elevated temperature on selected properties of mild steel and steel reinforcing 

materials. More detailed information on properties of steel materials at elevated temperatures has been compiled 

elsewhere [B1].

B.1 Mechanical Properties

Stress-Strain. Strength characteristics and mechanical properties of steels depend on several factors: amount and 

type of alloying constituents, heat treatment during manufacture, and retreatment in cold state (e.g., cold drawing). 

Figures B.1 and B.2 presents examples of the effect of elevated temperature on the stress-strain curve of a typical 

structural steel. 

Figure B.1 Stress-strain curves for structural steel (ASTM A36).

Source: SFPE Engineering Guide to Performance Based Fire Protection Analysis and Design of Buildings, Society 

of Fire Protection Engineers, Bethesda, Maryland, March 2000.
B-2

Figure B.2 Stress-strain curves for structural steel (ASTM A36) at high temperatures.

Source: T.Z. Harmathy and W.W. Stanzak, ―Elevated-Temperature Tensile and Creep Properties of Some 

Structural and Prestressing Steels,‖ Fire Test Performance, STP 464, American Society for Testing and 

Materials, West Conshohocken, Pennsylvania, pp. 186-208, 1970.

Stress-strain curves for hot-rolled 220 and 420 MPa ribbed steel rebars having diameters of 10 and 16 mm that were 

tested at room temperature after exposure to temperatures up to 950 ̊C for 3 hours are presented in Figures B.3. 

Additional information presenting results for a 343 MPa 51-mm diameter steel rebar tested at temperatures up to 

600 ̊C are presented in Figure B.4. 

Figure B.3 Stress-strain curves for steel reinforcement (S220 and S420).

Source: I.B. Topçu and C. Karakurt, ―Properties of Reinforced Concrete Steel Rebars Exposed to High 

Temperatures,‖ Research Letter in Materials Science, Article ID 814137, Hindawi Publishing Co., 4 p., 

2008.
B-3

Figure B.4 Stress-strain relationships of reinforcing bars at elevated temperature.

Source: M. Takeuchi, M. Hiramoto, N. Kumagai, N. Yamazaki, A. Kodaira, and K. Sugiyama, ―Material Properties 

of Concrete and Steel Bars at Elevated Temperatures,‖ 12th International Conference on Structural 

Mechanics in Reactor Technology, Paper H04/4, pp. 133–138, Elsevier Science Publishers, North-Holland, 

The Netherlands, 1993.

The effect of elevated temperature exposure on the stress-strain curves for uncoated stress-relieved steel wire 

(ASTM A421) has been investigated [B.2]. Results are presented in Figure B.5. The specimens were tested at 

temperature about 60 to 90 minutes after heating was initiated by loading at a rate of 0.635 mm/min. Additional 

specimens were tested at rates of 1.905 and 19.05 mm/min to evaluate the rate effect. For the strain rates 

investigated, the rate did not have a considerable effect on the shape of the stress-strain curves until the temperatures 

reached 371  ̊C. Figure B.6 presents a comparison of stress-strain curves obtained at high temperature and after

Figure B.5 Effect of temperature on stress-strain curves of a prestressing steel (ASTM A421).

Source: T.Z. Harmathy and W.W. Stanzak, ―Elevated-Temperature Tensile and Creep Properties of Some 

Structural and Prestressing Steels,‖ Fire Test Performance, STP 464, American Society for Testing and 

Materials, West Conshohocken, Pennsylvania, pp. 186-208, 1970.
B-4

cooling to room temperature for 1770 MPa 5-mm-diameter low-relaxation prestressing steel wire [B.3]. Specimens 

tested at temperature were loaded at 1/mm/min until the strain reached 0.012 at which the loading rate was increased 

to 5 mm/min causing a slight step increase in the results obtained at high temperature. Specimens tested at room 

temperature were subjected to the desired test temperature for 1 hour prior to permitting them to cool to room 

temperature. A comparison of results indicates that up to 300 ̊C the results for specimens tested at temperature and 

after cooling to room temperature were similar.

Figure B.6 Stress-strain curve for 1770 MPa 5-mm-diameter 

low-relaxation prestressing steel wire tested at temperature and at room temperature.

Source: W. Zheng, Q. Hu, and H. Zhang, ―Experimental Research on the Mechanical Property of Prestressing Steel 

Wire During and After Heating,‖ Frontiers in Architecture and Civil Engineering in China 1(2), pp. 247-

254, 2007.

It has been noted that the critical temperatures for reinforcing steel and prestressing steel are 593 ̊C and 426 ̊C, 

respectively [B.4]. The critical temperature is defined as the temperature at which the reinforcement loses much of 

its strength and can no longer support applied load.

Modulus of Elasticity. Figure B.7 presents the effect of elevated temperature on the modulus of elasticity for 

structural steel and steel reinforcing bars. The results have been normalized relative to room temperature results 

(i.e., E0 = 210 x 103 MPa). These results indicate that at a temperature of 550 ̊C the modulus of elasticity is reduced 

to about half of the value at ambient temperature. Figure B.8 presents the effect of temperature on elongation and 

Young’s modulus for a 343 MPa 51-mm diameter steel rebar tested at temperatures up to 800 ̊C [B.5]. Modulus 

Figure B.7 Modulus of elasticity at elevated temperatures for structural steel and steel reinforcing bars.

Source: SFPE Engineering Guide to Performance Based Fire Protection Analysis and Design of Buildings, Society 

of Fire Protection Engineers, Bethesda, Maryland, March 2000.
B-5

Figure B.8 Influence of temperature on Young's modulus and elongation of reinforcing bars.

Source: M. Takeuchi, M. Hiramoto, N. Kumagai, N. Yamazaki, A. Kodaira, and K. Sugiyama, ―Material Properties 

of Concrete and Steel Bars at Elevated Temperatures,‖ 12th International Conference on Structural 

Mechanics in Reactor Technology, Paper H04/4, pp. 133–138, Elsevier Science Publishers, North-Holland, 

The Netherlands, 1993.

data for 1770 MPa 5-mm-diameter low-relaxation prestressing steel wire tested at temperature are presented in 

Figure B.9. The results have been normalized relative to the room-temperature modulus. Also shown in the figure 

are calculated results for modulus as a function of temperature determined from:

, (B.1)

where Es(T) is the elastic modulus of steel wire at T ̊C, T is temperature in  ̊C, and Esis the elastic modulus of steel 

wire at room temperature. Modulus results for the steel prestressing wire after permitting it to cool to room 

temperature were similar to the reference room temperature value. It was therefore recommended that the room 

temperature modulus value for prestressing wire be used as the residual modulus value after heating.

Figure B.9 Normalized modulus of elasticity results for 1770 MPa 5-mm-diameter 

low-relaxation prestressing steel wire tested at temperature.

Source: W. Zheng, Q. Hu, and H. Zhang, ―Experimental Research on the Mechanical Property of Prestressing Steel 

Wire During and After Heating,‖ Frontiers in Architecture and Civil Engineering in China 1(2), pp. 247-

254, 2007.
B-6

Yield and Ultimate Strengths. The effect of elevated temperature on the strength of a typical structural steel (ASTM 

A36) is presented in Figure B.10. Normalized yield strength results for hot-rolled and cold-worked reinforcing steel 

at temperature and after cooling to room temperature are presented in Figure B.11. The loss of yield strength is 

significant while the steel remains hot but, except for cold-worked steel, the recovery is practically complete after 

cooling to room temperature from temperatures up to 700 ̊C for the hot-rolled steel. The yield strength of coldworked steel does not recover as well as that of hot-rolled steel due to release of sessile dislocations created initially 

by the amount of prior cold working of the crystal structure [B.6]. Strength data for hot-rolled steel, cold-drawn 

Figure B.10 Effect of elevated temperature on yield and ultimate strength of a typical structural steel (ASTM A36).

Source: T.Z. Harmathy and W.W. Stanzak, ―Elevated-Temperature Tensile and Creep Properties of Some 

Structural and Prestressing Steels,‖ Fire Test Performance, STP 464, American Society for Testing and 

Materials, West Conshohocken, Pennsylvania, pp. 186-208, 1970.

Figure B.11 Yield strength of steel reinforcement vs temperature.

Source: ―Assessment of Fire-Damaged Concrete Structures and Repair by Gunite,‖ Technical Report No. 15, 

Concrete Society, London, United Kingdom May 1978.
B-7

wire or strand, and high strength alloy bars are presented in Figure B.12. Residual tensile strength and elongation at 

failure results for #4 Grade 60 ASTM A615 steel rebar heated to temperatures from 500 ̊ to 802 ̊C, held at 

temperature for 1 hour, and then permitted to slowly cool to room temperature prior to testing are presented in 

Figure B.13. Yield and ultimate strength results for hot-rolled 220 and 420 MPa ribbed steel rebars having 

Figure B.12 Strength of selected steels at high temperature.

Source: Guide for Determining the Fire Endurance of Concrete Elements, ACI 216R-81, American Concrete 

Institute, Farmington Hills, Michigan, November 1982.

Figure B.13 Effect of elevated temperature on residual yield and 

ultimate strength, and elongation of ASTM A615 steel rebar.

Source: W.T. Edwards and W.L. Gamble, ‖Strength of Grade 60 Reinforcing Bars After Exposure to Fire 

Temperatures,‖ Concrete International 8(10), pp. 17-19, American Concrete Institute, Farmington Hills, 

Michigan, October 1986.
B-8

diameters of 10 and 16 mm that were tested at room temperature after exposure to temperatures up to 950 ̊C for 

3 hours are presented in Figures B.14 and B.15, respectively. Also shown in the figures are relations developed to 

fit the data. 

Figure B.14 Effect of exposure temperature on residual yield

strength of steel reinforcement (S220 and S420).

Source: I.B. Topçu and C. Karakurt, ―Properties of Reinforced Concrete Steel Rebars Exposed to High 

Temperatures,‖ Research Letter in Materials Science, Article ID 814137, Hindawi Publishing Co., 4 p., 

2008.

Figure B.15 Effect of exposure temperature on residual tensile

strength of steel reinforcement (S220 and S420).

Source: I.B. Topçu and C. Karakurt, ―Properties of Reinforced Concrete Steel Rebars Exposed to High 

Temperatures,‖ Research Letter in Materials Science, Article ID 814137, Hindawi Publishing Co., 4 p., 

2008.

Hot and cold (residual) yield and ultimate tensile strength results for a 343 MPa 51-mm diameter steel rebar are 

presented in Figure B.16. 
B-9

Figure B.16 Yield strength and ultimate tensile strength of reinforcing bars at elevated temperature.

Source: M. Takeuchi, M. Hiramoto, N. Kumagai, N. Yamazaki, A. Kodaira, and K. Sugiyama, ―Material Properties 

of Concrete and Steel Bars at Elevated Temperatures,‖ 12th International Conference on Structural 

Mechanics in Reactor Technology, Paper H04/4, pp. 133–138, Elsevier Science Publishers, North-Holland, 

The Netherlands, 1993.

The effect of elevated temperature on the residual yield and ultimate strength of high-strength (500 MPa) weldable 

8- and 12-mm diameter steel reinforcement was investigated [B.7]. The steels were produced by the Temcore 

process (thermochemical strengthening process), microalloying with vanadium, and work hardening. Reference 

room temperature yield and ultimate strengths for the steel ranged from 531 to 640 MPa and 655 to 740 MPa, 

respectively. The steels were subjected to the designated heat-soak temperature for 1 hour and then permitted to 

slowly cool to room temperature prior to testing. The effect of temperature on the residual yield and tensile 

strengths of the steels is presented in Figure B.17. Heating of the microalloyed steel resulted in a coarsening of the

Figure B.17 Residual yield and ultimate strength of high-strength weldable rebars.

Source: J. Nikolaou and G.D. Papadimitriou, ―Microstrnctures and Mechanical Properties After Heating of 

Reinforcing 500 MPa Class Weldable Steels Produced by Various Processes (Temcore, Microalloyed With 

Vanadium and Work-Hardened),‖ Construction and Building Materials 18, pp. 243-254, 2004.
B-10

existing precipitates along the ferritic grain boundaries leading to a slight increase in yield strength at temperatures 

between 100 ̊ and 400 ̊C. Work-hardened steel bars exhibited a continuous drop in yield and ultimate strength 

starting at 200 ̊C and 300 ̊C, respectively. The weakening of mechanical properties of the Tempcore was attributed 

to tempering that occurred at high temperatures.

Yield and ultimate strength data for uncoated stress-relieved steel wire (ASTM A421) tested at temperature are 

provided in Figure B.18.

Figure B.18 Effect of elevated temperature on yield and ultimate strength of a prestressing steel (ASTM A421).

Source: T.Z. Harmathy and W.W. Stanzak, ―Elevated-Temperature Tensile and Creep Properties of Some 

Structural and Prestressing Steels,‖ Fire Test Performance, STP 464, American Society for Testing and 

Materials, West Conshohocken, Pennsylvania, pp. 186-208, 1970.

Yield and tensile strength data for 1770 MPa 5-mm-diameter low-relaxation prestressing steel wire tested at 

temperature are presented in Figure B.19. The results have been normalized relative to the room-temperature 

values. Also shown in the figure are calculated results for yield and tensile strengths as a function of temperature 

[B.3]. Yield strength was determined from:

, (B.2)

where f0.2 (T) is yield strength of steel wire at T  ̊C, f0.2 is the yield strength of steel wire at room temperature, and T 

is temperature in  ̊C. Tensile strength was determined from:

, (B.3)

where fb(T) is the tensile strength limit of steel wire at T  ̊C, fb is the tensile strength limit at room temperature and T 

is the temperature in  ̊C. Residual yield and tensile strength data for the 1770 MPa 5-mm-diameter low-relaxation 
B-11

Figure B.19 Normalized yield and tensile strength results for 1770 MPa 5-mm-diameter 

low-relaxation prestressing steel wire tested at temperature.

Source: W. Zheng, Q. Hu, and H. Zhang, ―Experimental Research on the Mechanical Property of Prestressing Steel 

Wire During and After Heating,‖ Frontiers in Architecture and Civil Engineering in China 1(2), pp. 247-

254, 2007.

prestressing steel wire obtained at room temperature are presented in Figure B.20. The results have been 

normalized relative to the room-temperature values. Also shown in the figure are calculated results for the residual 

yield and tensile strengths as a function of exposure temperature [B.3]. Yield strength was determined from:

0 ̊C<T<300 ̊C, (B.4)

300 ̊C<T<400 ̊C, (B.5)

400 ̊C<T<700 ̊C, (B.6)

where is yield strength of steel wire after T  ̊C and T is temperature in  ̊C. Tensile strength was determined 

from:

, 0 ̊C<T<300 ̊C, (B.7)

300 ̊C<T<400 ̊C, (B.8)

400 ̊C<T<700 ̊C (B.9)

where is the tensile strength of steel wire after T  ̊C and T is the temperature in  ̊C.

The relationship between and tensile strength normalized to the room-temperature value and temperature for a colddrawn 7-wire prestressing strand is shown in Figure B.21. Also shown in the figure are normalized results obtained 

for a 25.4-mm diameter hot-rolled reinforcing steel alloy bar. It has been noted that special care must be taken when 

dealing with prestressing steel due to loss of tension caused by relaxation effects when creep occurs, with about 50% 

of normal yield strength likely achieved at 400 ̊C [B.8]. Additional results for 30 types of prestressing steels having 

strengths from 1500 to 2200 MPa and tested at temperature are presented in Figure B.22.
B-12

Figure B.20 Normalized residual yield and tensile strength results for 1770 MPa 

5-mm-diameter low-relaxation prestressing steel wire tested after heating.

Source: W. Zheng, Q. Hu, and H. Zhang, ―Experimental Research on the Mechanical Property of Prestressing Steel 

Wire During and After Heating,‖ Frontiers in Architecture and Civil Engineering in China 1(2), pp. 247-

254, 2007.

Figure B.21 Relationship between temperature and normalized tensile strength 

for hot-rolled reinforcing steel bar and cold-drawn prestressing steel wire.

Source: M.S. Abrams and C.R. Cruz, ―The Behavior at High Temperature of Steel Strand for Prestressed 

Concrete,‖ Journal of PCA Research and Development Laboratories 3(3), pp. 8-19, September 1961. 
B-13

Figure B.22 Tensile strength results of different prestressing steels in the heated state.

Source: U. Schneider, C. Diererichs, and C. Ehm, ―Effect of Temperature on Steel and Concrete for PCRV’s,‖ 

Nuclear Engineering and Design 67, pp. 245–258, 1981.

Toughness. Residual toughness results for hot-rolled 220 and 420 MPa ribbed steel rebars having diameters of 10 

and 16 mm that were tested at room temperature after exposure to temperatures up to 950 ̊C for 3 hours are 

presented in Figure B.23. Also shown in the figure are relations developed to fit the data. Results show that 

toughness values for both steel types decreased after elevated temperature exposure. However up to about 300 ̊C 

the toughness values increased somewhat due to the increased ductility of the steels. 

Figure B.23 Effect of exposure temperature on toughness of steel reinforcement (S220 and S420).

Source: I.B. Topçu and C. Karakurt, ―Properties of Reinforced Concrete Steel Rebars Exposed to High 

Temperatures,‖ Research Letter in Materials Science, Article ID 814137, Hindawi Publishing Co., 4 p., 

2008.

The effect of elevated temperature on the residual toughness (Chapry V-notch energy) for high-strength (500 MPa) 

weldable 8- and 12-mm diameter steel reinforcement was investigated [B.7]. The steels were produced by the 

Tempcore process (thermochemical strengthening process), microalloying with vanadium, and work hardening.
B-14

Reference room temperature yield and ultimate strengths for the steel ranged from 531 to 640 MPa and 655 to 740 

MPa, respectively. The steels were subjected to the designated heat-soak temperature for 1 hour and then permitted 

to slowly cool to room temperature prior to testing. Figure B.24 presents residual normalized Chapry V-notch 

absorbed energy of the steels after heating.

Figure B.24 Residual normalized absorbed energy of high-strength weldable rebars.

Source: J. Nikolaou and G.D. Papadimitriou, ―Microstrnctures and Mechanical Properties After Heating of 

Reinforcing 500 MPa Class Weldable Steels Produced by Various Processes (Temcore, Microalloyed With 

Vanadium and Work-Hardened),‖ Construction and Building Materials 18, pp. 243-254, 2004.

B.2 Physical Properties

Information on the density, mean specific heat, thermal conductivity, thermal diffusivity, and coefficient of thermal 

expansion of different steels is presented in Figures B.25-B.29, respectively [B.8]. Average thermal expansion data 

Figure B.25 Density of different steels.

Source: U. Schneider, C. Diererichs, and C. Ehm, ―Effect of Temperature on Steel and Concrete for PCRV’s,‖ 

Nuclear Engineering and Design 67, pp. 245–258, 1981.
B-15

Figure B.26 Mean specific heat of different steels.

Source: U. Schneider, C. Diererichs, and C. Ehm, ―Effect of Temperature on Steel and Concrete for PCRV’s,‖ 

Nuclear Engineering and Design 67, pp. 245–258, 1981.

Figure B.27 Thermal conductivity of different steels.

Source: U. Schneider, C. Diererichs, and C. Ehm, ―Effect of Temperature on Steel and Concrete for PCRV’s,‖ 

Nuclear Engineering and Design 67, pp. 245–258,1981.
B-16

Figure B.28 Thermal diffusivity of different steels.

Source: U. Schneider, C. Diererichs, and C. Ehm, ―Effect of Temperature on Steel and Concrete for PCRV’s,‖ 

Nuclear Engineering and Design 67, pp. 245–258, 1981.

Figure B.29 Coefficient of expansion of different steels.

Source: U. Schneider, C. Diererichs, and C. Ehm, ―Effect of Temperature on Steel and Concrete for PCRV’s,‖ 

Nuclear Engineering and Design 67, pp. 245–258, 1981. 
B-17

for ferritic steels over the temperature range of 200 ̊ to 650 ̊C is presented in Figure B.30. For this data the 

coefficient of thermal expansion can be approximated by [B.9]:

= (11 + 0. ) x 10-6/ ̊C

where is the temperature in degrees C.

Figure B.30 Thermal expansion of ferritic steel at high temperature.

Source: Manual of steel Construction, 7th Edition, American Institute of Steel Construction, New York, New York, 

1970.

B.3 Examples of Constitutive Relations at Elevated Temperature

Relations for the effect of elevated temperature on modulus of elasticity of rebars have been proposed [B.10]. The 

elastic modulus at elevated temperature can be estimated from:

(B.10)

where Es

T

and Es are the elastic modulus of rebar at elevated and room temperature, respectively. The elastic

modulus after high temperature exposure can be estimated from [B.10]:

. (B.11)

Stress-strain relationships for rebars at high temperatures have been proposed [B.10]:

(B.12)

where a and b are parameters related to different types of rebar, and fy is the yield strength of the rebar at room 

temperature. A relationship has also been proposed for the nominal yield strength of high strength wires at high 

temperatures [B.11]:
B-18

20 ̊ < T < 300 ̊C (B.13)

300 ̊ < T < 600 ̊C (B.14)

where and f0.2 are the nominal yield strength corresponding to 0.2% residual strain at elevated and room 

temperature, respectively.

Eurocode ENV 1993-1-2 provides the following relation for thermal conductivity ( a) (W/m  ̊C), thermal diffusivity 

(aa) (m2/hr), and specific heat (ca) (J/kg ̊C) of steel [B.12]:

 for a ≤ 800 ̊C, (B.15)

 for a > 800 ̊C, (B.16)

 for a ≤ 750 ̊C, (B.17)

ca = 425 + 0.773 a - 1.69 x 10-3

a

2

+ 2.22 x 10-6

a

3 for 20 ̊ < a ≤ 600 ̊C, (B.18)

 for 600 ̊ < a ≤ 735 ̊C, (B.19)

 for 735 ̊ < a ≤ 900 ̊C, and (B.20)

ca = 650, for 900 ̊ < a ≤ 1200 ̊C. (B.21)

B.4 References

B.1 Properties of Materials at High Temperatures – Steel, Y. Anderberg (Editor), Report LUTVDG/(TVBB2008), Lund Institute of Technology, Sweden, February 1983.

B.2 T.Z. Harmathy and W.W. Stanzak, ―Elevated-Temperature Tensile and Creep Properties of Some 

Structural and Prestressing Steels,‖ Fire Test Performance, STP 464, American Society for Testing and 

Materials, West Conshohocken, Pennsylvania, pp. 186-208, 1970.

B.3 W. Zheng, Q. Hu, and H. Zhang, ―Experimental Research on the Mechanical Property of Prestressing Steel 

Wire During and After Heating,‖ Frontiers in Architecture and Civil Engineering in China 1(2), pp. 247-

254, 2007.

B.4 Structural Fire Protection, Manuals and Reports on Engineering Practice, No. 78, T.T. Lie (Editor), 

American Society of Civil Engineers, New York, New York, 1992.

B.5 M. Takeuchi, M. Hiramoto, N. Kumagai, N. Yamazaki, A. Kodaira, and K. Sugiyama, ―Material Properties 

of Concrete and Steel Bars at Elevated Temperatures,‖ 12th International Conference on Structural 

Mechanics in Reactor Technology, Paper H04/4, pp. 133–138, Elsevier Science Publishers, North-Holland, 

The Netherlands, 1993.

B.6 P.D. Morley and R. Royles, ―The Influence of High Temperature on the Bond in Reinforced Concrete,‖ 

Fire Safety Journal 2, pp. 243-255, 1979/1980.

B.7 J. Nikolaou and G.D. Papadimitriou, ―Microstructures and Mechanical Properties After Heating of 

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