Irradiation-Assisted Stress 

Corrosion Cracking of 

Austenitic Stainless Steels in 

BWR Environments

Office of Nuclear Regulatory Research 

NUREG/CR-7018 

ANL-09/17
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Irradiation-Assisted Stress 

Corrosion Cracking of 

Austenitic Stainless Steels in 

BWR Environments 

Manuscript Completed: February 2009 

Date Published: June 2010 

 

Prepared by 

Y. Chen, O.K. Chopra, E.E. Gruber and W.J. Shack 

Argonne National Laboratory 

9700 South Cass Avenue 

Argonne, IL 60439 

A.S. Rao, NRC Project Manager 

NRC Job Code N6519 

Office of Nuclear Regulatory Research 

NUREG/CR-7018 

ANL-09/17 
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Abstract

The internal components of light water reactors are exposed to high-energy neutron irradiation and 

high-temperature reactor coolant. The exposure to neutron irradiation increases the susceptibility of 

austenitic stainless steels (SSs) to stress corrosion cracking (SCC) because of the elevated corrosion 

potential of the reactor coolant and the introduction of new embrittlement mechanisms through radiation 

damage. Various nonsensitized SSs and nickel alloys have been found to be prone to intergranular 

cracking after extended neutron exposure. Such cracks have been seen in a number of internal 

components in boiling water reactors (BWRs). The elevated susceptibility to SCC in irradiated materials, 

commonly referred to as irradiation-assisted stress corrosion cracking (IASCC), is a complex 

phenomenon that involves simultaneous actions of irradiation, stress, and corrosion. In recent years, as 

nuclear power plants have aged and irradiation dose increased, IASCC has become an increasingly

important issue. Post-irradiation crack growth rate and fracture toughness tests have been performed to 

provide data and technical support for the NRC to address various issues related to aging degradation of 

reactor-core internal structures and components. This report summarizes the results of the last group of 

tests on compact tension specimens from the Halden-II irradiation. The IASCC susceptibility of 

austenitic SSs and heat-affected-zone (HAZ) materials sectioned from submerged arc and shielded metal 

arc welds was evaluated by conducting crack growth rate and fracture toughness tests in a simulated 

BWR environment. The fracture and cracking behavior of HAZ materials, thermally sensitized SSs and 

grain-boundary engineered SSs was investigated at several doses (≤3 dpa). These latest results were 

combined with previous results from Halden-I and II irradiations to analyze the effects of neutron dose, 

water chemistry, alloy compositions, and welding and processing conditions on IASCC. The effect of 

neutron irradiation on the fracture toughness of austenitic SSs was also evaluated at dose levels relevant 

to BWR internals. 

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the requirements of the Paperwork Reduction Act of 1995 (44 U.S.C. 3501 et seq.).

Public Protection Notification

The NRC may not conduct or sponsor, and a person is not required to respond to, a request for 

information or an information collection requirement unless the requesting document displays a current 

valid OMB control number.
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Foreword

The internal components of light water reactors are exposed to high-energy neutron irradiation and hightemperature reactor coolant. The exposure to neutron irradiation increases the susceptibility of austenitic 

stainless steels (SSs) to irradiation-assisted stress corrosion cracking (IASCC). Similarly, the radiation 

damage may induce embrittlement and reduce the material’s resistance to crack propagation. Various 

non-sensitized SSs and nickel alloys have been shown to be susceptible to IASCC and embrittlement after 

extended neutron exposure in laboratory environments. Additionally, cracks caused by IASCC have been 

found in a number of internal components in boiling water reactors (BWRs). As nuclear power plants age 

and neutron irradiation dose increases, IASCC and embrittlement become more likely and are an 

important safety consideration in the long-term operation of nuclear power plants. 

IASCC is a complex phenomenon that results from the interaction of irradiation, stress, and corrosion. 

This report examines the effects of simulated light-water reactor environments, material chemistry, and 

irradiation damage on the IASCC susceptibility of a series of commercially available and laboratorymelted stainless steels. This report is the final report in its series dating back approximately 8 years. 

Some of the earlier publications in this series include NUREG/CR 5608, “Irradiation-Assisted Stress 

Corrosion Cracking of Model Austenitic Stainless Steels Irradiated in the Halden Reactor”; NUREG/CR6892, “Fracture Toughness and Crack Growth Rates of Irradiated Austenitic Stainless Steels”; 

NUREG/CR-6687, “Irradiation-Assisted Stress Corrosion Cracking of Model Austenitic Stainless Steel 

Alloys”; NUREG/CR- 6915, “Irradiation-Assisted Stress Corrosion Cracking of Austenitic Stainless 

Steels and Alloy 690 from Halden Phase-II Irradiations” and NUREG/CR-6960, “Crack Growth Rates 

and Fracture Toughness of Irradiated Austenitic Stainless Steels in BWR Environments.”

The Halden - Phase I irradiation principally evaluated stainless steels having a wide-range of chemical 

compositions (including commercial steels of typical chemistry) with conventional heat treatment and 

product form processing. The Halden – Phase II irradiation included a plurality of innovatively fabricated 

and engineered alloys intended to be more resistant to IASCC. Specially, these irradiations studied 

austenitic SSs, such as Types 304 and 316 (and their low carbon counterparts), grain boundary engineered 

(GBE) SSs, submerged arc and shield metal arc welds, and the heat-affected zone (HAZ) materials 

sectioned from these welds. The IASCC susceptibility of these materials was evaluated by conducting 

crack growth rate and fracture toughness tests in simulated BWR environments (fluence range up to 3 

dpa; and dissolved oxygen in the water 200 ppb to 8 ppm).

In NUREG/CR-6892, it was found that sulfur (S) content below 0.002 wt. % provides the best resistance 

to IASCC in SSs. In contrast, IASCC susceptibility increased dramatically in SSs that contain greater 

than 0.003 wt. % S. A sufficiently low S concentration is the most important factor for ensuring that a SS 

has good IASCC resistance. The earlier test data also suggest that moderately high carbon content (i.e., 

>0.03 wt%) ensures good resistance to IASCC content.. From the above two observations, it is 

reasonable to infer that both moderately high carbon (> 0.03 wt. %) and low sulfur (< 0.002 wt. %) are 

necessary to provide sufficient IASCC resistance in Type 304 and 316 steels. This is an important 

finding, given that many stainless steels (Type 304L or 316L) are decarburized to protect against thermal 

sensitization; too low a carbon level (<0.03 wt.%) could render such steels susceptible to IASCC. 

In this report, it is demonstrated that although highly irradiated SSs exhibit strain softening, the effect on 

the plastic zone size ahead of a crack tip is only minimally increased. The increase in yield strength due 

to irradiation more than compensates for this effect such that valid crack growth rate and fracture 

toughness measurements are achieved in small test specimens. Results also show that, for a Type 304 SS 
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sensitized at 600oC for 10.5 and 24 hours, the degree of sensitization has a negligible effect on IASCC 

resistance. Also, the GBE treatment (the fraction of coincident-site-lattice boundaries ~ 60%) did not 

improve IASCC susceptibility in the test materials and environment. Conversely, increasing the neutron 

fluence significantly decreased both the IASCC resistance and the fracture toughness.

The research described in this report provides relevant test data and technical support for NRC to analyze 

the integrity of reactor-core internal structures and components which are subjected to IASCC. 

Specifically, results of this final report can be used to support the evaluation of the acceptability of reactor 

internal aging management programs in license renewal applications. The results may also be used to 

determine appropriate inspection and flaw disposition procedures for reactor internals that can be used in 

ASME code development and to support any needed conditions within Title 10, Section 50.55a, “Codes 

and Standards,” of the Code of Federal Regulations (10CFR 5055a). Finally, the results are also useful 

for identifying issues related to long-term operation (i.e., greater than 60 years) in nuclear power plants.

Michael Case, Director

Division of Engineering

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission
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Contents 

Abstract................................................................................................................................................. iii

Foreword............................................................................................................................................... v

Figures .................................................................................................................................................. ix

Tables.................................................................................................................................................... xv

Executive Summary.............................................................................................................................. xvii

Acknowledgments................................................................................................................................. xxi

Acronyms and Abbreviations ............................................................................................................... xxiii

1 Introduction................................................................................................................................. 1

2 Experimental............................................................................................................................... 5

2.1 Compact Tension Specimens and Materials..................................................................... 5

2.2 Irradiation ......................................................................................................................... 7

2.3 Test Facility...................................................................................................................... 7

2.4 Test Procedures................................................................................................................. 10

2.4.1 Crack Growth Rate Test ......................................................................................... 10

2.4.2 Model for Crack Growth Rate Test Analysis......................................................... 11

2.4.3 Fracture Toughness J-R Curve Tests...................................................................... 12

3 Experimental Results.................................................................................................................. 15

3.1 Thermally Treated Type 304L SA Weld HAZ Irradiated to 2.4 dpa (GG3TA-TT)......... 15

3.2 Type 304 SMA Weld HAZ .............................................................................................. 20

3.2.1 As-welded Type 304 SMA Weld HAZ Irradiated to 0.75 dpa (85-7B) ................. 20

3.2.2 Thermally Treated Type 304 SMA Weld HAZ Irradiated to 2 dpa (85-2ATT)..... 24

3.3 Sensitized Type 304 SS Irradiated to 0.75 dpa................................................................. 30

3.3.1 Type 304 SS Sensitized at 600°C for 10.5 h (85-1TT)........................................... 30
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3.3.2 Type 304 SS Sensitized at 600°C for 24 h (85-5A1TT)......................................... 34

3.4 Type 304 SS with and without GBE Treatment ............................................................... 38

3.4.1 Type 304 SS without GBE Treatment (IT304-01) ................................................. 38

3.4.2 Type 304 SS with GBE Treatment (GB304-01)..................................................... 42

3.5 Type 316 SS with and without GBE Treatment ............................................................... 46

3.5.1 Type 316 without GBE Treatment (IT316-01)....................................................... 46

3.5.2 Type 316 SS with GBE Treatment (GB316-01)..................................................... 50

4 Discussion................................................................................................................................... 55

4.1 Specimen Size Criteria Relevant to Cyclic and Constant-Load CGR Tests..................... 55

4.2 Effects of Weld Process on Cracking Behavior ............................................................... 57

4.2.1 Cyclic CGR for Nonirradiated Weld HAZs ........................................................... 57

4.2.2 Cyclic CGR for Irradiated Weld HAZ Materials.................................................... 58

4.2.3 Constant-Load CGR for Nonirradiated and Irradiated Weld HAZ Specimens...... 59

4.2.4 Fracture Toughness of Weld HAZ Specimens....................................................... 61

4.3 Effect of Sensitization ...................................................................................................... 62

4.4 Effect of Material Types................................................................................................... 64

4.5 Effect of Water Chemistry................................................................................................ 68

4.6 Effect of Neutron Dose..................................................................................................... 71

4.7 Effect of GBE Treatment.................................................................................................. 73

4.7.1 Crack Growth Rates for GBE and Non-GBE Type 304 and 316 SSs.................... 73

4.7.2 Fracture Toughness of GBE and Non-GBE Type 304 and 316 SSs....................... 75

4.7.3 Hardness and Microstructure of Nonirradiated GBE and Non-GBE Type 304 

and 316 SSs ............................................................................................................ 75

5 Summary ..................................................................................................................................... 81

References............................................................................................................................................. 83
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Figures

1. Susceptibility of irradiated austenitic SSs to intergranular SCC as a function of fluence in 

high-DO water. Data from slow-strain-rate tensile tests............................................................ 2

2. Configuration of compact-tension specimen used in this study (dimensions in mm)................. 5

3. Weld microstructure in the (a) Grand Gulf SA weld and (b) laboratory-prepared SMA weld... 6

4. Mechanical test system for conducting CGR and J-R tests on radioactive specimens. .............. 8

5. A schematic diagram of the water recirculation system. ............................................................ 9

6. Crack-length-vs.-time plot for the thermally treated GG 304L SA HAZ specimen (≈2.4 dpa) 

in simulated BWR water at 289°C.............................................................................................. 16

7. Crack growth rate for GG 304L SA HAZ specimen under cyclic load in simulated BWR 

NWC water at ≈290°C................................................................................................................ 17

8. Crack growth rate versus stress intensity under constant-loading condition for the 304L SA 

HAZ specimen thermally treated at 500°C for 24 h (GG3TA-TT) in simulated BWR NWC 

and HWC water at ≈290°C. ........................................................................................................ 18

9. Load vs. load-line displacement curve for the thermally treated 304L SA HAZ specimen 

(GG3TA-TT) tested in NWC at 289°C....................................................................................... 19

10. Fracture toughness J-R result for the thermally treated 304L SA HAZ specimen (GG3TATT) tested in NWC at 289°C....................................................................................................... 19

11. Photograph of the fracture surface of the thermally treated 304L SA HAZ specimen 

(GG3TA-TT)............................................................................................................................... 20

12. Crack-length-vs.-time plot for the as-welded Type 304 SMA HAZ specimen (85-7B, ~0.75 

dpa) in BWR water at 289°C. ..................................................................................................... 21

13. Cyclic crack growth rate for the as-welded Type 304 SMA HAZ specimen (85-7B) irradiated 

to 0.75 dpa................................................................................................................................... 22

14. Crack growth rate versus stress intensity under constant-loading conditions for the as-welded 

Type 304 SMA HAZ specimen (85-7B) irradiated to 0.75 dpa.................................................. 23

15. Temperature dependence of crack growth rate for the as-welded 304 SMA HAZ specimen 

(85-7B, 0.75 dpa)........................................................................................................................ 23

16. Load vs. load-line displacement curve for the as-welded Type 304 SMA HAZ specimen (85-

7B) tested in NWC at 289°C....................................................................................................... 24
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17. Fracture toughness J-R curves for the as-welded Type 304 SMA HAZ specimen (85-7B) 

tested in NWC at 289°C.............................................................................................................. 24

18. Photograph of the fracture surface of the as-welded 304 SMA HAZ specimen (85-7B) 

irradiated to 0.75 dpa. ................................................................................................................. 24

19. Crack-length-vs.-time plot for the thermally treated Type 304 SMA HAZ specimen (85-

2ATT, ~2 dpa) in BWR water at 289°C. .................................................................................... 26

20. Cyclic CGRs for the thermally treated Type 304 SS SMA weld HAZ specimen (85-2ATT) 

irradiated to 2.1 dpa. ................................................................................................................... 27

21. Constant-load CGRs for the thermally treated Type 304 SS SMA weld HAZ specimen (85-

2ATT) irradiated to 2.1 dpa. ....................................................................................................... 28

22. (a) Load vs. load-line displacement and (b) fracture toughness J-R curve for the thermally 

treated 304 SMA weld HAZ specimen (85-2ATT). ................................................................... 28

23. Optical images of the thermally treated 304 SS SMA HAZ specimen (85-2ATT) After 

fracturing: (a) fracture surface and (b) side view........................................................................ 29

24. Crack-length-vs.-time plot for the Type 304 SS specimen sensitized at 600°C for 10.5 h (85-

1TT, ~0.75 dpa) in BWR water at 289°C. .................................................................................. 30

25. Cyclic CGR for the Type 304 specimen sensitized at 600°C for 10.5 h (85-1TT) and 

irradiated to 0.75 dpa. ................................................................................................................. 32

26. Crack growth rate versus stress intensity under constant-loading conditions for the Type 304 

specimen sensitized at 600°C for 10.5 h (85-1TT) and irradiated to 0.75 dpa. .......................... 32

27. Load vs. load-line displacement curve for the Type 304 specimen sensitized at 600°C for 

10.5 h (85-1TT) and tested in NWC at 289°C. ........................................................................... 33

28. Fracture toughness J-R curve for the Type 304 specimen sensitized at 600°C for 10.5 h (85-

1TT) and tested in NWC at 289°C.............................................................................................. 33

29. Photograph of the fracture surface of the Type 304 specimen sensitized at 600°C for 10.5 h 

(85-1TT) and irradiated to 0.75 dpa............................................................................................ 34

30. Crack-length-vs.-time plot for the Type 304 SS specimen sensitized at 600°C for 24 h (85-

5A1TT, ~0.75 dpa) in BWR water.............................................................................................. 35

31. Cyclic CGR for the Type 304 specimen sensitized at 600°C for 24 h (85-5A1TT) and 

irradiated to 0.75 dpa. ................................................................................................................. 36

32. Crack growth rate versus stress intensity under constant-loading conditions for the Type 304 

specimen sensitized at 600°C for 24 h (85-5A1TT) and irradiated to 0.75 dpa. ........................ 37
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33. Photograph of the fracture surface of the Type 304 specimen sensitized at 600°C for 24 h 

(85-5A1TT) and irradiated to 0.75 dpa....................................................................................... 37

34. Temperature dependence of CGR for the Type 304 specimen sensitized at 600°C for 24 h 

(85-5A1TT, 0.75 dpa)................................................................................................................. 38

35. Crack-length-vs.-time plot for the base heat of Type 304 SS specimen (IT304-01, ~2 dpa) in 

BWR water at 289°C. ................................................................................................................. 39

36. Cyclic CGR for the base heat of Type 304 SS specimen (IT304-01) irradiated to 2 dpa........... 40

37. Crack growth rate versus stress intensity under constant-loading conditions for the base heat 

of Type 304 SS specimen (IT304-01) irradiated to 2 dpa........................................................... 41

38. Load vs. load-line displacement curve for the base heat of Type 304 SS specimen (IT304-01) 

tested in NWC at 289°C.............................................................................................................. 41

39. Fracture toughness J-R curves for the base heat of Type 304 SS specimen (IT304-01) tested 

in NWC at 289°C........................................................................................................................ 42

40. Photograph of the fracture surface of the base heat of Type 304 SS (IT304-01) irradiated to 

2 dpa............................................................................................................................................ 42

41. Crack-length-vs.-time plot for the GBE-treated Type 304 SS specimen (GB304-01, ≈2 dpa) 

in BWR water at 289°C. ............................................................................................................. 43

42. Cyclic crack growth rate for the GBE-treated Type 304 SS specimen (GB304-01) irradiated 

to 2 dpa........................................................................................................................................ 44

43. Crack growth rate versus stress intensity under constant-loading conditions for the GBEtreated Type 304 SS specimen (GB304-01) irradiated to 2 dpa. ................................................ 44

44. Load vs. load-line displacement curve for the GBE-treated Type 304 SS (GB304-01, ~2 dpa) 

tested in NWC at 289°C.............................................................................................................. 45

45. Fracture toughness J-R curves for the GBE-treated Type 304 SS (GB304-01, ~2 dpa) tested 

in NWC at 289°C........................................................................................................................ 45

46. Photograph of the fracture surface of the GBE-treated Type 304 SS (GB304-01) irradiated to 

2 dpa............................................................................................................................................ 45

47. Crack-length-vs.-time plot for the base heat of Type 316 SS specimen (IT316-01, ≈2 dpa) in 

BWR water at 289°C. ................................................................................................................. 47

48. Cyclic CGR for the base heat of Type 316 SS (IT316-01) irradiated to 2 dpa........................... 48

49. Crack growth rate versus stress intensity under constant-loading conditions for the base heat 

of Type 316 SS (IT316-01) irradiated to 2 dpa........................................................................... 48
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50. Load vs. load-line displacement curve for the base heat of Type 316 SS (IT316-01) tested in 

NWC at 289°C............................................................................................................................ 49

51. Fracture toughness J-R curve for the base heat of Type 316 SS (IT316-01) tested in NWC at 

289°C. ......................................................................................................................................... 49

52. Photograph of the fracture surface of the base heat of Type 316 SS (IT316-01) irradiated to 2

dpa............................................................................................................................................... 49

53. Crack-length-vs.-time plot for the GBE-treated Type 316 SS (GB316-01, ≈2 dpa) tested in 

BWR water at 289°C. ................................................................................................................. 51

54. Cyclic CGR for the GBE-treated Type 316 SS (GB316-01) irradiated to 2 dpa. ....................... 52

55. Crack growth rate versus stress intensity under constant-loading conditions for the GBEtreated Type 316 SS (GB316-01) irradiated to 2 dpa.................................................................. 52

56. Load vs. load-line displacement curve for the GBE-treated Type 316 SS (GB316-01, ≈2 dpa) 

tested in NWC at 289°C.............................................................................................................. 53

57. Fracture toughness J-R curves for the GBE-treated Type 316 SS (GB316-01, ≈2 dpa) tested 

in NWC at 289°C........................................................................................................................ 53

58. Photograph of the fracture surface of the GBE-treated Type 316 SS (GB316-01) irradiated to 

2 dpa............................................................................................................................................ 53

59. Stress distribution ahead of a two-dimensional crack in a work-softened material.................... 56

60. Cyclic CGRs for non-irradiated SA and SMA weld HAZ specimens under the as-welded and 

thermally treated conditions........................................................................................................ 57

61. Cyclic CGRs for SA and SMA weld HAZ specimens under as-welded condition irradiated to 

(a) 0.75 dpa, and (b) ≈2 dpa........................................................................................................ 58

62. Cyclic CGRs for thermally treated and as-welded (a) SMA HAZ specimens at 0.75 dpa and 

(b) SA HAZ specimens at 2 dpa. ................................................................................................ 59

63. Constant-load CGR versus stress intensity for nonirradiated SA and SMA weld HAZ 

specimens with and without heat treatment. ............................................................................... 60

64. Constant-load CGR results versus stress intensity for SA and SMA HAZ specimens 

irradiated to (a) 0.75 dpa and (b) ≈2 dpa..................................................................................... 60

65. Fracture toughness of irradiated Type 304L SA and 304 SMA weld HAZ specimens tested in 

NWC BWR water. ...................................................................................................................... 62

66. Cyclic CGRs for sensitized Type 304 SS irradiated to 0.75 and 2 dpa....................................... 63
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67. Constant-load CGRs versus stress intensity of sensitized Type 304 SS irradiated to 0.75 and 

2 dpa............................................................................................................................................ 63

68. Cyclic CGRs for austenitic SSs irradiated to (a) 0.45 dpa, (b) 1.35 dpa, and (c) 3 dpa.............. 65

69. Constant-load CGRs versus stress intensity for austenitic SSs irradiated to (a) 0.45 dpa, (b) 

1.35 dpa, and (c) 3 dpa................................................................................................................ 66

70. Cyclic CGRs for Type 304L and 316L SSs in NWC and HWC tests. ....................................... 68

71. Constant-load CGRs versus stress intensity in NWC and HWC for (a) weld HAZ, (b) 

sensitized 304 SS, and (c) wrought SSs...................................................................................... 69

72. Dose dependence of cyclic CGRs for Type 304L and 316 SSs.................................................. 71

73. Constant-load CGRs versus stress intensity for Type 304L, 316 and 316L SSs at different 

doses............................................................................................................................................ 72

74. Dose dependence of fracture toughness for austenitic SSs. Dashed lines represent the scatter 

band for the fast reactor data at 350-450°C. ............................................................................... 73

75. Cyclic CGRs for irradiated non-GBE and GBE treated (a) Type 304 SSs and (b) Type 316 

SSs tested in BWR NWC............................................................................................................ 74

76. Constant-load CGRs versus stress intensity for non-GBE and GBE treated Type 304 and 316 

SSs............................................................................................................................................... 74

77. Fracture toughness J-R curves for non-GBE and GBE (a) Type 304 SSs and (b) Type 316 

SSs............................................................................................................................................... 75

78. Hardness measurements of non-GBE and GBE Type 304 and 316 SSs..................................... 76

79. Grain morphology revealed by electron backscattered diffraction. ............................................ 76

80. Grain size distribution (left) and CSL length fraction (right) for GBE and non-GBE Type 304 

and 316 SSs................................................................................................................................. 77

81. Transmission electron microscopy images for (a) non-GBE and (b) GBE Type 316 SSs. ........ 78

82. TEM images for (a) non-GBE and (b) GBE 304 SSs. ................................................................ 79
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Tables

1. Composition for the Halden-II irradiation materials................................................................... 5

2. Fractions of CSL boundaries in Types 304 and 316 SSs with and without GBE processing. .... 6

3. Irradiation conditions for the Halden-II CT specimens............................................................... 7

4. Halden phase II specimens and type of tests performed. ............................................................ 15

5. Test results for Specimen GG3TA-TT of Type 304L SS weld HAZ in BWR water at 289°C. . 16

6. Test results for the Type 304 SMA HAZ specimen at as-welded condition (85-7B, 0.75 dpa) 

in BWR water.............................................................................................................................. 21

7. Test results for a thermally treated 304 SMA HAZ specimen irradiated to 2 dpa (85-2ATT). .. 25

8. Test results for a Type 304 specimen sensitized at 600°C for 10.5 h (85-1TT, 0.75 dpa) in 

BWR water.................................................................................................................................. 30

9. Test results for a Type 304 specimen sensitized at 600°C for 24 h (85-5A1TT, 0.75 dpa) in 

BWR water.................................................................................................................................. 34

10. Test results for the base heat of Type 304 specimen (IT304, ≈2 dpa) in BWR water................ 39

11. Test results for the GBE-treated Type 304 specimen (GB304-01, ≈2 dpa) in BWR water. ....... 43

12. Test results for the base heat of Type 316 specimen (IT316-01, ≈2 dpa) in BWR water........... 46

13. Test results for the GBE-treated Type 316 specimen (GB316-01, ≈2 dpa) in BWR water. ....... 50

14. Composition of austenitic stainless steels (wt%)........................................................................ 64
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Executive Summary

Stress corrosion cracking (SCC) is an important degradation mechanism for austenitic stainless 

steels (SSs) in high-temperature aqueous environments. Premature failures resulting from SCC are 

commonly found in thermally sensitized austenitic SSs under corrosive environments. In light water 

reactors (LWRs), reactor core internal components are exposed to high-temperature coolant and also 

subjected to high-energy irradiation by fission neutrons. The exposure to fast neutron irradiation can 

further increase the susceptibility of SSs to SCC by elevating the corrosion potential of the reactor coolant 

and by introducing new embrittlement mechanisms through radiation damage. Various nonsensitized 

austenitic SSs and nickel alloys have been found to be prone to intergranular cracking after extended 

irradiation exposure. The elevated susceptibility of irradiated materials to SCC, commonly termed

irradiation-assisted stress corrosion cracking (IASCC), has become an increasingly important issue in 

recent years as nuclear power plants continue to age and irradiation dose increases. A comprehensive 

understanding of the degradation mechanism of IASCC is important for determining the structural

integrity of reactor internal components and the safety and economy of aging nuclear power plants. 

Post-irradiation tests of crack growth and fracture toughness have been performed on various 

austenitic SSs to provide data and technical support for addressing age-related degradation of reactor-core 

internal structures and components. In the present study, crack growth rate (CGR) tests were conducted 

in simulated boiling water reactor (BWR) environments on 1/4 inch thick compact tension (1/4T-CT)

specimens. Heat-affected-zone (HAZ) materials sectioned from Type 304L submerged arc (SA) and 

Type 304 shielded metal arc (SMA) welds, sensitized Type 304 SSs, and Type 304 and 316 SSs with and 

without grain boundary engineering (GBE) treatment were included in the study. The specimens were 

irradiated in the Halden reactor to 0.75 and ≈2 dpa. The loading condition in all irradiated constant-load 

CGR tests was validated by the size criterion specified in American Society for Testing and Materials 

(ASTM) standard E399 using irradiated yield stress. Based on cyclic and constant-load CGR results, 

IASCC susceptibility was evaluated for the tested materials as a function of neutron dose, processing 

conditions, and water chemistry. Fracture toughness tests were also carried out in the BWR environment 

after the CGR tests. These latest results were combined with previous results from Halden-I and -II 

irradiations to analyze the effects of neutron dose, water chemistry, alloy compositions, and welding and 

processing conditions on IASCC. The effect of neutron irradiation on the fracture toughness of austenitic 

SSs was also evaluated at dose levels relevant to BWR internals. The main findings for this analysis 

follow.

1. Effect of welding process

The post-irradiation tests demonstrated that the cyclic CGRs increase significantly with increases in 

neutron irradiation dose for all HAZ specimens regardless of weld type or post-weld processing. The 

constant-load CGRs for as-welded SA and SMA HAZ specimens are approximately one order of 

magnitude greater than the disposition curve in NUREG-0313 for sensitized SSs in water with 8 ppm 

dissolved oxygen (DO). No significant difference was present between as-welded and thermally treated 

HAZ materials with SA and SMA welds. The effect of low-temperature sensitization also appears to be 

insignificant for the IASCC susceptibility of SA and SMA weld HAZ materials. 

2. Effect of sensitization

Type 304 SS specimens were thermally sensitized at 600°C for 10.5 h and 24 h to investigate the 

effect of degree of sensitization on IASCC. The specimens sensitized at the two conditions exhibited

similar cracking behavior under both cyclic and constant-load test conditions. Therefore, it appears that
xviii

the degree of sensitization does not play a significant role in the cracking susceptibility of irradiated

materials. Irradiation effects appear to be the dominating factor for IASCC behavior.

3. Effect of material type and composition

The influence of material type and material chemistry on IASCC behavior was studied by analysis 

of earlier CGR and slow strain rate test (SSRT) results. For the CGR tests, cyclic and constant-load 

CGRs of three austenitic SSs were compared at three dose levels in normal water chemistry (NWC) and 

hydrogen water chemistry (HWC). Within the scatter of the data, IASCC susceptibility appears to be 

insensitive to the material type in NWC. In HWC for dose levels up to 3 dpa, any effect of materials type 

is unclear due to the low CGRs observed in the low potential environment for all material types. For 

SSRT tests, the fraction of intergranular cracking was used to characterize the IASCC susceptibility of 

different alloys. The SSRT results indicate that minor alloying elements such as carbon, sulfur, and 

oxygen may play an important role in the IASCC process. Low-carbon and high-sulfur SSs were found 

to be prone to IASCC. Sulfur content below 0.002 wt.% was found to be critical for IASCC resistance in 

austenitic SSs in high-DO BWR water. It is also reported in the literature that high silicon content has a 

detrimental effect on the SCC behavior of austenitic SSs. However, data are limited and it is unclear at 

present if silicon segregation resulting from neutron irradiation does elevate the cracking susceptibility. 

4. Effect of water chemistry

Stress corrosion cracking is a strong function of corrosion potential. It has been repeatedly 

demonstrated that a low corrosion potential can effectively mitigate IASCC. Both cyclic and constantload CGR results have shown at least one order of magnitude reduction in CGR for HWC compared with 

NWC. However, a few tests conducted within the high-K (stress intensity factor) range showed that the 

beneficial effect of HWC is absent in some high-dose specimens (>3 dpa). This anomalous cracking 

behavior in HWC has often been disregarded in the literature because the loading conditions typically do 

not meet proposed size criteria for irradiated materials. However, the insensitiveness to low corrosion 

potential demonstrated by some high-dose specimens should not be ignored solely based on the violation 

of the proposed size criteria whose technical basis has not been adequately justified. Tests conducted 

under K-constraint conditions that do meet standard validity criteria have demonstrated that the HWC 

beneficial effect is lost with Type 304L SS at doses as low as 3 dpa. Since the dose dependence of the 

HWC effect is poorly understood, additional data are needed for doses higher than 3 dpa.

5. Effect of neutron dose

Halden-I crack growth results in NWC at three doses were used to illustrate the dose dependence of 

IASCC behavior. Environmental enhancement observed in the cyclic CGR tests increased continuously

with dose. The constant-load CGRs, however, rose sharply above 0.45 dpa and saturated in the high-dose 

region. The constant-load CGRs for the three SSs examined were approximately a factor of six higher 

than the NUREG-0313 curve between 1.35 and 3 dpa. The fracture toughness of austenitic SSs also

decreased rapidly with neutron dose. The threshold dose for a sharp decrease in fracture toughness was 

about 0.3 dpa. Above 5 dpa, a slower reduction in fracture toughness with increasing dose is expected for 

austenitic SSs. 

6. Effect of grain boundary engineering

Crack growth tests were also performed on GBE and non-GBE Type 304 and 316 SSs in NWC. The

GBE treatments increased the fraction of coincident-site-lattice (CSL) boundaries to approximately 60% 

in these materials. However, the beneficial effect on IASCC resistance was not readily revealed. Further

examination of the unirradiated materials showed large grain size and precipitation microstructure for 

both GBE Type 304 and 316 SSs. The ineffectiveness of GBE treatment suggests that the beneficial 
xix

effect provided by a high fraction of CSL boundaries may be overwhelmed by detrimental effects 

introduced by brittle precipitates or large grain size.
xx

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xxi

Acknowledgments

The authors would like to thank T. M. Karlsen, OECD Halden Reactor Project, Halden, Norway, 

for specimen irradiations in the Halden reactor; D. O. Pushis for specimen retrieval; and L. A. Knoblich, 

E. E. Gruber, E. J. Listwan, W. K. Soppet, D. L. Rink, and R. Clark, for their contributions to the

experimental effort. The authors would also like to thank Dr. Yong Yang and Prof. Todd Allen for their 

contributions to microstructural examination. The authors are also grateful to Dr. R. Tregoning for 

reviewing the manuscript and for many helpful comments. Further more, the authors are thankful to Drs. 

B. Alexandreanu, K. Natesan, W. H. Cullen, Jr., S. Crane, and A. S. Rao for many discussions. This 

work is sponsored by the Office of Nuclear Regulatory Research, U.S. Nuclear Regulatory Commission, 

under Job Code N-6519; Program Manager: Appajosula S. Rao.
xxii

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xxiii

Acronyms and Abbreviations

ASTM American Society for Testing and Materials 

BWR Boiling Water Reactor

CGR Crack Growth Rate

CSL Coincident Site Lattice

CT Compact Tension

DCPD Direct Current Potential Drop 

DO Dissolved Oxygen

dpa Displacements per Atom

EBSD Electron Backscattered Diffraction 

ECP Electrochemical Potential

GBE Grain Boundary Engineered

GG Grand Gulf 

HAZ Heat Affected Zone

HWC Hydrogen Water Chemistry

IASCC Irradiation-Assisted Stress Corrosion Cracking

IG Intergranular

IML Irradiated Materials Laboratory

LWR Light Water Reactor

NWC Normal Water Chemistry

PWR Pressurized Water Reactor

RIS Radiation-Induced Segregation

SA Submerged Arc

SCC Stress Corrosion Cracking

SHE Standard Hydrogen Electrode

SMA Shielded Metal Arc

SS Stainless Steel

SSRT Slow Strain Rate Tensile

TG Transgranular

TEM Transmission Electron Microscopy
xxiv

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1

1 Introduction

Stress corrosion cracking (SCC) is a major degradation mechanism for austenitic stainless steels 

(SSs) in high-temperature aqueous environments. Premature failures resulting from SCC are commonly

found in thermally sensitized austenitic SSs under corrosive environments.

1,2 In light water reactors

(LWRs), reactor core internal components are not only exposed to high-temperature coolant but also

subjected to high-energy irradiation by fission neutrons. The exposure to fast neutron irradiation can 

further increase the susceptibility of SSs to SCC by elevating the corrosion potential of the reactor coolant 

and by introducing new embrittlement mechanisms through radiation damage. Various nonsensitized 

austenitic SSs and nickel alloys have been found to be prone to intergranular (IG) cracking after extended 

neutron exposure. Service failures have been seen in the core shroud, jet pump assembly, top guide, and 

core plate in boiling water reactors (BWRs).

3-9 Failures of baffle former bolts have also been observed in 

pressurized water reactors (PWRs).10 The elevated susceptibility to SCC in irradiated materials,

commonly referred to as irradiation-assisted stress corrosion cracking (IASCC), is a complex 

phenomenon that involves simultaneous actions of irradiation, stress, and corrosion. In recent years, as 

nuclear power plants have aged and irradiation dose has increased, IASCC has become an increasingly

important issue. A comprehensive understanding of IASCC is important for determining the structural

integrity of reactor internal components and the safety and economy of aging nuclear power plants. 

Neutron irradiation affects material microstructure, microchemistry, and local water chemistry 

significantly. Point defects resulting from ballistic collisions between energetic particles and lattice atoms

are the basis of radiation damage. At representative LWR operating temperatures, faulted dislocation 

loops, network dislocations, and, on occasion, voids can form in irradiated austenitic SSs.

11-13 These 

microstructural changes are also accompanied by local microchemistry change during irradiation. 

Through the inverse Kirkendal effect, solute redistribution can occur around point defect sinks.

14 

Radiation-induced segregation (RIS) of alloy elements can, therefore, be seen in irradiated SSs at LWR 

coolant temperatures.

15 Neutron irradiation has a strong effect on water chemistry by radiolysis as well. 

Water decomposes into oxidant species under ionizing irradiation and thus elevates corrosion potential.

16 

Because of the unstable nature of radiolytic products, the radiolysis effect on water chemistry is short

lived and depends on neutron flux rather than fluence. 

All irradiation effects on microstructure, microchemistry, and water chemistry contribute to 

IASCC.

16,17 However, the occurrence of IG cracking in post-irradiation tests indicates that “persistent”

irradiation effects on microstructure and microchemistry are the root cause of the elevated cracking

susceptibility.18 For this reason, IASCC susceptibility depends on the accumulated irradiation dose, i.e.,

neutron fluence. A characteristic rise in IASCC susceptibility is indeed observed in post-irradiation tests 

above a “threshold” neutron fluence, as shown in Fig. 1 for commercial and high-purity heats.

19,20 

Because of the complexity of the IASCC dependence on material and environmental variables, the 

threshold dose for IASCC varies with different test conditions. The proposed threshold is 

≈5 x 1020 n/cm2 (E >1 MeV)*

 

*All references to fluence levels are calculated for fast neutron E ≥1 MeV.

(≈0.75 dpa) for BWR environments21,22 and approximately one order of 

magnitude higher for PWRs.23 Some studies have shown that IASCC could occur at a dose as low as 

≈2 x 1020 n/cm2 (≈0.3 dpa) for high-purity heats tested in water with high dissolved oxygen (DO)

concentrations.

24 A better understanding of the onset of IASCC is of great technical importance. 

Meanwhile, the dose dependence of IASCC susceptibility and fracture behavior is also a central topic for 

issues related to aging reactor internal components. 
2

0

20

40

60

80

100

1020 1021 1022

0.1 1 10

Type 304 SS

Type 316 SS

HP 304 SS

%IGSCC

Neutron Fluence (n/cm2

, E> 1 MeV)

Dose (dpa)

Austenitic Stainless Steels

High-Purity Water ~8ppm DO

Commercial Heats

High-Purity Heats

Figure 1. Susceptibility of irradiated 

austenitic SSs to intergranular SCC as a 

function of fluence in high-DO water. Data 

from slow-strain-rate tensile tests.

21, 24-26

The IASCC susceptibility is a complex function of environmental, material, and testing variables. 

Because SCC susceptibility depends strongly on the corrosion potential between –100 and 0 mV with 

respect to a standard hydrogen electrode (SHE),27 a low potential in BWR hydrogen water chemistry

(HWC) or PWR primary water chemistry is considered to be beneficial. The threshold fluence for 

IASCC is, therefore, higher under low potential conditions. An effective way to mitigate SCC is to 

control the electrochemical corrosion potential (ECP) in BWRs.28,29 However, low corrosion potential 

does not provide complete immunity to IASCC. Intergranular SCC has also been observed in coldworked, irradiated SS baffle bolts in PWRs.

30,31 In the present study, the cracking and fracture behavior 

of austenitic SSs under different water chemistries have been evaluated for different dose levels. 

Early mechanistic studies of IASCC focused mainly on the depletion of Cr at grain boundaries and 

its influence on the grain boundary oxidation.32–34 Radiation-induced segregation was considered as the 

main reason for the sharp Cr depletion at grain boundaries. It is believed that the absence of protective 

oxide layers stimulates cracking along grain boundaries in the IASCC process. More recently, the roles 

of impurity elements such as Si, P, and S on IASCC susceptibility have also been investigated.16,27 

Although redistribution of impurities to grain boundaries due to RIS clearly does occur, no obvious

correlations have yet been found between the original concentration of impurities in the bulk material and 

IASCC susceptibility. Recent studies on IASCC susceptibility have also attributed a greater role to 

radiation hardening and the development of localized deformation modes.35 The localized plastic 

deformation induced by irradiation contributes to the rupture of the oxide film near the grain boundary 

and, therefore, is a critical factor for IASCC susceptibility.36 Based on the progress in the mechanistic 

understanding of IASCC, various engineering methods have been proposed to mitigate the effect of 

IASCC on LWR core internal components. In addition, because of the predominately IG nature of 

IASCC, considerable attention has been paid to the role of grain boundary properties. It is believed that 
3

coincident site lattice (CSL) grain boundaries with low Σ numbers*

This report summarizes the test results for the last group of Halden-II compact tension (CT) 

specimens. The IASCC susceptibility of austenitic SSs and weld heat-affected-zone (HAZ) materials was

evaluated by crack growth rate (CGR) tests in simulated BWR environments. Post-irradiation CGR and 

J-R fracture toughness tests were performed to provide data and technical support for helping the NRC 

address issues related to aging degradation of reactor core internal structures and components. The 

fracture and cracking behavior of submerged-arc (SA) and shielded-metal-arc (SMA) HAZ materials, 

thermally sensitized SSs, and grain-boundary-treated SSs were investigated at several dose levels. These 

latest results were combined with previous results from Halden-I and II irradiations to analyze the effects 

of neutron dose, water chemistry, alloy compositions, and welding and processing conditions on IASCC. 

The effect of neutron irradiation on the fracture toughness of austenitic SSs was also evaluated at dose 

levels relevant to BWR internals. 

(≤29) are more resistant to IG 

cracking.37-41 Grain boundary engineering (GBE), a thermomechanical treatment that systematically 

increases the population of low Σ CSL grain boundaries, is considered as another potential approach to 

improve a material’s resistance to IASCC.42-45 

 

* Σ is the reciprocal of the density of coincident sites. For example, Σ of 17 means that one in every 17 sites in one grain 

coincides with a site from the neighboring grain, and the density of coincident sites is 1/17.
4

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5

2 Experimental

2.1 Compact Tension Specimens and Materials 

Crack growth rate and J-R fracture toughness tests were conducted on 1/4T-CT specimens. Figure 

2 shows a schematic of a 1/4T-CT specimen used in the present study. Besides two loading pin holes on 

a normal CT specimen, four additional holes were added to the specimen design to install electrical leads

for crack length measurement using the potential drop method. The two through holes in the back of the 

specimen are designed for current leads, while the two inclined holes across the machine notch are for

potential leads. Side grooves were machined for each specimen to ensure in-plane crack propagation 

during testing. 

Figure 2. Configuration of compact-tension specimen used in this study (dimensions in mm).

Materials sectioned from the HAZ of an SA weld and SMA weld, sensitized 304 SSs, a cast SS, 

and GBE-treated Type 304 and 316 SSs are included in this study. Table 1 provides the chemical 

composition of these alloys.

Table 1. Composition for the Halden-II irradiation materials

Steel Type Material ID Descript Ni Si P S Mn C N Cr Mo O

304 10285 HAZ, Sensitized 8.45 0.60 0.015 0.007 1.90 0.070 0.084 18.56 0.51 0.013

304L GG Top Shell HAZ 9.05 0.53 0.027 0.016 1.84 0.013 0.064 18.23 0.44 0.010

304L GG Bottom Shell HAZ 8.95 0.55 0.023 0.008 1.80 0.015 0.067 18.62 0.31 0.014

304 Non-GBE 304 Base a 8.46 0.40 0.013 0.014 1.56 0.065 0.086 18.32 0.36 0.010

304 GBE 304 GBE a 8.42 0.46 0.014 0.002 1.54 0.065 0.088 18.38 0.50 0.006

316 Non-GBE 316 Base a 10.30 0.43 0.013 0.014 1.53 0.055 0.054 16.42 2.19 0.006

316 GBE 316 GBE a 11.12 0.56 0.010 0.022 1.80 0.070 0.056 16.57 2.28 0.007

CF-8M 75 Cast SS 9.12 0.67 0.022 0.012 0.53 0.065 0.052 20.86 2.58 -

a

The HAZ specimens used in this study were taken from a core-shroud weld of the cancelled Grand 

Gulf (GG) reactor and a laboratory-prepared weld. The GG weld was obtained from the H5 weld of the 

core shroud whose top and bottom shroud shells were fabricated from ASME SA240 Type 304L hotAverage values of two specimens. 
6

rolled plates. The GG weld is a double-V butt joint SA weld with ER308L filler metal. Details of the GG 

weld have been given in previous NUREG reports.19,20 The laboratory-prepared weld is a SMA weld 

manufactured from 30-mm-thick plates of Type 304 SS (Heat 10285). The weld has a single-V butt joint 

design and was produced by 31 weld passes using E308 filler rods. Fabrication details of this laboratory 

weld can also be found in these NUREG reports.19,20 

The GG SA weld and the laboratory-prepared SMA weld were sliced into 9.5-mm thick blanks. 

Some blanks were thermally treated for 24 h at 500°C to simulate low-temperature sensitization. The CT 

specimens were cut from both as-welded and thermally treated weld blanks. All slices of the weldment 

were polished and etched to reveal the weld microstructure, as shown in Fig. 3. The machine notch of a 

CT specimen was located about 1 mm away from the fusion line on the HAZ side of a weldment. The 

crack plane is parallel to the fusion line and the expected direction of crack propagation is toward the 

center of the weld. Side grooves were machined for each CT specimen to ensure crack advance within 

the weld HAZ. 

(a) (b)

Figure 3. Weld microstructure in the (a) Grand Gulf SA weld and (b) laboratory-prepared SMA weld. 

Sensitized specimens were obtained from the same Type 304 SS plate (Heat 10285) that was used 

for making the laboratory-prepared weld. The as-received mill-annealed Type 304 SS was sensitized at 

600°C for 10.5 h and 24 h to provide two different degrees of sensitization. The principal loading 

direction in all the sensitized CT specimens is perpendicular to the rolling direction, and the direction of 

crack advance is along the plate thickness (the T-S orientation). 

Table 2. Fractions of CSL boundaries in Types 304 and 316 SSs with and without GBE processing.

Boundary Type (Σ)

Heat 1 3 5 7 9 11 13 15 17 19 21 23 25 27 29 ≤29

non-GBE 304 2.1 28.9 1.0 0.7 5.4 0.3 0.5 0.8 0.4 0.3 0.4 0.2 0.4 2.4 0.2 44.1

GBE 304 2.2 52.1 0.4 0.1 7.3 0.2 0.2 0.1 0.1 0.2 0.2 0.4 0.2 3.2 0.3 67.3

non-GBE316 4.9 37.8 0.4 0.4 2.6 0.4 0.4 0.3 0.3 0.5 0.5 0.1 0.3 1.4 0.3 50.7

GBE 316 3.2 45.5 0.2 0.4 4.8 0.7 0.4 0.2 0.3 0.3 0.3 0.2 0.1 1.5 0.2 58.4

Type 304 and 316 SSs with and without GBE treatment were also included in the present study. 

The GBE processing is a thermomechanical treatment that involves one or several iterations of strainanneal or recrystallization-anneal schedules to increase the population of low-Σ number CSL boundaries. 

The chemical compositions of the GBE and non-GBE materials are given in Table 1. The fractions of 

low-Σ CSL boundaries (≤29) for non-GBE Type 304 and 316 SSs are 44% and 51%, respectively,

according to the Brandon criterion. The GBE processing increased the fraction of low-Σ boundaries to 
7

67% in GBE Type 304 SS and to 58% in GBE Type 316 SS. The fractions of CSL boundaries in these 

materials are given in Table 2. 

The CF-8M SS (Heat 75 in Table 1) was obtained from a static cast plate approximately 610 x 610 

x 76 mm. The cast SS has a duplex ferrite-austenite structure consisting of lacy ferrite morphology. 

Ferrite phase forms elongated islands among austenite matrix as a result of casting.

46 The total ferrite 

content in Heat 75 is about 28%. Prior to irradiation, the cast SS was aged for 10,000 h at 400°C in air. 

2.2 Irradiation

All specimens were irradiated in helium-filled irradiation capsules in the Halden reactor, a natural 

circulation boiling heavy water reactor. The maximum thermal power of the Halden reactor is 25 MW,

and its normal operation conditions are approximately 235°C and 2.8 MPa (~406 psig). The specimen

temperature was monitored by two sets of melting alloy temperature monitors installed in each irradiation 

capsule. The actual neutron doses received by the specimens were determined after irradiation by activity 

analysis on fluence monitor wires mounted at different positions in the irradiation capsules. Iron and 

nickel wires were used to record the fast neutron fluence (E≥1MeV), while activity analysis of Al/1%Co 

wires was used to estimate thermal neutron fluence (E<0.625 eV). The fast neutron fluence was used to 

estimate the displacement damage (dpa) on the irradiated specimens. Table 3 summarizes the irradiation 

conditions for the Halden-II CT specimens. 

Table 3. Irradiation conditions for the Halden-II CT specimens.

Material Heat Treatment

Number 

of

Specimens

Fast Neutron Fluence

E >1 MeV 

(x1021 n/cm2)

Displacement 

Damage

(dpa)

Irradiation 

Temperature

(°C)

GG 304L SA HAZ As-welded 2 0.50 0.75 ~2902 1.44 2.15 290-296

24 h @ 500°C 2 1.63 2.43 ~290

304 SS SMA HAZ

As-welded 2 0.50 0.75 ~290

2 1.44 2.15 290-296

24 h @ 500°C 2 0.50 0.75 ~290

2 1.44 2.15 290-296

CF-8M Cast SS 10,000 h @ 400°C 2 1.63 2.43 ~290

Sensitized 304 SS

10.5 h @ 600°C 2 0.5 0.75 ~290

2 1.44 2.15 290-296

24 h @ 600°C 2 0.5 0.75 ~290

2 1.44 2.15 290-296

304 SS non-GBE 1 1.31 1.96 290-296

GBE 1 1.31 1.96 290-296

316 SS non-GBE 1 1.31 1.96 290-296

GBE 1 1.31 1.96 290-296

2.3 Test Facility

Irradiated specimens were tested in a mechanical test facility located in the Irradiated Materials 

Laboratory (IML) at Argonne National Laboratory. The IML is a radiological-controlled laboratory

maintained at a negative pressure with respect to the surroundings and has four shielded hot cells 

designed for handling highly radioactive materials. The mechanical test facility consists of a loading 

frame, an Instron 8500+ Dynamic Materials Testing System, and a water recirculation system. The load 

train, actuator, furnace, autoclave, load cell, and linear variable displacement transducer of the test system 

are located inside a hot cell. The control console, data acquisition system, hydraulic pump, and water 

recirculation loop are kept outside the hot cell for the convenience of operation. Figure 4 is a schematic 
8

of the loading frame used in the present study. A 22-kN load cell is installed on the top of the pull rod. A 

1-liter SS autoclave is installed inside the furnace to provide the simulated BWR coolant environment. 

The furnace is mounted on a pneumatic cylinder and can be raised to enclose the autoclave with the load 

cage during tests. 

The hydraulic actuator is located on top of the load frame, with the test train components suspended 

beneath it. The load cage consists of the cover plate and a 12.7-mm-thick bottom plate separated by four

compression rods. The lower two-piece clevis assembly is fastened to the bottom plate of the cage with 

the two sections connected by an Zircaloy pin oxidized at 500°C for 24 h. The gap between the two 

pieces of clevis is electrically insulated by use of mica washers. The top clevis assembly is connected to 

the pull rod and is also insulated with the same types of Zircaloy pin and mica washers. The CT 

specimen is mounted in the clevises with 17-4 PH SS pins. Platinum wires are used for DC potential drop 

measurements during CGR and J-R tests. Two platinum wires attached to SS split pins are inserted into 

the through holes on the 1/4T-CT specimens (Fig. 2) to be used as current leads. Potential leads are two 

platinum wires attached to the SS pins that are threaded into the inclined holes at the machine notch. 

Hydraulic

Actuator

Furnace

Load Cell

Pneumatic

Cylind er

Figure 4. Mechanical test system for conducting CGR and J-R tests on radioactive specimens.

The water recirculation system that provides simulated BWR water is located outside the hot cell. 

The recirculation system consists of a 135-liter water retention tank, a high pressure pump, two 

regenerative heat exchangers, an autoclave and preheater, an ECP cell and preheater, a Mity Mite backpressure regulator, an ion-exchange cartridge, two 0.2-micron filters, a demineralizer resin bed, a pressure 

transducer, pH and conductivity sensors, and numerous pressure relief valves. A schematic diagram of 

the water recirculation system is given in Fig. 5. 
9

The simulated BWR environment in this study consisted of high-purity deionized water with either

250-700 ppb DO, corresponding to normal water chemistry (NWC) BWR water or <30 ppb DO, 

corresponding to HWC BWR water. The ECPs for SS are in the range of 160 to 240 mV (vs. SHE) in 

NWC and -200 to -500 mV (vs. SHE) in HWC. The deionized water is prepared by passing purified 

water through a set of filters that comprise a carbon filter, an Organex-Q filter, two ion exchangers, and a 

0.2-mm (8-mil) capsule filter. The DO level in water is established by maintaining a cover gas of 

nitrogen plus 1% oxygen above the supply tank and initially bubbling the gas mixture through the 

deionized water. The ECP of a Pt electrode and an SS rod (1/8” in diameter) located at the exit of the 

autoclave was monitored continuously during the test, and water samples were taken periodically to 

measure pH, resistivity, and DO concentration. The DO level was measured in the in-cell facility by the 

colorimetric technique using CHEMets sampling ampoules. 

V7

V25

V26

V23

V22

V24

V19

V20 V21

V27

V8 V9

V13

28 V11

V14

29

30

26

25

23

18

24

V12

27

15

16

V15

32

13 14

23

6

11

33

20

17

19

OUTSIDE HOT CELL INSIDE HOT CELL

21 21

21

22 21

31

35

11

V16

13

V17

1

2 3

4

8

7

6

5

9

10

11

12 36

V3

V2

V1

V4

V6

V5

13

33

31

15

34

V10

V18

V28

V29

V30

21

1. Cover gas supply cylinder

2. Two stage high-pressure regulator

3. Low-pressure regulator

4. Flow meter

5. Vacuum & pressure gauge

6. Pressure relief valve

7. Vent to air

8. Feed water storage tank

9. Bubbling tube

10. Feed water fill port

11. Water sample port

12. Solenoid valve

13. 0.2-micro filter

14. High-pressure pump

15. Check valve

16. Rupture disk

17. Accumulator

18. Heat exchanger

19. Heat exchanger drain port

20. Heat changer bleed port

21. Thermocouples

22. Pressure transducer

23. System bleed port

24. High-pressure gauge

25. Autoclave preheater

26. Autoclave

27. Thermocouple well

28. ECP cell preheater

29. ECP cell

30. ECP cell bypass line

31. Water-cooled heat exchanger

32. Back-pressure regulator

33. Ion-exchange cartridge

34. Conductivity sensor

35. pH sensor

36. Recirculation pump

Figure 5. A schematic diagram of the water recirculation system.

All tests in the simulated BWR environment were started in high-purity water that contained 250-

500 ppb DO (NWC). A gas mixture of 1% O2 balanced with N2 was normally used to achieve the 

required DO level for NWC in the feed water. In some tests, after the initial CGR test in NWC, the DO 

level of the feed water was decreased to <30 ppb by switching to a N2+4% H2 gas mixture for cover gas. 

Because of the low water flow rates, it can take several days for the environmental conditions to be 

stabilized for the in-cell tests. In general, the changes in ECP were slower in the SS sample than in the Pt

electrode. 
10

2.4 Test Procedures 

2.4.1 Crack Growth Rate Test

The CGR tests were performed in accordance with the American Society for Testing and Materials 

(ASTM) E-647, “Standard Test Method for Measurement of Fatigue Crack Growth Rates,” and ASTM E1681, “Standard Test Method for Determining a Threshold Stress Intensity Factor for EnvironmentAssisted Cracking of Metallic Materials under Constant Load.” All CGR tests in the present study were 

performed on 1/4T-CT specimens whose dimensions are given in Fig. 2. With the DC potential drop 

(DCPD) method, crack extension was monitored throughout a CGR test while the sample was under

cyclic or constant loading. 

A CGR test began with a fatigue loading stage in the simulated BWR environment with load ratio 

(R) around 0.2-0.3, frequency of 1-5 Hz, and maximum stress intensity factor (Kmax) between 10 and 

16 MPa m1/2. The objective of this test period is to generate a sharp fatigue crack at the machine notch

and to extend the crack front beyond the damaged region immediately next to the machine notch. After

an initial crack advance of several hundred micrometers, a prescribed loading sequence was followed to 

introduce environmentally enhanced cracking. To facilitate the transition of a transgranular (TG) fatigue 

crack to an IG stress corrosion crack, a series of test periods was implemented with a slow/fast sawtooth 

waveform at a constant Kmax. During these test periods, R was increased incrementally to 0.7 with 

increasing rise times up to 1000 s. Once the environmentally enhanced cracking was established, the test 

condition was changed to a constant load with or without periodic unloading to R = 0.7 in every 1 or 2 h. 

Crack growth rates (m/s) were measured and reported for each test stage. With the increase of the crack 

length, the load was gradually reduced to maintain a constant K throughout each stage during the constant 

load test. 

The stress intensity factor K for a CT specimen was calculated as follows: 

( )

1/2 3/2

N

a 2

P a W K f

BB W a W

1

W

    +     = ⋅⋅         −  

(1)

where P is applied load; B is the specimen thickness; BN is the net specimen thickness (or distance 

between the roots of the side grooves); a is crack length; and W is specimen width (measured from the 

load line to the back edge of the specimen). The geometry factor for a CT specimen is:

f a

W











 = 0.886 + 4.64 a

W











 −13.32 a

W













2

+14.72 a

W













3

− 5.60 a

W













4

(2)

During a test in water, the friction between the autoclave seal and the pull rod exerted a force of

about 20 to 30 lb (89 – 133 N) on the pull rod. This fraction load was measured for each test to correct 

the real load acting on a specimen. The applied K and the load ratio were determined by subtracting the 

frictional load from the measured maximum load and adding frictional load to the measured minimum 

load. The most significant effect of this correction is on the waveform for the cyclic tests. For triangular 

or sawtooth waveforms generated by the test machine, the actual loading waveforms applied on a 

specimen are trapezoidal because the top and bottom parts of the loading or unloading cycles are 

truncated to overcome the friction at the sealing surface. For example, for a test intended to be conducted 
11

at R = 0.7 and a sawtooth waveform with 300-s rise time and 12-s return time, the actual loading 

waveform was trapezoidal with 112-s hold time at minimum load, 188-s rise time, 6-s hold time at 

maximum load, and 6-s return time. 

After each CGR test, optical images were taken of the fracture surface, and the final crack length 

was measured to compare with the DCPD result. Because of unbroken ligaments or an uneven crack 

front, crack extension is usually underestimated by the DCPD method. Thus, the crack length obtained 

by the DCPD method for each test period was scaled proportionally with the optical measurement of the 

final crack length. The corrected plot for crack length vs. time was used to determine the crack growth 

rate for each test period. For cyclic loading, only the rise time was used to determine the crack growth 

rate. 

In linear elastic fracture mechanics, the stress intensity factor is a single parameter that

characterizes the stress-strain field at a crack tip. For K-controlled fracture, the plastic zone ahead of the 

crack tip must be embedded within an elastic singularity zone. To ensure a K-controlled fracture so that 

the test results are independent of sample size, the following sample/crack size requirement is adopted,

Beff and (W - a) ≥2.5 (K/σy)2 (3)

where K is the applied stress intensity factor, σy is the yield stress of the material, a is crack length, W is 

specimen width, and the Beff is the specimen effective thickness, defined as (B·BN)0.5. Joyce and 

Tregoning47 have demonstrated that this size criterion is a very conservative requirement to ensure a 

small-scale yielding condition. For materials with high strain-hardening capacity, i.e., (σu/σy) ≥1.3 (e.g.,

nonirradiated SSs), the flow stress defined as σf = (σu + σy)/2 rather than the yield stress can be used to 

scale the plastic zone size. Unless otherwise noted, all CGRs reported in this study were obtained under 

conditions where Eq. (3) is satisfied.

2.4.2 Model for Crack Growth Rate Test Analysis

To reveal the environmental effects on cracking, a superposition model 48 was used to analyze the 

cyclic CGR test results. The model assumes that a cyclic CGR in environment ( env a ) includes three 

components representing mechanical fatigue in air ( air a ), corrosion fatigue in the environment ( cf a ), and 

stress corrosion cracking under constant load ( scc a ), i.e.

env air cf scc a a aa    = ++ (4)

The mechanical fatigue growth rate, air a (m/s), can be determined from a correlation developed by 

James and Jones: 49

3.3

air SS r a C (T) S(R) K / t  = ⋅ ⋅∆ (5)

where ∆K = Kmax - Kmin in MPa m1/2, and tr is the rise time (s) of the loading waveform. Furthermore, 

CSS is a function of temperature only and is given by a third-order polynomial function as 

-12 -15 -17 2 -20 3 C = 1.9142 10 + 6.7911 10 T - 1.6638 10 T + SS ×× × × 3.9616 10 T (6)

The S(R) is a piecewise function of the load ratio (R):
12

S(R) = 1.0 R < 0

S(R) = 1.0 + 1.8R 0 < R <0.79

S(R) = -43.35 + 57.97R 0.79 < R <1.0, (7)

where the load ratio R is defined as Kmin/Kmax. To evaluate environmental enhancement, the mechanical 

fatigue CGR in air was estimated for each cyclic loading condition, and was compared with the measured 

CGR in the environment. Under the rapid cyclic loading typically used for precracking, the crack growth 

is dominated by mechanical fatigue. The CGR in environment is very similar to that observed in air. 

Environmentally enhanced cracking is more evident under high load ratios and long rise times. Usually, 

significantly higher cyclic CGRs in high-purity water can be seen for irradiated SSs under load ratios 

greater than 0.5 and rise time longer than 30 s. 

A model developed by Shack and Kassner48 has been used successfully to characterize the 

environmental effects on fatigue crack growth for nonirradiated austenitic SSs. In the absence of a 

significant contribution of SCC to growth rate, the CGRs in water with ≈0.3 ppm DO are best represented 

by:

5 0.5

env air air a a 4.5 10 (a ) −   = +× (8)

and the CGRs in water with ≈8 ppm DO is,

4 0.5

env air air a a 1.5 10 (a ) −   = +× (9)

Under SCC conditions ( scc a ), the CGR can be represented by a correlation given in NUREG-0313, 

Rev. 2: 50

2.161 scc a AK  = ⋅ (10)

The magnitude of the constant A depends on the water chemistry and composition and structure of the 

steel. An A value of 2.1 x 10-13 is proposed in NUREG-0313 for sensitized SS in water with 8 ppm DO. 

The value of constant A is smaller in low-DO environments, such as HWC BWR or PWR environments. 

For water with 0.2 ppm DO, the constant A is taken as one-third that given in NUREG-0313. In the 

present study, CGRs measured under constant-loading conditions are plotted as a function of stress 

intensity, with the NUREG-0313 curve used as a reference. With the K-dependence given in Eq. (10) as 

a guideline, the SCC susceptibility was evaluated for different irradiated SSs and for different water 

chemistries (NWC vs. HWC). 

2.4.3 Fracture Toughness J-R Curve Tests

For some specimens, a fracture toughness J-R test was performed at 289°C in NWC after the CGR 

test. The J-R test was conducted at a constant extension rate of ≈0.43 μm/s (0.017 mil/s) in accordance 

with ASTM E-1820 for “Standard Test Method for Fracture Toughness.” The test was interrupted 

periodically (by holding the specimen at constant extension) to measure the crack length. Prior to a 

DCPD measurement, the specimen was held at a constant strain for ≈30 s for stress relaxation. Since the 

specimen extension was monitored and controlled outside the high-temperature zone, the actual 

displacement of load points was determined by subtracting the extension of the load train from the 

measured extension. The load train displacement was determined as a function of applied load by using a 

specimen with high stiffness. 
13

The J-integral was calculated from the load (P) vs. load-line displacement (v) curves according to 

ASTM Specification E 1820. The J is the sum of the elastic and plastic components, Jel and Jpl,

J = Jel + J

pl (11)

At a hold point i corresponding to crack length ai, and vi and Pi on the load vs. load-line displacement 

curve, the elastic component Jel(i) is given by:

( ) ( ) 2 2

(i)

el(i)

K 1

J

E

− υ = (12)

where ν is Poisson’s ratio, and the stress intensity K(i) is calculated from Eqs. 1 and 2. The plastic 

component Jpl(i) is given by:

J

pl(i) = Jpl(i−1) +

η(i−1)

b

(i−1)

















A

pl(i)

− A

pl(i−1)

BN















 1− γ(i−1)

a

(i)

− a

(i−1)

b

(i−1)















 (13)

where b(i-1) is the remaining ligament (distance from the physical crack front to the back edge of the 

specimen) at a point i-1; Apl(i) is the area under the load vs. load-line displacement curve; and BN is the 

net specimen thickness. In addition, γ(i) and η(i)

η(i−1) = 2.0 + 0.552

b

(i−1)

W

are factors that account for the crack growth effects on J 

during the test and are expressed as:

(14)

γ(i−1) = 1.0 + 0.76

b

(i−1)

W (15)

In Eq. (13), the quantity Apl(i) - Apl(i-1) is the increment of plastic area under the load vs. load-line 

displacement curve between lines of constant plastic displacement at points i-1 and i. The quantity Jpl(i)

represents the total crack-growth-corrected plastic J at point i and is obtained by first incrementing the 

existing Jpl(i-1) and then modifying the total accumulated result to account for the crack growth 

increment. Accurate evaluations of Jpl(i) require small uniform increments in crack growth. The plastic 

area under the load vs. load-line displacement record is given by 

[ i i 1 pl(i) pl(i 1) ]

pl(i) pl(i 1)

PP v v

A A

2

− −

−

+ −     = + (16)

where the total and plastic components of the load-line displacement, v(i) and vpl(i), respectively, are 

expressed as

pl(i) (i) i LL(i) v v PC = − (17)
14

where CLL(i) is the compliance, (∆v/∆P)i, required to give the current crack length ai. For test methods 

that do not use the elastic compliance techniques, CLL(i) can be determined from knowledge of ai/W, as 

follows: 

2 3

ii i

LL(i) ' 2

e i

1.62 17.80(a / W) 4.88(a / W) 1.27(a / W)

C

E B [1 (a / W) ]

  +−+   = − (18)

where Be is specimen effective thickness given by B - (B - BN)2/B and E′ =E/(1 - υ2). 

After the test, the final crack size was marked by fatigue cycling in air at room temperature. The 

specimens were then fractured, and the fracture surface of both halves of the specimen was photographed 

with a telephoto lens through the hot cell window. The final crack length of each half of the fractured 

specimen was determined from the optical photograph by the 9/8 averaging technique. In this technique, 

nine measurements were taken across the width of the specimen at equal intervals, the two near-surface 

measurements were averaged, and the resultant value was averaged with the remaining seven 

measurements. The crack extensions determined from the DCPD method were proportionately scaled to 

match the final optically measured crack length.

The experimental results from the J-R curve test were analyzed in accordance with ASTM E-1820 to 

obtain the fracture toughness J-R curve. The DC potential data were corrected to account for the effects 

of plasticity on the measured potential, since large crack-tip plasticity can increase the measured 

potentials due to resistivity increases without crack extension. The change in potential before crack 

initiation was ignored, and the remainder of the potential change was used to establish the J-R curve. The 

normalized potential varies linearly with load-line displacement until the onset of crack extension. For all 

data prior to the loss in linearity, crack extension was expressed as ao + ∆aB, where ao is the initial crack 

length, and the crack extension ∆aB is calculated from the blunting line relationship ∆a = J/(4σf), where

σf is flow stress and defined as the average of yield strength (σy) and ultimate tensile strength (σu). The 

use of the blunting line given by ∆a = J/(4σf) is not consistent with ASTM E 1820, which specifies a 

slope of two times the effective yield stress (or flow stress) for the blunting line. However, for highstrain-hardened materials, such as austenitic SSs, a slope that is four times the flow stress (4σf) represents 

the blunting line better than the slope of 2σf defined in ASTM E 1820.

51,52 In irradiated materials,

irradiated yield strength (σy-irr) is used to replace σf. In this study, although strain hardening is low or 

even absent in irradiated SSs, the blunting line of 4σy-irr is still used to be more conservative in evaluating 

J values.
15

3 Experimental Results

Heat-affected zone materials from Type 304 and 304L SS welds, sensitized Type 304 SS, thermally 

aged cast SS, and Type 304 and 316 SSs with and without GBE treatment were included in the Halden-II 

test matrix. Two dose levels of the irradiated specimens, ~0.75 dpa and ~ 2 dpa, were available for 

testing. Table 4 summarized all specimens tested in the Halden-II study. Most J-R tests were performed 

in simulated BWR environments in NWC or HWC. Two J-R tests were performed in air. Test results on 

all nonirradiated and some irradiated specimens have been reported in previous NUREG reports. 19,20 

This report summarizes the results on specimens of weld HAZ (GG3TA-TT, 85-7B and 85-2ATT),

sensitized SS (85-1TT, 85-5A1TT, and 85-5A2TT), and GBE materials (IT304-01, GB304-01, IT316-01 

and GB316-01).

Table 4. Halden phase II specimens and type of tests performed. 

Material Type

Dose Specimen Test Performed Main 

Condition (dpa) ID CGR J-R Objectives

GG 304L SA HAZ As welded - GG5BA √ -

GG 304L SA HAZ As welded 0.75 GG5TA, GG5TB √ - Effect of SA weld

GG 304L SA HAZ As welded 2.15 GG6TA, GG6TB a √ √

GG 304L SA HAZ Heat treated, 24 h @ 500°C - GG3BA-TT √ - Effect of heat 

treatment on SA 

weld GG 304L SA HAZ Heat treated, 24 h @ 500°C 2.43 GG3TA-TT √ √

304 SMA HAZ As welded - 85-YA √ - Effect of SMA

Weld 304 SMA HAZ As welded 0.75 85-7A, 85-7B √ √

304 SMA HAZ As welded 2.15 85-XA, 85-XB √ √

304 SMA HAZ Heat treated, 24 h @ 500°C - 85-3ATT √ - Effect of heat 

treatment on SMA 

weld

304 SMA HAZ Heat treated, 24 h @ 500°C 0.75 85-1ATT √ -

304 SMA HAZ Heat treated, 24 h @ 500°C 2.15 85-2ATT √ √

CF-8M Cast SS Thermally aged, 10,000 h @ 400°C 2.43 75-11TT, 75-11TM √ √ Effect of thermal 

aging

304 SS Sensitized, 10.5 h @ 600°C 0.75 85-1TT √ √

Effect of 

sensitization

304 SS Sensitized, 10.5 h @ 600°C 2.15 85-3TT √ √

304 SS Sensitized, 24 h @ 600°C 0.75 85-5A1TT √ -

304 SS Sensitized, 24 h @ 600°C 2.15 85-5A2TT a - √

304 SS Non-GBE 1.96 IT304-01 √ √

Effect of GBE 

treatment

304 SS GBE treatment 1.96 GB304-01 √ √

316 SS Non-GBE 1.96 IT316-01 √ √

316 SS GBE treatment 1.96 GB316-01 √ √

a J-R test was performed in air.

3.1 Thermally Treated Type 304L SA Weld HAZ Irradiated to 2.4 dpa (GG3TA-TT)

To simulate the effect of low-temperature sensitization, a Type 304L SA weld was thermally

treated at 500°C for 24 h. The specimen, GG3TA-TT, was sectioned from weld HAZ with the machine 

notch parallel to the fusion line. The specimen was irradiated to ≈2.4 dpa at approximately 290°C. After 

being exposed to the test environment for over 80 h, fatigue precracking was started on this specimen in 

high-purity water with 600 ppb DO at 289°C. A triangular waveform was applied to the specimen at a 

frequency of 1 Hz, R ratio of 0.2 to 0.35, and Kmax ≈ 15 MPa m1/2. Following about 120-μm crack 

advance, the R ratio was increased, and a slow/fast sawtooth waveform was employed to transfer the

cracking mode from transgranular to intergranular. This initial attempt to reveal environmental 

enhancement was unsuccessful, and the crack growth rate became negligibly low. Several cyclic loading 

periods with triangular waveforms were repeated with an increased Kmax to re-establish a desired crack 

growth. The environmental and test conditions, crack length, and CGR for each test stage are 

summarized in Table 5. Figure 6 is a crack-length-vs.-time plot for this test. The green line is crack 
16

length, and the red line is the K value. The cyclic CGRs obtained in this weld HAZ specimen are shown 

in Fig. 7. 

Table 5. Test results for Specimen GG3TA-TT of Type 304L SS weld HAZ in BWR water at 289°C.

Test

Test 

Time,

Test

Temp.,

ECP,

a

mV (SHE)

O2

Conc.,a RLoad

Rise 

Time,

Return 

Time,

Hold 

Time,b Kmax, ∆K,

CGR in 

Env.,

CGR in 

Air,

Crack 

Length,

Period h °C Pt Steel ppb Ratio s s s MPa m1/2 MPa m1/2 m/s m/s mm

5.761

Pre a 7 290 290 235 600 0.21 0.42 0.42 0.08/0.08 15.4 12.2 1.12E-8 4.31E-18 5.866

Pre b 9 289 600 0.35 0.41 0.41 0.09/0.09 15.3 10.0 7.27E-9 2.68E-8 5.888

Pre c 22 289 600 0.49 19.5 19.5 5.5/5.5 15.0 7.6 negligible 2.71E-10 5.894

Pre d 28 289 600 0.49 0.39 0.39 0.11/0.11 15.3 7.8 3.62E-9 1.45E-8 5.920

Pre e 49 289 600 0.43 8.0 1.6 2.0/0.4 15.1 8.6 negligible 9.27E-10 5.921

Pre f 56 289 600 0.29 0.84 0.84 0.16/0.16 15.7 11.1 negligible 1.75E-8 5.933

Pre g 81 289 600 0.24 0.42 0.42 0.08/0.08 15.6 12.1 negligible 4.34E-8 5.930

Pre h 123 289 600 0.23 0.43 0.43 0.07/0.07 16.0 12.3 negligible 4.58E-8 5.941

Pre i 145 289 600 0.25 0.43 0.43 0.07/0.07 17.5 13.2 5.03E-9 5.77E-8 5.968

Pre j 152 289 600 0.25 0.85 0.85 0.15/0.15 17.1 12.9 4.18E-9 2.66E-8 6.019

Pre k 166 289 600 0.24 4.26 4.26 0.74/0.74 17.1 12.9 1.01E-9 5.41E-9 6.042

Pre l 170 289 600 0.24 0.85 0.85 0.15/0.15 17.1 13.0 1.65E-9 2.72E-8 6.051

1a 189 290 600 0.26 4.24 0.85 0.76/0.15 17.0 12.6 2.83E-10 5.09E-9 6.069

1b 195 290 600 0.31 2.11 0.84 0.39/0.16 17.7 12.3 1.39E-9 9.89E-9 6.084

1c 206 290 600 0.32 2.11 0.84 0.39/0.16 18.4 12.5 9.71E-10 2.68E-9 6.136

1d 222 290 600 0.31 42.3 10.2 7.7/1.8 18.3 12.6 2.63E-10 5.45E-10 6.148

1e 272 290 271 207 600 0.30 254 10.2 46/1.8 18.6 13.0 3.76E-10 9.97E-11 6.197

2 291 290 270 203 600 0.69 213 8.5 78/3.5 18.8 5.7 5.46E-10 1.17E-11 6.226

3 335 290 282 218 600 0.69 711 8.5 289/3.5 18.8 5.7 3.89E-10 3.51E-12 6.265

4 391 290 284 223 600 0.68 696 8.4 304/3.6 17.1 5.5 2.22E-10 2.98E-12 6.291

5 c 466 289 600 1.00 - - - 17.3 - 8.18E-10 - 6.540

6 527 289 600 1.00 - - - 10.4 - 3.94E-10 - 6.730

7 598 289 284 226 600 1.00 - - - 6.0 - 1.78E-10 - 6.774

8 678 289 -497 -204 <30 1.00 - - - 6.2 - negligible - 6.773

9 796 289 <30 1.00 - - - 10.0 - 1.06E-11 - 6.784

10 985 290 <30 1.00 - - - 17.5 - 1.13E-11 - 6.884

11 1107 Changing back to NWC 1.00 - - - 18.0 - 1.64E-10 - 7.073

12 1153 289 285 199 600 1.00 - - - 7.9 - 1.61E-10 - 7.120 a Represents values in the effluent. In the feedwater, conductivity was ≈0.08 μS/cm, and DO was ≈800 to 1000 ppb. b Hold periods at minimum load during the loading cycle and at maximum load during the unloading cycle. c An unexpected rapid crack growth possibly due to broken ligaments. An average value of this test stage is reported.

5.90

6.00

6.10

6.20

6.30

10

20

30

40

50

150 200 250

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

GG 304L SA HAZ, 24 h @ 500oC (Spec. GG3TA-TT)

Test CGRIJR–43

Dose 2.4 dpa

289°C

High–Purity Water

Pre j

Pre k

Pre l 1a

1b 1c 1d

1e

2 Crack length

Kmax

(a)

Figure 6. Crack-length-vs.-time plot for the thermally treated GG 304L SA HAZ specimen ( ≈2.4 

dpa) in simulated BWR water at 289°C. 
17

6.00

6.20

6.40

6.60

6.80

7.00

10

20

30

40

50

320 360 400 440 480 520

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

289°C

High–Purity Water

3 4

5

6

GG 304L SA HAZ, 24 h @ 500o

C (Spec. GG3TA-TT)

Test CGRIJR–43

Dose 2.4 dpa

Crack length

Kmax

Rapid crack growth

(b)

6.50

6.60

6.70

6.80

6.90

7.00

7.10

7.20

7.30

10

20

30

40

50

560 640 720 800 880 960 1040 1120 1200

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

289°C

High–Purity Water

7 8

10

GG 304L SA HAZ, 24 h @ 500oC (Spec. GG3TA-TT)

Test CGRIJR–43

Dose 2.4 dpa

Crack length

Kmax

11

Change back 

to NWC

12

9

HWC

(c)

Figure 6. (Contd.)

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-13 10-12 10-11 10-10 10-9 10-8 10-7

CGRenv (m/s)

CGRair (m/s)

CGRI JR-43, GG3TA-TT

304L SA HAZ, 24 h @ 500oC 

2.4 dpa

289°C, 600 ppb DO

Figure 7. Crack growth rate for GG 304L SA 

HAZ specimen under cyclic load in simulated 

BWR NWC water at ≈290°C.
18

After nearly 340-μm crack advance, environmental enhancement started to appear in test period 1e 

at Kmax ≈ 19 MPa m1/2 and R ≈ 0.3 with a slow/fast sawtooth waveform. Further increases in the load 

ratio (up to ≈0.7) and rise time (up to ≈700 s) did not yield significantly higher CGRs (test period 2-4). 

Starting from period 5, the CGR test was transferred to the constant-load mode. No periodical unloading 

was employed for this sample except at the transitions between two constant load stages. One or two 

unloading cycles to R ≈ 0.7 were carried out between two constant-load test stages to break unbroken 

ligaments. The CGRs in NWC were measured for the test periods at ≈6, 10, and 17 MPa m 1/2, 

respectively. In period 5, a sudden increase in CGR was observed (as shown in Fig. 6), perhaps due to 

broken ligaments. An average value for the entire period 5 is, therefore, considered to be a better 

representation of the CGR for this stage. After the tests in NWC, the DO level in the water environment 

was reduced to simulate HWC by bubbling 4% hydrogen balanced with nitrogen in the system. Under 

the HWC, no crack propagation was detectable at ≈6 MPa m1/2, and CGRs at ≈10 and 17.5 MPa m1/2

were at least one order magnitude lower than the corresponding CGRs in NWC. Figure 8 shows all 

constant-load CGRs obtained from this specimen in both NWC and HWC. 

After the tests in HWC, the DO level in water was recovered by bubbling 1% O2 balanced with N2

in the system during period 11. The environmentally enhanced cracking in NWC was reconfirmed with 

limited data in period 12 at K ≈ 8 MPa m1/2, as shown in Fig. 8. 

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30 35 40

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

CGRI JR-43, GG3TA-TT

GG 304L SA HAZ, 24h @ 500oC, 2.4 dpa 

Closed Symbols: NWC BWR Env.

Open Symbols: HWC BWR Env.

8 x NUREG-0313

Curve

Figure 8. Crack growth rate versus stress 

intensity under constant-loading condition for the

304L SA HAZ specimen thermally treated at 

500°C for 24 h (GG3TA-TT) in simulated BWR 

NWC and HWC water at ≈290°C.

Following the CGR test, a J-R fracture toughness test was performed on the specimen at 289°C in 

NWC. The test was conducted at an extension rate of 0.43 μm/s. Figures 9 and 10 are the load vs. loadline displacement and J-R curve, respectively. The J value of the onset of crack extension (JIC) was 

determined to be 73 kJ/m2 for this specimen. After the J-R test, the final crack size was marked by 

fatigue cycling at room temperature. The specimen was then fractured. Figure 11 shows the final 

fracture surface and a side view of the broken specimen. The final crack length was about 12% greater 

than the DCPD measurement. Accordingly, all DCPD results were scaled up proportionally with the final 

crack length. 
19

0.0

0.5

1.0

1.5

2.0

2.5

3.0

0.0 0.5 1.0 1.5 2.0

Measured Displacement

Loading Pin Displacement

Load (kN)

Displacement (mm)

CGRI JR-43, GG3TA-TT

GG 304L SA HAZ, 24h @ 500oC, 2.4 dpa 

Figure 9. Load vs. load-line 

displacement curve for the

thermally treated 304L SA 

HAZ specimen (GG3TA-TT)

tested in NWC at 289°C.

0.0

50.0

100.0

150.0

200.0

250.0

0.0 0.5 1.0 1.5 2.0

J (kJ/m2)

Crack Extension (mm)

J=140*∆a0.43

JIC=73 kJ/m2

CGRI JR-43, GG3TA-TT

GG 304L SA HAZ, 24h @ 500oC, 2.4 dpa 

Figure 10. Fracture 

toughness J-R result for the 

thermally treated 304L SA 

HAZ specimen (GG3TA-TT)

tested in NWC at 289°C.

Following the CGR test, a J-R fracture toughness test was performed on the specimen at 289°C in 

NWC. The test was conducted at an extension rate of 0.43 μm/s. Figures 9 and 10 are the load vs. loadline displacement and J-R curve, respectively. The J value of the onset of crack extension (JIC) was 

determined to be 73 kJ/m2 for this specimen. After the J-R test, the final crack size was marked by 

fatigue cycling at room temperature. The specimen was then fractured. Figure 11 shows the final 

fracture surface and a side view of the broken specimen. The final crack length was about 12% greater 

than the DCPD measurement. Accordingly, all DCPD results were scaled up proportionally with the final 

crack length. 
20

Figure 11. Photograph of the fracture surface of

the thermally treated 304L SA HAZ specimen 

(GG3TA-TT).

3.2 Type 304 SMA Weld HAZ

3.2.1 As-welded Type 304 SMA Weld HAZ Irradiated to 0.75 dpa (85-7B)

A Type 304 SMA weld HAZ specimen irradiated to ≈0.75 dpa (85-7B) was tested in the simulated 

BWR environment. The objective of this test was to evaluate the cracking behavior of SMA HAZ at the 

IASCC threshold dose level. After the specimen was soaked in simulated BWR water with 600 ppb DO 

for 150 h, fatigue precracking with a triangular waveform was started at 1 Hz. The initial load ratio was 

0.3, and Kmax was about 10 to 15 MPa m1/2. The load ratio was increased to 0.5 after a few hours, and

Kmax was reduced to 10 MPa m1/2. The measured CGR decreased slowly with the increasing load ratio 

and rise time. At around 5×10-10 m/s, environmental enhancement became evident. Details of the test 

conditions are given in Table 6. Figure 12 shows the crack-length-vs.-time curve of this test. Figure 13 is 

a plot of the cyclic CGR data showing enhanced cracking in simulated BWR water.

Constant-load CGR tests began at period 4. Figure 14 shows the constant-load CGR results 

obtained from this specimen. In NWC, the constant-load CGR for K ≈ 5 MPa m1/2 is about 

2.39×10-10 m/s at ≈290°C, much higher than that of nonirradiated sensitized 304 SSs (NUREG -0313 

curve). Additional tests were also performed in NWC on the specimen at ≈10, 15 and 19 MPa m 1/2. All 

constant-load CGR results for the current specimen are plotted in Fig. 14 along with the NUREG-0313 

disposition curve. The CGRs of this SMA weld HAZ specimen are high even at this low dose level. 

Besides the tests at 290°C, the specimen was also tested in NWC at ≈270°C and 280°C with K ≈ 

15 MPa m1/2. For every temperature change, sufficient time was allowed to re-stabilize the system (e.g., 

periods 7a, 8a, and 9a). The DCPD measurement was also re-initiated to eliminate any effect introduced 

by the low test temperature. An increasing trend of the CGR was observed with rising test temperature. 

The temperature dependence of CGR is plotted in Fig. 15. For this specimen, the apparent activation 

energy for environmentally assisted cracking was determined from Fig. 15 to be 188 kJ/mol in NWC at 

270-290°C.
21

Table 6. Test results for the Type 304 SMA HAZ specimen at as-welded condition (85-7B, 0.75 dpa) in 

BWR water.

Test

Test 

Time,

Test

Temp.,

ECP,a

mV (SHE)

O2

Conc.,a RLoad

Rise 

Time,

Return 

Time,

Hold 

Time,b Kmax, ∆K,

CGR in 

Env.,

CGR in 

Air,

Crack 

Length,

Period h °C Pt Steel ppb Ratio s s s MPa m1/2 MPa m1/2 m/s m/s mm

5.834

Pre a 3 289 600 0.32 0.41 0.41 0.09/0.09 9.8 6.7 negligible 6.95E-09 5.832

Pre b 4 289 0.30 0.44 0.44 0.06/0.06 15.3 10.7 3.23E-08 3.03E-08 5.913

1a 5 289 0.58 0.80 0.80 0.20/0.20 15.2 6.3 2.57E-08 3.89E-09 5.946

1b 9 289 0.50 0.76 0.76 0.24/0.24 10.1 5.1 6.57E-09 1.80E-09 5.975

1c 10 289 0.70 0.66 0.66 0.34/0.34 10.1 3.0 7.19E-09 4.57E-10 5.991

2a 13 289 0.72 3.22 0.64 1.78/0.36 10.0 2.8 1.34E-09 7.40E-11 5.999

2b 21 290 273 226 0.72 6.45 0.64 3.55/0.36 10.0 2.8 6.22E-10 3.85E-11 6.001

2c 26 289 277 225 0.69 19.9 7.94 10.1/4.0 10.1 3.1 5.40E-10 1.59E-11 6.016

2d 44 289 281 232 0.69 200 8.00 100/4.0 10.1 3.2 4.48E-10 1.71E-12 6.034

3a 68 290 0.68 154 6.16 146/5.8 5.2 1.7 4.89E-10 2.69E-13 6.055

3b 92 289 281 235 600 0.74 458 5.50 542/6.5 5.1 1.3 4.79E-10 4.68E-14 6.073

4 141 289 281 238 1.0 - - - 5.2 - 2.39E-10 - 6.116

5 171 289 283 242 1.0 - - - 9.9 - 5.58E-10 - 6.200

6 199 289 283 244 1.0 - - - 15.3 - 9.38E-10 - 6.263

7 220 289 281 243 1.0 - - - 19.1 - 1.33E-09 - 6.486

7a 249 Decrease temperature 1.0 - - - 3.0 - - - -

8 288 269 285 247 1.0 - - - 15.4 - 2.10E-10 - 6.547

8a 308 Increase temperature 1.0 - - - 4.8 - - - -

9 337 280 285 245 1.0 - - - 15.3 - 4.24E-10 - 6.597

9a 355 Increase temperature 1.0 - - - 4.8 - - - -

9b 358 289 1.0 - - - 15.3 - 7.74E-10 - 6.630

9c 373 Reduce DO 1.0 - - - 15.3 - - - -

10 537 289 -488 -183 <50 1.0 - - - 15.4 - 6.36E-12 - 6.662

10a 574 Increase DO 1.0 - - - 15.2 - - - -

10b 585 289 259 236 600 1.0 - - - 15.4 - 7.77E-10 - 6.700 a Represents values in the effluent. In the feedwater, conductivity was ≈0.08 μS/cm, and DO was ≈800 to 1000 ppb. b Hold periods at minimum load during the loading cycle and at maximum load during the unloading cycle. 

5.80

5.85

5.90

5.95

6.00

6.05

6.10

4

8

12

16

20

24

28

0 5 10 15 20 25 30 35 40

Crack Length (mm)

K (MPa m0.5)

Time (h)

Type 304 SMA HAZ, as-welded (Spec. 85-7B)

Test CGRIJR–45

Dose 0.75 dpa

289°C

High–Purity Water

Pre a

Pre b 1a

2d 2c 2b 2a

1b

1c

Kmax

Crack length

(a)

Figure 12. Crack-length-vs.-time plot for the as-welded Type 304 SMA HAZ specimen (85-7B, ~0.75 

dpa) in BWR water at 289°C. 
22

5.90

6.00

6.10

6.20

6.30

6.40

6.50

6.60

5

10

15

20

25

30

35

40

50 100 150 200 250

Crack Length (mm)

K (MPa m0.5)

Time (h)

Type 304 SMA HAZ, as-welded (Spec. 85-7B)

Test CGRIJR–45

Dose 0.75 dpa

289°C

High–Purity Water

7

6

5

4

3a

K

Crack length

3b

7a,

Decrease

T to 270oC

(b)

6.40

6.45

6.50

6.55

6.60

6.65

6.70

6.75

6.80

5

10

15

20

25

30

35

40

250 300 350 400 450 500 550 600

Crack Length (mm)

K (MPa m0.5)

Time (h)

Type 304 SMA HAZ, as-welded (Spec. 85-7B)

Test CGRIJR–45

Dose 0.75 dpa

289°C

High–Purity Water

10a,

10, Increase DO

T=290o 9, C, HWC

T=280oC

8, 

T=270oC

K

Crack length

8a, 

Increase 

T to 280oC

9a, 9b, 9c,

Increase T=290oC, 

and reduce DO

10b

(c)

Figure 12. (Contd.)

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-13 10-12 10-11 10-10 10-9 10-8 10-7

CGRenv (m/s)

CGRair (m/s)

CGRI JR-45, 85-7B

304 SMA HAZ, as-welded

0.75 dpa

289°C, 600 ppb DO

Figure 13. Cyclic crack growth rate for the 

as-welded Type 304 SMA HAZ specimen 

(85-7B) irradiated to 0.75 dpa.
23

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25

NWC, 290oC

HWC, 290oC

NWC, 280oC

NWC, 270oC

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

CGRI JR-45, 85-7B

304 SS SMA, as-welded

0.75 dpa

Test Conditions:

Figure 14. Crack growth rate versus stress 

intensity under constant-loading conditions 

for the as-welded Type 304 SMA HAZ 

specimen (85-7B) irradiated to 0.75 dpa.

10-10

10-9

0.00172 0.00176 0.0018 0.00184 0.00188

310 300 290 280 270 260 250

Crack growth rate (m/s)

1/T (1/K)

Temperature (oC)

Q=188 kJ/mol

304 SMA HAZ, as-welded, 0.75 dpa

Tested in NWC

Figure 15. Temperature dependence of 

crack growth rate for the as-welded 304 

SMA HAZ specimen (85-7B, 0.75 dpa)

The effect of HWC on constant-load CGR was investigated at ≈290°C with this specimen. After 

CGR tests in NWC, the DO level in the simulated environment was reduced by bubbling 4% hydrogen 

mixed with nitrogen in the system. Consequently, the CGR at ≈15 MPa m1/2 was nearly two orders of 

magnitude lower than that in NWC. 

After the CGR test, a J-R test was performed to obtain the fracture toughness of this specimen in 

NWC. The test was conducted under displacement control at a constant extension rate of 0.43 μm/s. 

During the test, loading was interrupted periodically to measure the crack length. A relaxation time of 

30 s was allowed before each DCPD measurement. Figures 16 and 17 show the obtained load vs. loadline displacement and the corresponding J-R curve, respectively. The power-law fitting of the data 

suggests a JIC of 94 kJ/m2 at 0.2-mm crack extension. 

The final crack front was marked by fatigue cycling at room temperature in air. The specimen was 

then fractured, and optical images of the fracture surface were taken (Fig. 18). The final crack length 

measured on the fracture surface was approximately 13% greater than the value determined by the DCPD 

method. Accordingly, the previous crack extension results obtained from DCPT measurement were 

scaled proportionately. 
24

0.0

1.0

2.0

3.0

4.0

0 0.5 1 1.5 2 2.5 3 3.5

Loading Pin Displacement

measured Displacement

Load (kN)

Displacement (mm)

CGRI JR–45, 85-7B

304 SMA HAZ, as-welded, 0.75 dpa

289°C High-Purity Water

Figure 16. Load vs. load-line

displacement curve for the as-welded 

Type 304 SMA HAZ specimen (85-7B)

tested in NWC at 289°C.

0

100

200

300

400

500

0 0.5 1 1.5 2 2.5 3

J (kJ/m2)

Crack Extension (mm)

CGRIJR–45, 85-7B

304 SMA HAZ, as-welded, 0.75 dpa 

289°C High-Purity Water

J = 322∆a0.86

JQ = 94 kJ/m2

Figure 17. Fracture toughness J-R 

curves for the as-welded Type 304 

SMA HAZ specimen (85-7B) tested in 

NWC at 289°C.

Figure 18. Photograph of the fracture surface of 

the as-welded 304 SMA HAZ specimen (85-7B) 

irradiated to 0.75 dpa.

3.2.2 Thermally Treated Type 304 SMA Weld HAZ Irradiated to 2 dpa (85-2ATT)

A CGR test was performed on a Type 304 SMA weld HAZ specimen (85-2ATT) irradiated to 2.1 

dpa. The specimen was thermally treated at 500°C for 24 hours prior to irradiation. The objective of this 

test was to investigate the effect of low-temperature sensitization and to provide additional information on

the dose dependence of IASCC susceptibility for the Type 304 SMA weld HAZ. Table 7 summarizes 

details of test periods, and Fig. 19 shows the crack-extension-vs.-time curve for this specimen. 
25

Table 7. Test results for a thermally treated 304 SMA HAZ specimen irradiated to 2 dpa (85-2ATT). 

Test

Test 

Time,

Test

Tem

p

ECPa

mV (SHE)

O2

Conc.,a RLoad

Rise 

Time,

Return 

Time,

Hold 

Time,b Kmax, ∆K,

CGR in 

Env., CGR in 

air

Crack 

Length,

Period h °C Pt Steel ppb Ratio s s s MPa m1/2 MPa m1/2 m/s m/s mm

5.771

Pre a 3 290 - - 600 0.43 0.16 0.16 0.09/0.09 6.2 3.5 - 2.24E-9 5.771

Pre b 6 290 256 223 0.32 0.39 0.39 0.11/0.11 10.0 6.8 - 7.59E-9 5.773

Pre c 7 290 259 226 0.30 0.43 0.43 0.07/0.07 15.3 10.8 1.15E-08 3.07E-8 5.802

Pre d 23 290 268 234 0.29 0.43 0.43 0.07/0.07 15.5 10.9 3.14E-08 3.21E-8 5.855

1a 25 290 269 235 0.29 0.86 0.86 0.14/0.14 15.3 10.8 1.79E-08 1.55E-8 5.903

1b 30 290 267 233 0.49 0.807 0.81 0.19/0.19 15.4 7.8 3.46E-09 6.95E-9 5.937

2a 46 290 267 233 0.51 2.4 0.8 0.6/0.2 15.2 7.5 - 2.05E-9 5.944

2b 51 290 268 234 0.49 1.61 1.61 0.39/0.39 15.3 7.8 - 3.40E-9 5.943

2c 53 290 267 232 0.49 0.40 0.40 0.1/0.1 15.3 7.8 - 1.37E-8 5.947

3a 55 290 266 231 0.29 0.43 0.43 0.07/0.07 15.3 10.8 - 3.13E-8 5.949

3b 74 290 271 235 0.30 0.45 0.45 0.05/0.05 15.0 10.5 - 2.77E-8 5.936

3c 75 290 269 235 0.29 0.46 0.46 0.04/0.04 16.6 11.8 2.36E-08 3.94E-8 5.989

3d 76 290 269 234 0.30 0.90 0.90 0.1/0.1 15.1 10.6 - 1.42E-8 5.994

3e 78 290 269 234 0.28 0.91 0.91 0.09/0.09 16.7 11.9 1.27E-08 2.04E-8 6.029

3f 80 290 269 235 0.51 0.88 0.88 0.12/0.12 16.7 8.3 6.81E-09 8.01E-9 6.059

4a 114 290 269 235 0.31 4.54 0.91 0.46/0.09 16.6 11.4 5.97E-11 3.67E-9 6.069

4b 167 290 274 239 0.31 2.72 0.91 0.28/0.09 16.7 11.4 1.41E-10 6.16E-9 6.084

4c 215 290 273 240 600 0.32 9.08 0.91 0.92/0.09 16.9 11.5 7.61E-11 1.90E-9 6.099

4d 244 290 274 240 0.51 8.74 0.85 1.26/0.15 16.9 8.2 9.12E-11 7.86E-10 6.109

4e 272 290 275 243 0.71 8.04 0.80 1.96/0.20 16.9 4.9 1.74E-10 1.81E-10 6.122

5a 312 290 - - 0.70 24.32 4.05 5.68/0.95 16.9 5.1 1.24E-10 6.83E-11 6.137

5b 335 290 - - 0.70 244 9.74 56/2.26 16.9 5.1 1.50E-10 7.07E-12 6.145

5c 337 290 - - 0.66 829 9.95 171/2.05 17.0 5.8 4.22E-09 3.01E-12 6.160

5d 345 290 264 218 0.69 723 8.68 277/3.32 10.1 3.1 5.30E-10 4.62E-13 6.169

5e 358 290 273 231 0.68 616 7.40 384/4.60 6.1 1.9 5.73E-10 1.10E-13 6.185

6 431 290 272 236 600 1.0 - - - 6.2 - 2.00E-10 - 6.243

7 502 290 267 239 1.0 - - - 10.3 - 2.76E-10 - 6.327

8 526 290 - - 1.0 - - - 15.3 - 7.28E-10 - 6.455

8a 550 Transition to 270°C

9 622 270 269 236 600 1.0 - - - 15.2 - 2.32E-10 - 6.488

9a 655 Transition to 280°C

10 790 280 274 244 600 1.0 - - - 15.1 - 7.26E-11 - 6.529

10a 815 Transition to 290°C

11 1200 289 269 244 600 1.0 - - - 15.3 - 5.71E-11 - 6.619

11a 1207 290 -487 -128 1.0 - - - 15.4 - negligible - 6.620

12 1420 290 -485 -291 <30 1.0 - - - 15.3 - negligible - 6.616 a Represents values in the effluent. In the feedwater, conductivity was ≈0.08 μS/cm and DO was ≈800 to 1000 ppb. b Hold periods at minimum load during the loading cycle and at maximum load during the unloading cycle. 

The initial fatigue precracking for this sample was performed with a triangular waveform, at 

Kmax=10-15 MPa m1/2, R≈0.3, and 1 Hz frequency. After about 250-μm crack extension, the triangular

waveform was changed to slow/fast sawtooth waveform with 5-s rise time and 1-s return time. 

Environmentally enhanced cracking was difficult to establish in this specimen, and the transition from 

triangular to slow/fast sawtooth loading had to be repeated several times between test periods 2 and 4. 

Finally, after another ≈100 μm crack extension, environmental cracking started to appear in test period 4e. 

During test periods 5a-5e, cyclic CGRs were obtained at Kmax≈17, 10, and 6 MPa m1/2. After the cyclic 

CGR test, the test condition was switched to constant load, and CGRs were obtained at K≈6, 10, and 15 

MPa m1/2 in NWC at 290°C. Constant-load CGR tests were also performed at ≈270 and 280°C in NWC. 

In HWC, no measurable crack extension was observed for K≈15 MPa m1/2 at 290°C after more than 200 

hours. 
26

6.00

6.05

6.10

6.15

6.20

4

8

12

16

20

24

28

32

80 120 160 200 240 280 320 360

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

Type 304 weld HAZ, 24 h @ 500oC (Spec. 85-2ATT)

Test CGRIJR–46

Dose 2.1 dpa

289°C

High–Purity Water

4a 4b 4c 4d 4e

5a

5b

5e

5d

(a)

6.10

6.20

6.30

6.40

6.50

4

8

12

16

20

24

28

32

360 400 440 480 520

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

Type 304 weld HAZ, 24 h @ 500oC (Spec. 85-2ATT)

Test CGRIJR–46

Dose 2.1 dpa

289°C

High–Purity Water

6

7

8 8a,

Reduce T

to 270o

C

(b)

6.20

6.30

6.40

6.50

6.60

6.70

4

8

12

16

20

24

28

32

550 600 650 700 750 800

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

Type 304 weld HAZ, 24 h @ 500oC (Spec. 85-2ATT)

Test CGRIJR–46

Dose 2.1 dpa

289°C

High–Purity Water

9

9a

increase T

to 280o

C

10

10a

increase T

to 290o

C

(c)

Figure 19. Crack-length-vs.-time plot for the thermally treated Type 304 SMA HAZ specimen (85-2ATT, 

~2 dpa) in BWR water at 289°C.
27

6.30

6.40

6.50

6.60

6.70

6.80

4

8

12

16

20

24

28

32

900 1000 1100 1200 1300 1400

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

Type 304 weld HAZ, 24 h @ 500oC (Spec. 85-2ATT)

Test CGRIJR–46

Dose 2.1 dpa

289°C

High–Purity Water

11

11a

Reduce DO

12

(d)

Figure 19. (Contd.)

Figures 20 and 21 show the cyclic and constant-load CGRs obtained from this HAZ specimen. For 

the cyclic CGR test, environmentally enhanced cracking was evident at about 10-10 m/s. The influence of 

environment on cyclic CGR increased rapidly with the increasing rise time. For the constant-load CGR 

test, the CGRs at ≈290°C were about eight times higher than the NUREG-0313 curve. The CGRs 

obtained at 270°C and 280°C were significantly lower. No measurable crack extension was observed in 

the constant-load CGR test at 290°C in HWC. 

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-13 10-12 10-11 10-10 10-9 10-8 10-7

CGRenv (m/s)

CGRair (m/s)

304 SMA HAZ, 24 h @ 500oC 

2.1 dpa

289°C, 600 ppb DO

Figure 20. Cyclic CGRs for the thermally 

treated Type 304 SS SMA weld HAZ 

specimen (85-2ATT) irradiated to 2.1 dpa.
28

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30

290oC

270oC

280oC

290oC

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NWC

NUREG-0313

Curve

304 SS SMA, 24 h @ 500oC

~2.15 dpa

8 x NUREG-0313

Curve

HWC

Figure 21. Constant-load CGRs for the

thermally treated Type 304 SS SMA weld 

HAZ specimen (85-2ATT) irradiated to 2.1

dpa.

A fracture toughness J-R test was performed on the specimen in NWC after the CGR test. The J-R 

test was conducted at a constant extension rate of 0.43 μm/s. During the test, loading was interrupted 

periodically and crack extension was measured by the DCPD method. A 30-s relaxation time was 

allowed prior to each DCPD measurement. Figure 22 shows the load vs. load-line displacement curve 

and the J-R data. A power-law curve fitting on the J-R result yielded a JIC of about 34 kJ/m2 (Fig. 22b). 

0.0

1.0

2.0

3.0

4.0

0 0.5 1 1.5 2 2.5

Loading Pin Displacement

Measured Displacement

Load (kN)

Displacement (mm)

304 SMA, 24h+500oC (Spec. 85-2ATT)

Test CGRIJR-46

Dose: ~ 2 dpa

289°C High-Purity Water

0

50

100

150

200

0 0.5 1 1.5 2 2.5

J (kJ/m2)

Crack Extension (mm)

304 SMA, 24h+500oC (Spec. 85-2ATT)

Test CGRI JR-46

Dose: ~ 2 dpa

J = 86∆a0.6

JIC = 34 kJ/m2

(a) (b)

Figure 22. (a) Load vs. load-line displacement and (b) fracture toughness J-R curve for the thermally 

treated 304 SMA weld HAZ specimen (85-2ATT). 

After the J-R test, the specimen was fatigued at room temperature in air to mark the final crack 

front. The specimen was then fractured, and optical images of the fracture surface were taken. The final 

crack length measured on the fracture surface was approximately 11% greater than the value determined 

previously by the DCPD method. Accordingly, the crack extension results obtained from DCPT 

measurement were scaled proportionately. As shown in Fig. 23, an out-of-plane crack propagation had 
29

occurred during the J-R test. The complex weld microstructure along the fusion line may have 

contributed to the out-of-plane crack propagation.

(a) (b)

Figure 23. Optical images of the thermally treated 304 SS SMA HAZ specimen (85-2ATT) After 

fracturing: (a) fracture surface and (b) side view. 
30

3.3 Sensitized Type 304 SS Irradiated to 0.75 dpa

3.3.1 Type 304 SS Sensitized at 600°C for 10.5 h (85-1TT)

A CGR test was performed on a Type 304 SS specimen sensitized at 600°C for 10.5 h (Specimen 

85-1TT). The objective of this test was to investigate the effect of sensitization on IASC susceptibility 

around “threshold” dose level. The specimen was irradiated to 0.75 dpa at ≈290°C in the Halden reactor. 

The test was started in high-purity water with ≈350 ppb DO and a flow rate of ≈ 20 mL/min. The 

specimen was fatigue precracked at R = 0.25, Kmax = 18.0 MPa m1/2, triangular waveform, and 2-Hz 

frequency. After 130-μm crack advance, R was increased incrementally to 0.7, and the waveform was 

changed to a slow/fast sawtooth with rise times of 300 or 1000 s and a return time of 12 s. The test 

conditions and results are summarized in Table 8, and the crack-extension-vs.-time plots are shown in 

Fig. 24. Environmental enhancement started to appear during period 2 at ≈70 h test duration (Fig. 24a). 

Environmental effects on CGR are more evident in the subsequent test stages. Figure 25 summarizes 

cyclic CGR results on this specimen.

Table 8. Test results for a Type 304 specimen sensitized at 600°C for 10.5 h (85-1TT, 0.75 dpa) in BWR 

water.

Test

Test 

Time,

Test

Temp.,

ECP,

a

mV (SHE)

O2

Conc.,a RLoad

Rise 

Time,

Return 

Time,

Hold 

Time,b Kmax, ∆K,

CGR in 

Env.,

CGR in 

Air,

Crack 

Length,

Period h °C Pt Steel ppb Ratio s s s MPa m1/2 MPa m1/2 m/s m/s mm

5.804

Pre 30 288 - - 350 0.25 0.22 0.22 0.03/0.03 18.0 13.5 9.43E-8 1.21E-7 5.964

1a 38 288 - - 350 0.51 8.1 1.6 1.9/0.4 17.3 8.4 4.99E-10 9.29E-10 5.976

1b 48 286 - - 350 0.52 8.1 1.6 1.9/0.4 17.3 8.4 negligible 9.12E-10 5.976

2 56 286 - - 350 0.42 4.2 0.8 0.8/0.2 17.5 10.2 4.93E-09 3.00E-09 6.057

3 120 286 - - 350 0.56 235 9.4 65/2.6 17.3 7.6 8.44E-10 2.37E-11 6.167

4 175 287 - - 350 0.75 674 8.1 326/3.9 17.6 4.5 4.63E-10 1.66E-12 6.226

5 288 287 - - 350 1.0 - - - 17.9 - 2.19E-10 - 6.315

6 456 287 - - 350 1.0 - - - 16.3 - 1.51E-10 - 6.417

7 504 287 275 230 350 1.0 - - - 20.4 - 2.89E-10 - 6.467 a Represents values in the effluent. In the feedwater, conductivity was ≈0.08 μS/cm, and DO was ≈800 to 1000 ppb. b Hold periods at minimum load during the loading cycle and at maximum load during the unloading cycle. 

Pre

5.80

5.85

5.90

5.95

6.00

6.05

6.10

6.15

6.20

14

16

18

20

22

24

40 60 80 100 120

Crack Length (mm)

K (MPa m0.5)

Time (h)

Type 304 SS, Sensitized 10.5 h at 600°C (Spec. 85-1TT)

Test CGRI JR–38

Dose 0.75 dpa

289°C

High–Purity Water

Kmax

Crack Length

2

3

1a 1b

(a)

Figure 24. Crack-length-vs.-time plot for the Type 304 SS specimen sensitized at 600°C for 10.5 h 

(85-1TT, ~0.75 dpa) in BWR water at 289°C. 
31

6.10

6.15

6.20

6.25

6.30

6.35

14

16

18

20

22

24

120 160 200 240 280

Crack Length (mm)

K (MPa m0.5)

Time (h)

289°C

High–Purity Water

Kmax

5

Crack Length

Type 304 SS, Sensitized 10.5 h at 600°C (Spec. 85-1TT)

Test CGRI JR–38

Dose 0.75 dpa

4

(b)

6

6.25

6.30

6.35

6.40

6.45

6.50

14

16

18

20

22

24

26

28

280 320 360 400 440 480 520

Crack Length (mm)

K (MPa m0.5)

Time (h)

Type 304 (Heat 10285) Sensitized 10.5 h at 600°C

Test CGRIJR–38 (Spec. 85-1TT)

Fluence 0.5 x 1021 n/cm2

289°C

K High–Purity Water max

Crack Length

7

(c)

Figure 24. (Contd.)
32

10-12

10-11

10-10

10-9

10-8

10-7

10-12 10-11 10-10 10-9 10-8 10-7

CGRenv (m/s)

CGRair (m/s)

CGRI JR-38, 85-1TT

304 SS, sensitized at 600oC for 10.5 h

0.75 dpa

289oC, 500 ppb DO

Figure 25. Cyclic CGR for the Type 304 

specimen sensitized at 600°C for 10.5 h (85-1TT) 

and irradiated to 0.75 dpa.

Starting from test period 5, the specimen was subjected to constant load with Kmax =

17.6 MPa m1/2. After ≈100 h, Kmax was decreased to 15.8 MPa m1/2. After another ≈170 h, Kmax was 

increased to 19.7 MPa m1/2. The CGR test was terminated at 504 h. Figure 26 shows the CGR results 

under constant loading, and the NUREG-0313 curve is also plotted for reference. 

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30 35 40

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

CGRI JR-38, 85-1TT

304 SS, sensitized 10.5 h at 600°C

0.75 dpa

NWC, 289oC

Figure 26. Crack growth rate versus stress 

intensity under constant-loading conditions for 

the Type 304 specimen sensitized at 600°C for 

10.5 h (85-1TT) and irradiated to 0.75 dpa.

After the CGR test, a J-R curve test was performed on the specimen at 289°C in NWC (≈350 ppb 

DO). The test was conducted at a constant extension rate of 0.43 μm/s; the DC potential drop method was 

used to measure crack extension. The load vs. load-line displacement curve is shown in Fig. 27. The J-R 

data were analyzed to obtain a power-law curve for the material. The DC potential data were corrected to 

account for the effects of plasticity on the measured potential. The fracture toughness J-R curve for 

Specimen 85-1TT in NWC is shown in Fig. 28, and the JQ*

 

* A provisional value of JIC.

is 182 kJ/m2.
33

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

0 0.5 1 1.5 2 2.5 3 3.5

Loading -Pin Displacement

Measured Displacement

Load (kN)

Displacement (mm)

CGRIJR–38, 85-1TT

304 SS, Sensitized 10.5 h at 600°C, 0.75 dpa

289oC, High-Purity Water

Figure 27. Load vs. load-line 

displacement curve for the Type 304 

specimen sensitized at 600°C for 

10.5 h (85-1TT) and tested in NWC at 

289°C.

0

100

200

300

400

500

600

700

0 0.5 1 1.5 2 2.5

J (kJ/m2)

Crack Extension (mm)

Test CGRI JR–38, 85-1TT

304 SS, Sensitized 10.5 h at 600°C, 0.75 dpa

289oC, High-Purity Water

J = 404∆a0.64

JQ = 182 kJ/m2

Figure 28. Fracture toughness J-R 

curve for the Type 304 specimen 

sensitized at 600°C for 10.5 h (85-1TT)

and tested in NWC at 289°C.

The final crack front was marked by fatigue cycling at room temperature after the J-R test. 

Specimen 85-1TT was then fractured, and the final crack length of both halves of the fractured specimen 

was measured from the photograph of the fracture surface (Fig. 29). The actual crack extension was 

≈24% greater than the value determined from the DC potential measurements. The crack extensions 

estimated from the DC potential method were adjusted accordingly. 
34

Figure 29. Photograph of the fracture surface of the Type 

304 specimen sensitized at 600°C for 10.5 h (85-1TT) and 

irradiated to 0.75 dpa.

3.3.2 Type 304 SS Sensitized at 600°C for 24 h (85-5A1TT)

A CGR test was performed in simulated BWR environment on a sensitized Type 304 SS specimen

irradiated to ≈0.75 dpa at ≈290°C (85-5A1TT). The specimen was sensitized at 600°C for 24 h prior to 

irradiation. The objective of this test was to compare with the previous test (85-1TT) to evaluate IASCC 

susceptibility at a higher degree of sensitization. Table 9 shows all test conditions and CGR results of this 

specimen, and Fig. 30 is the crack-length-vs.-time plot of this test. 

Table 9. Test results for a Type 304 specimen sensitized at 600°C for 24 h (85-5A1TT, 0.75 dpa) in 

BWR water.

Test

Test 

Time,

Test

Temp.,

ECP,

a

mV (SHE)

O2

Conc.,a RLoad

Rise 

Time,

Return 

Time,

Hold 

Time,b Kmax, ∆K,

CGR in 

Env.,

CGR in 

Air,

Crack 

Length,

Period h °C Pt Steel ppb Ratio s s s MPa m1/2 MPa m1/2 m/s m/s mm

5.860

Pre 4 291 294 232 700 0.30 0.90 0.90 0.10/0.10 17.3 12.0 1.79E-8 2.16E-8 5.935

1a 6 291 291 231 700 0.52 0.86 0.86 0.14/0.14 17.0 8.2 1.11E-8 7.99E-9 5.975

1b 10 291 291 231 700 0.50 0.85 0.85 0.15/0.15 15.0 7.5 4.93E-9 5.88E-9 5.999

1c 21 291 293 233 700 0.50 8.53 8.53 1.47/1.47 15.1 7.6 1.87E-9 6.02E-10 6.031

1d 29 290 700 0.50 25.6 10.2 4.4/1.8 15.1 7.6 8.69E-10 2.05E-10 6.043

2a 33 290 700 0.70 23.2 9.30 6.7/2.7 15.0 4.5 1.10E-9 4.82E-11 6.053

2b 45 290 700 0.70 223 9.33 76.7/2.7 15.1 4.6 6.13E-10 5.09E-12 6.071

2c 68 290 700 0.70 777 9.32 223/2.7 15.1 4.6 3.17E-10 1.53E-12 6.092

3 126 290 285 223 700 1.0 - - - 15.4 - 1.34E-10 - 6.142

4 173 290 289 228 700 1.0 - - - 10.3 - negligible - 6.124

5 341 290 281 227 700 1.0 - - - 18.2 - 8.16E-11 - 6.192

6 480 290 281 233 700 1.0 - - - 10.3 - 6.25E-11 - 6.215

7 549 290 -495 -240 20 1.0 - - - 10.2 - negligible - 6.209

8 651 290 20 1.0 - - - 15.5 - negligible - 6.236

9 838 290 -414 -413 20 1.0 - - - 17.5 - negligible - 6.241

9a 960 Increase DO - - - - 17.5 - negligible - 6.269

10 1197 290 270 216 600 1.0 - - - 17.5 - 3.65E-11 c - 6.315

10a 1223 Reduce temperature 1.0 - - - 17.6 - 4.64E-10 - 6.350

11 1390 272 292 232 600 1.0 - - - 17.7 - 5.70E-11 - 6.385 a Represents values in the effluent. In the feedwater, conductivity was ≈0.08 μS/cm, and DO was ≈800 to 1000 ppb. b Hold periods at minimum load during the loading cycle and at maximum load during the unloading cycle. c Average value of the whole test stage. CGR at the end of the test stage was 6.50E-11 m/s.
35

5.85

5.90

5.95

6.00

6.05

6.10

6.15

6.20

6.25

0

5

10

15

20

25

30

35

40

0 50 100 150

Crack Length (mm)

K (MPa m0.5)

Time (h)

Type 304 SS, Sensitized @ 600oC for 24 h (Spec. 85-5A1TT)

Test CGRI–44

Dose 0.75 dpa

289°C

High–Purity Water

Pre, 1a

1b 1c 1d 2a 2b

2c

3

4

Crack length

Kmax

(a)

6.05

6.10

6.15

6.20

6.25

6.30

6.35

0

5

10

15

20

25

30

35

40

200 300 400 500 600

Crack Length (mm)

K (MPa m0.5)

Time (h)

Type 304 SS, Sensitized @ 600oC for 24 h (Spec. 85-5A1TT)

Test CGRI–44

Dose 0.75 dpa

289°C

High–Purity Water

5

6

Crack length 7, HWC

K

8, HWC

(b)

6.15

6.20

6.25

6.30

6.35

6.40

6.45

6.50

0

5

10

15

20

25

30

35

40

700 800 900 1000 1100 1200 1300 1400

Crack Length (mm)

K (MPa m0.5)

Time (h)

Type 304 SS, Sensitized @ 600oC for 24 h (Spec. 85-5A1TT)

Test CGRI–44

Dose 0.75 dpa

289°C

High–Purity Water

9, HWC 10

11,

T=272oC

Crack length

K

9a, 

Increase DO

10a,

Reduce Temp.

A load trip

(c)

Figure 30. Crack-length-vs.-time plot for the Type 304 SS specimen sensitized at 600°C for 24 h 

(85-5A1TT, ~0.75 dpa) in BWR water. 
36

The specimen was first exposed to high-purity water with 600 ppb DO for over 70 h before the test. 

After the environmental condition was stabilized, fatigue precracking was initiated with a triangular 

waveform at R = 0.3, Kmax = 17 MPa m1/2, and 0.5-Hz frequency. After ≈70-μm crack advance, the load 

ratio was increased to 0.5, and the Kmax was reduced to ≈15 MPa m1/2. Environmentally enhanced 

cracking was readily achieved in this sensitized specimen. After about 170-μm crack advance, the 

triangular waveform was replaced by slow/fast sawtooth loading, and the load ratio was further increased 

to 0.7. With additional ≈60-μm crack extension, the test condition was switched to a constant load at test 

period 3. 

The constant-load CGR is about ≈1.3 4×10-10 m/s for test period 3 at K ≈ 15 MPa m1/2. An 

unexpected overload caused by a load trip occurred at the end of test period 3. Subsequently, no crack 

extension was observed during test period 4 at K ≈ 10 MPa m1/2 even with periodical unloading. To reestablish crack growth, a higher K value was applied on the specimen in test period 5. After about 50 h

and several loading and unloading cycles, crack extension was re-activated, and an average CGR of 

8.16×10-11 m/s was registered at K ≈ 18 MPa m1/2 for test period 5. An average CGR of 6.25×10-11 m/s 

was also recorded in test period 6 for K ≈ 10 MPa m1/2. 

A constant-load CGR test was also carried out in HWC on this sensitized specimen. No crack 

extension was detectable in HWC for stress levels of 10, 15, and 17 MPa m1/2. Following the CGR test 

in HWC, the DO level in the system was recovered to NWC during test period 9a. Test period 10 was 

used to re-establish the CGR observed during the earlier test periods under NWC. Once a comparable 

CGR was achieved, the autoclave temperature was reduced to prepare for a test at a lower temperature. In 

test period 11, the CGR was measured at ≈270°C and K≈17.7 MPa m1/2. All constant-load CGR results 

are plotted in Fig. 32. 

10-12

10-11

10-10

10-9

10-8

10-7

10-12 10-11 10-10 10-9 10-8 10-7

CGRenv (m/s)

CGRair (m/s)

CGRI-44, 85-5A1TT

304 SS, sensitized at 600oC for 24 h

0.75 dpa

289°C, 500 ppb DO

Figure 31. Cyclic CGR for the Type 304 

specimen sensitized at 600°C for 24 h (85-

5A1TT) and irradiated to 0.75 dpa.

After the CGR test, a DCPD pin was found broken in the subsequent J-R test. The J-R test had to 

be terminated. Before the specimen was pulled open, the final crack front was marked by fatigue cycling 

at room temperature. The actual crack length of both halves of the fractured specimen was measured 

from the photograph of the fracture surface (Fig. 33). The crack extension was ≈8% greater than the 
37

value determined from the DC potential measurements. Accordingly, the crack extensions estimated 

from the DC potential method were adjusted proportionally.

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25

NWC, ~ 290oC

HWC, ~ 290oC

NWC, ~ 270oC

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313 Test Conditions:

Curve

CGRI-44, 85-5A1TT

304 SS, Sensitized 24 h at 600oC

0.75 dpa

K2.161 power

curve fit

Figure 32. Crack growth rate 

versus stress intensity under 

constant-loading conditions for 

the Type 304 specimen sensitized 

at 600°C for 24 h (85-5A1TT) and 

irradiated to 0.75 dpa.

The temperature dependence of the crack growth rate is shown in Fig. 34 at K = 17.4 MPa m1/2 for 

the sensitized 304 SS. Since the CGRs measured at low and high temperatures were at different K values 

for this specimen (as shown in Fig. 32), an expression obtained from a power curve fit (~K2.161) with all 

constant-load data points at 290°C was used to extrapolate a CGR at the same value of K that the CGR 

was obtained at 270°C. Based on this analysis, the apparent activation energy for this irradiated 

sensitized Type 304 SS is approximately 42 kJ/mol (Fig. 34).

Figure 33. Photograph of the fracture

surface of the Type 304 specimen 

sensitized at 600°C for 24 h (85-5A1TT) 

and irradiated to 0.75 dpa.
38

10-11

10-10

0.00172 0.00176 0.0018 0.00184 0.00188

300 290 280 270 260 250

Crack growth rate (m/s)

1/T (1/K)

Temperature (oC)

Q=42 kJ/mol

304 SS, sensitized @ 600oC for 24 h, 0.75 dpa

Tested in NWC

Figure 34. Temperature dependence 

of CGR for the Type 304 specimen 

sensitized at 600°C for 24 h (85-

5A1TT, 0.75 dpa)

3.4 Type 304 SS with and without GBE Treatment

3.4.1 Type 304 SS without GBE Treatment (IT304-01)

A CGR test on a Type 304 SS specimen (IT304-01) irradiated to ≈2 dpa was performed in 

simulated BWR water at ≈290°C. This specimen is the baseline material for a companion specimen given 

the GBE treatment. The test was started in high-purity water with ≈350 ppb DO and a flow rate of 

≈19 mL/min. The specimen was fatigue precracked at R = 0.36, Kmax ≈ 18.0 MPa m1/2, triangular 

waveform, and 2-Hz frequency. After about 240-μm crack advance, R was increased incrementally, and 

the triangular waveform was changed to a slow/fast sawtooth waveform. Environmentally enhanced

cracking was revealed at Kmax ≈ 19 MPa m1/2 with 10-s rise time in test period 2. 

After the cyclic CGRs were obtained, the specimen was subjected to a constant load in test period

6, and the SCC CGR was about 1.16×10-10 m/s at K = 19.8 MPa m1/2. Following a power outage in test 

period 7, another constant load test was performed, and a significantly lower CGR was measured during 

this period. An overload caused by the power outage might have contributed to this low growth rate. The 

loading conditions during the first three or four test periods were repeated in the subsequent test periods

to ensure that the test results were comparable to those before the interruption. Finally, another constantload test was carried out in test period 12 at K = 20.1 MPa m1/2. The SCC CGR was determined to be

5.06×10-11 m/s, about half that observed in test period 6. Table 10 summarizes all test conditions and 

CGR results, and Fig. 35 shows the crack-length-vs.-time curve of this specimen. The CGRs for this 

specimen, under both cyclic and constant-load conditions, are somewhat unusual; the cyclic CGRs are a 

factor of ≈ 5 lower than those typically observed for austenitic SSs irradiated to similar dose levels

(Fig. 36), and the SCC CGRs are either comparable or below the NUREG-0313 curve for nonirradiated 

sensitized SSs (Fig. 37). 
39

Table 10. Test results for the base heat of Type 304 specimen (IT304, ≈2 dpa) in BWR water.

Test

Test 

Time,

Test

Temp.,

ECP,

a

mV (SHE)

O2

Conc.,a RLoad

Rise 

Time,

Return 

Time,

Hold 

Time,b Kmax, ∆K,

CGR in 

Env.,

CGR in 

Air,

Crack 

Length

Period h °C Pt Steel ppb Ratio s s s MPa m1/2 MPa m1/2 m/s m/s mm

6.002

Pre 23 290 - - 350 0.36 0.22 0.22 0.03/0.03 17.8 11.3 5.95E-08 7.81E-08 6.242

1a 69 290 - - 350 0.47 25.7 3.4 4.3/0.6 17.7 9.4 6.44E-12 4.02E-10 6.259

1b 75 290 248 232 350 0.49 8.5 1.7 1.5/0.3 17.7 9.1 6.47E-11 1.10E-09 6.263

1c 78 290 248 234 350 0.46 4.3 0.9 0.7/0.1 18.9 10.2 2.31E-09 3.13E-09 6.281

2 103 290 251 238 350 0.46 26 3.5 4.0/0.5 19.5 10.4 1.01E-09 5.56E-10 6.346

3 148 290 255 - 350 0.45 261 10.4 39/1.6 19.9 11.0 4.79E-10 6.52E-11 6.399

4 212 289 212 202 350 0.68 237 9.5 63/2.5 - - 3.37E-10 - 6.463

5 308 288 215 210 350 0.67 790 9.5 210/2.5 19.7 6.4 2.22E-10 4.46E-12 6.523

6 369 288 224 220 350 1.0 - - - 19.8 0.2 1.16E-10 - 6.573

7 371 288 Power outage - - - - - - - - 6.583

8 489 288 204 208 350 1.0 - - - 19.5 0.2 2.92E-11 - 6.598

9 513 288 202 204 350 0.50 254 10.0 46/2.0 19.5 9.7 4.29E-10 4.66E-11 6.643

10 537 289 200 203 350 0.70 231 9.2 69/2.8 19.6 5.9 7.64E-11 1.20E-11 6.653

11 544 289 200 204 350 0.56 250 10.0 50/2.0 20.0 8.8 4.61E-10 3.62E-11 6.663

12 633 289 200 204 350 1.0 - - - 20.1 0.2 5.06E-11 - 6.682 a Represents values in the effluent. In the feedwater, conductivity was ≈0.08 μS/cm, and DO was ≈800 to 1000 ppb. b Hold periods at minimum load during the loading cycle and at maximum load during the unloading cycle. 

6.00

6.10

6.20

6.30

6.40

6.50

16

17

18

19

20

21

22

23

24

20 40 60 80 100 120 140

Crack Length (mm)

K (MPa m0.5)

Time (h)

Type 304 SS, base heat, (Spec. IT304-CT01)

Test CGRIJR–39

Dose 2 dpa

289°C

High–Purity Water

Kmax

Crack Length

1a

Pre

1b

1c 2

3

(a)

6.30

6.35

6.40

6.45

6.50

6.55

6.60

6.65

6.70

18

20

22

24

26

28

150 200 250 300 350 400

Crack Length (mm)

K (MPa m0.5)

Time (h)

Kmax

Crack Length

Test 

Interrupted

Type 304 SS, base heat, (Spec. IT304-CT01)

Test CGRIJR–39

Dose 2 dpa

4

5

6

7

8

289°C

High–Purity Water

(b)

Figure 35. Crack-length-vs.-time plot for the base heat of Type 304 SS specimen (IT304-01, ~2 dpa) 

in BWR water at 289°C. 
40

6.45

6.50

6.55

6.60

6.65

6.70

6.75

6.80

18

20

22

24

26

28

350 400 450 500 550 600 650

Crack Length (mm)

K (MPa m0.5)

Time (h)

289°C

High–Purity Water

Kmax

Crack Length

Type 304 SS, base heat, (Spec. IT304-CT01)

Test CGRIJR–39

Dose 2 dpa

8

9

10

11

12

(c)

Figure 35. (Contd.)

10-12

10-11

10-10

10-9

10-8

10-7

10-12 10-11 10-10 10-9 10-8 10-7

CGRenv (m/s)

CGRair (m/s)

CGRI JR-39, IT304-01

304 SS, base heat 

~2 dpa

290oC, 350 ppb

Figure 36. Cyclic CGR for the base heat of 

Type 304 SS specimen (IT304-01) 

irradiated to 2 dpa.
41

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30 35 40

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

CGRI JR-39, IT304-01

304 SS, base heat

2 dpa

NWC, 290oC

Figure 37. Crack growth rate versus stress 

intensity under constant-loading conditions 

for the base heat of Type 304 SS specimen

(IT304-01) irradiated to 2 dpa.

After the CGR test, a J-R curve test was performed on the specimen at 289°C in high-DO water 

(≈350 ppb DO). The test was conducted at a constant extension rate of 0.43 μm/s. The test was 

interrupted periodically to measure crack length by DCPD measurements. The measured load vs. loadline displacement and the J-R curves are shown in Figs. 38 and 39, respectively. The JIC determined for

this non-GBE Type 304 SS is 83 kJ/m2. 

0.0

1.0

2.0

3.0

4.0

0 0.5 1 1.5 2 2.5

Loading Pin Displacement

Measured Displacement

Load (kN)

Displacement (mm)

Type 304 (Spec. IT304-CT01)

Test CGRIJR–39

Fluence 1.31 x 1021 n/cm2

289°C High-Purity Water

Figure 38. Load vs. load-line displacement 

curve for the base heat of Type 304 SS 

specimen (IT304-01) tested in NWC at 

289°C

After the J-R test, the final crack size was marked by fatigue cycling at room temperature. The 

specimen was then fractured, and the final crack length of both halves of the fractured specimen was 

measured from the photograph of the fracture surface (Fig. 40). The actual crack extension was about 

10% greater than the value determined from the DCPD measurements. Crack extensions, estimated from 

the DCPD method, were adjusted accordingly. 
42

0

50

100

150

200

250

300

0 0.5 1 1.5 2 2.5 3

J (kJ/m2)

Crack Extension (mm)

Type 304 (Spec. IT304-CT01)

Test CGRIJR–39

Fluence 1.31 x 1021 n/cm2

J = 132∆a0.3

JIC = 83 kJ/m2

Figure 39. Fracture toughness J-R curves 

for the base heat of Type 304 SS specimen 

(IT304-01) tested in NWC at 289°C.

Figure 40. Photograph of the fracture surface of the 

base heat of Type 304 SS (IT304-01) irradiated to 

2 dpa.

3.4.2 Type 304 SS with GBE Treatment (GB304-01)

A CGR test on the GBE-treated Type 304 SS specimen, GB304-01, was carried out in BWR water 

at 289°C. The specimen was irradiated to ≈2 dpa at 290°C in the Halden reactor. The test was started in 

high-purity water with ≈400 ppb DO and a flow rate of ≈19 mL/min. Fatigue precracking was performed

with a triangular waveform at R = 0.37, Kmax = 18.8 MPa m1/2, and 2-Hz frequency. After ≈100-μm 

crack advance, the load ratio was increased to ≈ 0.56, and the waveform was changed to a slow/fast 

sawtooth with a rise time of 60 s and a return time of 12 s. After cyclic CGRs were obtained, the 

specimen was tested under constant load at 17.8, 13.2, and 11.0 MPa m1/2. At 96 h, the DO level was 

decreased from ≈400 ppb to <30 ppb by bubbling the feedwater with a gas mixture of N2 and 4% H2. 

Accordingly, the ECP values for both Pt and SS electrodes were reduced significantly, and the constantload CGRs decreased by a factor of 10. After ≈50 h, crack growth became negligible at ≈11 MPa m1/2. 

To ensure that the DCPD measurements were not influenced by any unbroken ligaments, the specimen 

was subjected to cyclic loading at R = 0.54, including a sawtooth waveform with ≈ 300-s rise time for a 

relatively short time in test period 8. The cyclic CGR in this test period was determined to be ≈3 x 10-9 

m/s. Finally, the specimen was subjected to a constant load, and the CGR was measured at K = 13.9 MPa 

m1/2 in HWC. The CGR test was terminated after 200 h. Table 11 presents all the test conditions and 

CGR results on this GBE-treated Type 304 specimen, and Fig. 41 shows the crack-length-vs.-time plot. 
43

Cyclic CGR data are shown in Fig. 42, and constant-load CGRs are given in Fig. 43. Environmental 

enhancement is evident in this material, and the constant-load CGRs are significantly higher than the 

NUREG-0313 curve for nonirradiated sensitized SSs.

Table 11. Test results for the GBE-treated Type 304 specimen (GB304-01, ≈2 dpa) in BWR water.

Test

Test 

Time,

Test

Temp.,

ECP,

a

mV (SHE)

O2

Conc.,a RLoad

Rise 

Time,

Return 

Time,

Hold 

Time,b Kmax, ∆K,

CGR in 

Env.,

CGR in 

Air,

Crack 

Length,

Period h °C Pt Steel ppb Ratio s s s MPa m1/2 MPa m1/2 m/s m/s mm

6.043

Pre 25 288 - - 400 0.37 0.22 0.22 0.03/0.03 18.8 11.9 9.68E-08 9.11E-08 6.331

1 29 289 - - 400 0.53 4.9 0.8 1.1/0.2 18.3 8.6 9.53E-09 1.69E-09 6.413

2 45 289 251 238 400 0.53 49.2 9.8 10.8/2.2 18.6 8.8 2.96E-09 1.76E-10 6.532

3 56 288 255 238 400 0.58 47.6 9.5 12.4/2.5 17.1 7.2 2.85E-09 1.01E-10 6.609

4 69 288 234 228 400 1.0 - - - 17.8 0.3 2.72E-09 - 6.732

5 77 288 215 210 400 1.0 - - - 13.2 0.3 1.40E-09 - 6.786

6 96 289 196 154 400 1.0 - - - 10.9 0.2 1.16E-09 - 6.863

7 166 288 -524 -265 <30 1.0 - - - 11.0 0.2 negligible - 6.908

8 170 289 -527 -273 <30 0.54 217 8.7 83/3.3 12.2 5.7 3.02E-09 9.45E-12 6.948

9 272 288 -465 -374 <30 1.0 - - - 13.9 0.3 2.04E-10 - 7.016 a Represents values in the effluent. In the feedwater, conductivity was ≈0.1 μS/cm, and DO was ≈800 to 1000 ppb. b Hold periods at minimum load during the loading cycle and at maximum load during the unloading cycle. 

6.00

6.20

6.40

6.60

6.80

7.00

10

15

20

25

30

35

20 30 40 50 60 70 80 90 100

Crack Length (mm)

K (MPa m0.5)

Time (h)

GBE 304 (Spec. GB304-CT01)

Test CGRI JR–40

Dose 2 dpa

289°C

High–Purity Water

Kmax

Crack Length

Pre

1

2

3

4

5

6

(a)

6.70

6.75

6.80

6.85

6.90

6.95

7.00

7.05

8

12

16

20

24

28

32

36

100 150 200 250 300

Crack Length (mm)

K (MPa m0.5)

Time (h)

289°C

High–Purity Water

K

Crack Length

400 ppb DO <30 ppb DO

GBE 304 (Spec. GB304-CT01)

Test CGRI JR–40

Dose 2 dpa

6

7

8

9

(b)

Figure 41. Crack-length-vs.-time plot for the GBE-treated Type 304 SS specimen (GB304-01, 

≈2 dpa) in BWR water at 289°C. 
44

10-12

10-11

10-10

10-9

10-8

10-7

10-12 10-11 10-10 10-9 10-8 10-7

CGRenv (m/s)

CGRair (m/s)

CGRI JR-39, GB304-01

GBE treated Type 304 SS, ~2 dpa

289°C, High-Purity water

Figure 42. Cyclic crack growth rate for the 

GBE-treated Type 304 SS specimen (GB304-

01) irradiated to 2 dpa.

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30 35 40

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

CGRI JR-40, GB304-01

GBE 304, ~2 dpa

289oC

Open Symbols: NWC BWR Env.

Closed Symbols: HWC BWR Env.

Figure 43. Crack growth rate versus stress 

intensity under constant-loading conditions for 

the GBE-treated Type 304 SS specimen 

(GB304-01) irradiated to 2 dpa.

The DO level in the system was increased to ≈400 ppb after the CGR test. A J-R curve test was

then performed on the specimen at 289°C in the high-purity water. The test was conducted at a constant 

extension rate of ≈ 0.43 μm/s, and the test was interrupted periodically to measure crack length by the 

DCPD method. The load vs. load-line displacement curve is shown in Fig. 44, and the corresponding J-R 

curve is given in Fig. 45. The final crack size was marked by fatigue cycling at room temperature. The 

specimen was then fractured, and the final crack length of both halves of the fractured specimen was 

measured from the photograph of the fracture surface (Fig. 46). The actual crack extension was ≈ 27% 

greater than the value determined from the DCPD measurement. Crack extensions, estimated from the 

DC potential method, were adjusted accordingly. 
45

0.0

1.0

2.0

3.0

4.0

5.0

0 0.5 1 1.5 2 2.5 3 3.5

Loading Pin Displacement

Measured Displacement

Load (kN)

Displacement (mm)

CGRI JR-40, GB304-01

GBE 304 SS, 2 dpa

289oC, High-Purity Water

Figure 44. Load vs. load-line displacement 

curve for the GBE-treated Type 304 SS 

(GB304-01, ~2 dpa) tested in NWC at 

289°C.

0

100

200

300

400

500

600

700

0 0.5 1 1.5 2 2.5 3

J (kJ/m2)

Crack Extension (mm)

J = 392∆a0.74

JIC = 138 kJ/m2

Test CGRI JR-40, GB304-01

GBE 304 SS, 2 dpa

289oC, High-Purity Water

Figure 45. Fracture toughness J-R curves 

for the GBE-treated Type 304 SS (GB304-

01, ~2 dpa) tested in NWC at 289°C.

Figure 46. Photograph of the fracture surface of the 

GBE-treated Type 304 SS (GB304-01) irradiated to 2

dpa.
46

3.5 Type 316 SS with and without GBE Treatment

3.5.1 Type 316 without GBE Treatment (IT316-01)

A CGR test was performed in simulated BWR water at 289°C on a Type 316 SS specimen 

(IT316-01) that was irradiated to ≈2 dpa. The specimen is the baseline material for a companion 

specimen that has been given the GBE treatment. The test was started in high-purity water with ≈700 ppb 

DO and a flow rate of ≈ 20 mL/min. Fatigue precracking was conducted with a triangular waveform at

R ≈ 0.35, Kmax ≈ 18 MPa m1/2, and frequency of 2 Hz. After about 150-μm crack advance, R was 

increased to 0.5, and the waveform was changed to a slow/fast sawtooth with rise times of 1 to 60 s and

return times of 1 to 12 s. The first attempt to introduce environmentally enhanced cracking was 

unsuccessful, and the cyclic CGRs in the test environment were generally comparable to those in air under 

the same loading conditions. The Kmax was then increased to ≈ 20 MPa m1/2, and the initial loading 

sequence was repeated. Environmental enhancement was achieved during test period 8. Subsequently, 

the specimen was subjected to constant load. Table 12 summarizes all the test conditions and CGR 

results, and Fig. 47 shows the crack-length-vs.-time plot of this specimen. 

Cyclic CGR data are shown in Fig. 48, and constant-load CGRs are given in Fig. 49. Little 

environmental effect can be seen for the cyclic CGRs. The CGRs in test period 2, 3, and 5a are not

included in the figure due to insignificant crack advance. Under the constant-load conditions, the 

measured CGRs are very close to the NUREG-0313 curve, as shown in Fig. 49. 

Table 12. Test results for the base heat of Type 316 specimen (IT316-01, ≈2 dpa) in BWR water.

Test

Test 

Time,

Test

Temp.,

ECP,

a

mV (SHE)

O2

Conc.,a RLoad

Rise 

Time,

Return 

Time,

Hold 

Time,b Kmax, ∆K,

CGR in 

Env.,

CGR in 

Air,

Crack 

Length,

Period h °C Pt Steel ppb Ratio s s s MPa m1/2 MPa m1/2 m/s m/s mm

6.022

Pre 21 288 266 208 600 0.34 0.22 0.22 0.03/0.03 18.3 12.1 5.37E-08 9.33E-08 6.185

1 26 288 266 208 600 0.37 0.87 0.87 0.13/0.13 17.7 11.1 2.69E-08 1.87E-08 6.369

2 66 288 271 209 600 0.53 49.4 9.90 10.6/2.1 16.8 8.0 3.85E-11 1.27E-10 6.382

3 70 288 271 208 600 0.48 25.2 3.40 4.8/0.6 17.4 9.0 negligible 3.59E-10 6.383

4 163 288 274 208 600 0.51 25.5 3.40 4.5/0.6 19.9 9.7 3.54E-10 4.68E-10 6.439

5a 165 288 274 208 600 0.37 0.88 0.88 0.12/0.12 19.4 12.2 negligible 2.49E-08 6.439

5b 167 288 273 208 600 0.33 0.22 0.22 0.03/0.03 21.2 14.2 9.00E-08 1.57E-07 6.668

6 187 288 273 208 600 0.48 25.5 3.40 4.5/0.6 20.4 10.6 2.69E-10 6.10E-10 6.707

7 210 289 274 208 600 0.45 25.7 3.40 4.3/0.6 20.5 11.3 2.00E-10 7.13E-10 6.720

8 234 289 275 208 600 0.43 10.3 1.70 1.7/0.3 20.9 11.8 3.16E-09 2.05E-09 6.820

9 332 289 275 209 600 0.40 51.7 3.40 8.3/0.6 20.5 12.2 2.59E-10 4.41E-10 6.881

10 356 289 275 209 600 0.43 103 3.40 17.2/0.6 20.9 11.9 2.41E-10 2.09E-10 6.903

11 380 289 274 208 600 0.44 257 10.3 43.3/1.7 21.3 11.8 2.76E-10 8.33E-11 6.921

12 405 289 274 209 600 0.50 253 10.1 47.4/1.9 21.4 10.7 2.87E-10 6.46E-11 6.943

13 478 289 273 209 600 1.0 - - - 20.6 0.2 1.49E-10 - 6.990

14 547 289 273 210 600 1.0 - - - 16.5 0.2 6.70E-11 - 6.997 a Represents values in the effluent. In the feedwater, conductivity was ≈0.1 μS/cm, and DO was ≈800 to 1000 ppb. b Hold periods at minimum load during the loading cycle and at maximum load during the unloading cycle. 
47

6.00

6.10

6.20

6.30

6.40

6.50

6.60

6.70

10

15

20

25

30

35

40 80 120 160 200

Crack Length (mm)

K (MPa m0.5)

Time (h)

Type 316 SS, base heat, (Spec. IT316-CT01)

Test CGRIJR–41

Dose 2 dpa

289°C

High–Purity Water

Kmax

4

2

Crack Length

5b

6

Pre

1

3

5a

(a)

6.60

6.70

6.80

6.90

7.00

10

15

20

25

30

35

200 250 300 350

Crack Length (mm)

K (MPa m0.5)

Time (h)

289°C

High–Purity Water

Kmax

9

7 Crack Length

10

Type 316 SS, base heat, (Spec. IT316-CT01)

Test CGRIJR–41

Dose 2 dpa

8

11

(b)

6.70

6.80

6.90

7.00

7.10

7.20

10

15

20

25

30

35

40

400 440 480 520

Crack Length (mm)

K (MPa m0.5)

Time (h)

289°C

High–Purity Water

K

12

Crack Length

14

Type 316 SS, base heat, (Spec. IT316-CT01)

Test CGRIJR–41

Dose 2 dpa

13

(c)

Figure 47. Crack-length-vs.-time plot for the base heat of Type 316 SS specimen (IT316-01, ≈2 dpa) 

in BWR water at 289°C. 
48

10-12

10-11

10-10

10-9

10-8

10-7

10-12 10-11 10-10 10-9 10-8 10-7

CGRenv (m/s)

CGRair (m/s)

CGRI JR-41, IT316-01

316 SS, base hear

~ 2 dpa

289oC, 600 ppb DO

Figure 48. Cyclic CGR for the base heat of Type 

316 SS (IT316-01) irradiated to 2 dpa.

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30 35 40

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

CGRI JR-41, IT316-01

316 SS, base heat, ~2 dpa

289oC

Figure 49. Crack growth rate versus stress 

intensity under constant-loading conditions for the 

base heat of Type 316 SS (IT316-01) irradiated to 

2 dpa.

After the CGR test, a J-R curve test was performed on the specimen at 289°C in high-DO water 

(≈600 ppb DO). The test was conducted at a constant extension rate of 0.43 μm/s. The test was 

interrupted periodically to measure crack length by DCPD measurements. The load vs. load-line 

displacement curve is shown in Fig. 50, and the corresponding J-R curve is shown in Fig. 51. The J1C

determined for this specimen is 89 kJ/m2.
49

0.0

1.0

2.0

3.0

4.0

0 0.5 1 1.5 2 2.5

Loading Pin Displacement

Measured Displacement

Load (kN)

Displacement (mm)

CGRI JR-41, IT316-01

316 SS base heat, 2 dpa

289oC, High-Purity water

Figure 50. Load vs. load-line 

displacement curve for the base heat 

of Type 316 SS (IT316-01) tested in 

NWC at 289°C

0

50

100

150

200

250

300

0 0.5 1 1.5 2 2.5 3

J (kJ/m2)

Crack Extension (mm)

Test CGRI JR-41, IT316-01

316 SS base heat, 2 dpa

289oC, High-Purity water

J = 163∆a0.42

JIC = 89 kJ/m2

Figure 51. Fracture toughness J-R 

curve for the base heat of Type 316 SS 

(IT316-01) tested in NWC at 289°C.

After the J-R test, the final crack length was marked by fatigue cycling at room temperature. The 

specimen was then fractured, and the final crack length of both halves of the fractured specimen was 

measured from the photograph of the fracture surface (Fig. 52). The actual crack extension was ≈6% 

greater than the value determined from the DC potential measurements. Crack extensions, estimated 

from the DC potential method, were adjusted accordingly. 

Figure 52. Photograph of the fracture surface of the 

base heat of Type 316 SS (IT316-01) irradiated to 2

dpa.
50

3.5.2 Type 316 SS with GBE Treatment (GB316-01)

A CGR test was initiated in BWR water at 289°C on a GBE specimen of Type 316 SS (GB316-01) 

that had been irradiated to ≈2 dpa at 290°C in the Halden reactor. The test was started in high-purity 

water with effluent DO content of ≈700 ppb and a flow rate of ≈ 20 mL/min. The specimen was fatigue 

precracked at R = 0.25, Kmax = 16-17 MPa m1/2, a triangular waveform, and 0.2-1 Hz frequency. After 

≈340-μm crack advance, R was increased to ≈ 0.4, Kmax decreased to ≈15 MPa m1/2, and the waveform 

was changed to a slow/fast sawtooth with a rise time of 30 s and a return time of 4 s. Environmental 

enhancement was not revealed with the change in loading condition; instead, the crack growth stopped. 

To reestablish crack growth, the R was decreased, Kmax was increased to ≈ 17 MPa m1/2, and frequency 

increased to 1 Hz. The CGRs were generally lower than those in air under the same loading conditions. 

Once again, an attempt to increase R and rise time resulted in insignificant crack growth, and little or no 

crack growth was noted at load ratios of 0.40-0.55 and ∆K values less than 10 MPa m1/2. The Kmax was

increased to ≈19.5 MPa m1/2, and the initial loading sequence was repeated for the third time. Finally, 

environmental enhancement was observed after 340 h during test period 10. After the cyclic CGR test, 

the specimen was tested under constant loads at 18.9 and 16.5 MPa m1/2 in test periods 12 and 13,

respectively. 

Table 13 summarizes the test conditions and CGR results, and Fig. 53 shows the crack-length-vs.-

time plots of this GBE-treated Type 316 SS specimen. The cyclic CGRs are plotted in Fig. 54, and 

environmental enhancement can be seen at around 10-9 m/s. The CGRs for test periods 1, 2, and 6 are not 

included in Fig. 54 due to insignificant crack growth. Figure 55 shows the constant-load CGRs obtained 

in this specimen in high-DO (≈600 ppb) water . The result from test period 8 is ignored due to 

insignificant crack extension. The CGRs measured at two constant load conditions are much higher than 

the NUREG-0313 curve, and this finding suggests a strong irradiation effect on the SCC cracking in this 

material.

Table 13. Test results for the GBE-treated Type 316 specimen (GB316-01, ≈2 dpa) in BWR water.

Test

Test 

Time,

Test

Temp.,

ECP,

a

mV (SHE)

O2

Conc.,a RLoad

Rise 

Time,

Return 

Time,

Hold 

Time,b Kmax, ∆K,

CGR in 

Env.,

CGR in 

Air,

Crack 

Length,

Period h °C Pt Steel ppb Ratio s s s MPa m1/2 MPa m1/2 m/s m/s mm

6.033

Pre a 4 288 270 203 700 0.25 0.44 0.44 0.06/0.06 16.1 12.1 1.91E-08 4.22E-08 6.091

Pre b 22 288 271 207 700 0.25 0.44 0.44 0.06/0.06 17.1 12.8 1.13E-07 5.15E-08 6.316

Pre c 24 288 275 198 700 0.26 2.2 2.2 0.3/0.3 16.1 11.9 2.85E-08 8.07E-09 6.384

1 42 288 270 205 700 0.42 24.6 3.3 5.4/0.7 14.7 8.5 negligible 2.89E-10 6.393

2a 52 288 270 206 700 0.31 4.3 0.8 0.7/0.2 15.7 10.9 negligible 3.28E-09 6.396

2b 68 288 271 208 700 0.31 4.3 0.8 0.7/0.2 15.6 10.7 negligible 3.14E-09 6.402

2c 77 288 272 209 700 0.33 0.85 0.85 0.15/0.15 15.9 10.7 negligible 1.61E-08 6.408

3 95 288 272 210 700 0.26 0.44 0.44 0.06/0.06 18.4 13.6 4.19E-08 6.31E-08 6.562

4 100 288 272 210 700 0.28 2.2 2.2 0.3/0.3 18.8 13.5 1.79E-08 1.26E-08 6.690

5 114 289 271 210 700 0.39 8.5 1.7 1.5/0.3 20.2 12.2 1.09E-08 2.68E-09 7.047

6a 139 290 270 210 700 0.49 39.8 3.2 10.2/0.8 17.3 8.8 negligible 2.13E-10 7.052

6b 161 290 700 0.49 40.6 3.2 9.4/0.8 19.2 9.7 6.89E-11 2.91E-10 7.065

7 174 287 273 213 700 0.41 4.2 1.7 0.8/0.3 19.7 11.5 5.94E-10 4.54E-09 7.090

8 217 288 280 220 700 1.0 - - - 20.0 - 3.56E-10 - 7.096

9 330 288 257 199 600 0.43 8.30 8.3 1.7/1.7 20.1 11.5 1.27E-09 2.32E-09 7.223

10 377 288 256 196 600 0.43 81.2 9.7 18.8/2.3 19.0 10.8 1.52E-09 1.92E-10 7.364

11 387 288 255 195 600 0.48 398 9.5 102/2.5 18.5 9.7 6.08E-10 2.90E-11 7.379

12 404 288 255 195 600 0.98 - - - 18.9 0.4 9.04E-10 - 7.437

13 426 288 255 195 600 0.98 - - - 16.5 0.3 7.57E-10 - 7.499 a Represents values in the effluent. In the feedwater, conductivity was ≈0.1 μS/cm, and DO was ≈800 to 1000 ppb.

b Hold periods at minimum load during the loading cycle and at maximum load during the unloading cycle
51

6.20

6.40

6.60

6.80

7.00

7.20

10

15

20

25

30

35

20 40 60 80 100 120 140

Crack Length (mm)

K (MPa m0.5)

Time (h)

GBE 316 SS, (Spec. GB316-01)

Test CGRIJR–42

Dose 2 dpa

289°C

High–Purity Water

Kmax

Crack Length

Pre b

Pre c

1 2a

2b

2c

3

4

5

6a

(a)

6.70

6.80

6.90

7.00

7.10

7.20

7.30

7.40

7.50

12

16

20

24

28

32

36

160 200 240 280 320

Crack Length (mm)

K (MPa m0.5)

Time (h)

289°C

High–Purity Water

Kmax

6b

GBE 316 SS, (Spec. GB316-01)

Test CGRIJR–42

Dose 2 dpa

7

8

9

Crack Length

(b)

7.10

7.20

7.30

7.40

7.50

12

16

20

24

28

32

36

280 300 320 340 360 380 400 420 440

Crack Length (mm)

K (MPa m0.5)

Time (h)

289°C

High–Purity Water

Kmax Crack Length

11

GBE 316 SS, (Spec. GB316-01)

Test CGRIJR–42

Dose 2 dpa

9

10

12

13

(c)

Figure 53. Crack-length-vs.-time plot for the GBE-treated Type 316 SS (GB316-01, ≈2 dpa) tested in 

BWR water at 289°C.
52

10-12

10-11

10-10

10-9

10-8

10-7

10-12 10-11 10-10 10-9 10-8 10-7

CGRenv (m/s)

CGRair (m/s)

CGRI JR-42, GB316-01

GBE 316 SS

~ 2 dpa

289°C, 700 ppb DO

Figure 54. Cyclic CGR for the GBE-treated Type 

316 SS (GB316-01) irradiated to 2 dpa.

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30 35 40

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

CGRI JR-42, GB316-01

GBE 316 SS, ~2 dpa

289oC

Figure 55. Crack growth rate versus stress 

intensity under constant-loading conditions for the 

GBE-treated Type 316 SS (GB316-01) irradiated 

to 2 dpa.

After the CGR test, a J-R curve test was performed on the GBE 316 SS specimen at 289°C in highDO water (≈600 ppb DO). The test was conducted at a constant extension rate of 0.43 μm/s. The test 

was interrupted periodically to measure crack length by DCPD measurements. The obtained load vs. 

load-line displacement curve is shown in Fig. 56, and the corresponding J-R curve is shown in Fig. 57. A 

relatively high fracture toughness was obtained for this specimen with a JIC of about 178 kJ/m2. 
53

0.0

1.0

2.0

3.0

4.0

0 0.5 1 1.5 2 2.5

Loading Pin Displacement

Measured Displacement

Load (kN)

Displacement (mm)

CGRI JR–42, GB316-01

GBE 316 SS, 2 dpa 

289oC, High-Puity water

Figure 56. Load vs. load-line displacement 

curve for the GBE-treated Type 316 SS 

(GB316-01, ≈2 dpa ) tested in NWC at 

289°C.

J = 290∆a0.36

0

100

200

300

400

500

600

0 0.5 1 1.5 2 2.5 3

J (kJ/m2)

Crack Extension (mm)

Test CGRI JR-42, GB316-01

GBE 316 SS, 2 dpa

289oC, High-Purity water

JIC = 178 kJ/m2

Figure 57. Fracture toughness J-R curves 

for the GBE-treated Type 316 SS (GB316-

01, ≈2 dpa) tested in NWC at 289°C.

Following the J-R test, the final crack length was marked by fatigue cycling at room temperature in 

air. The specimen was then fractured, and the final crack length of both halves of the fractured specimen 

was measured from the photograph of the fracture surface (Fig. 58). The actual crack extension was ≈4% 

greater than the value determined from the DCPD measurements. Crack extensions, estimated from the 

DCPD method, were adjusted accordingly.

Figure 58. Photograph of the fracture surface of the 

GBE-treated Type 316 SS (GB316-01) irradiated to 

2 dpa.
54

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55

4 Discussion

In addition to the data presented in Section 3, the test results from previous Halden-II tests 20 are 

considered in this section to derive an overall view of IASCC susceptibility. The effects of neutron dose, 

water chemistry, alloy type, and welding and processing conditions on IASCC behavior are discussed. 

The influence of neutron irradiation on the fracture toughness of austenitic SSs, weld HAZs, and a cast SS 

material is also presented. 

4.1 Specimen Size Criteria Relevant to Cyclic and Constant-Load CGR Tests

Before the test results can be used to evaluate the IASCC behavior, loading conditions in the crack 

growth rate test need to be validated. Both cyclic and constant-load CGR tests are intended to be 

performed under a K-dominated condition in which the stress intensity factor (K) defines the amplitude of 

the crack-tip singularity.

53-55 If K characterizes the crack-tip loading conditions (stress and strain) 

adequately, the test results will be independent of sample size or geometry, even for materials that exhibit 

plastic deformation. Stress-intensity-factor dominance implies that the plastic zone ahead of the crack 

front must be small compared to all length dimensions in a test specimen. For a given specimen 

geometry, the applied load must be less than a limiting value to ensure that a relatively small plastic zone 

ahead of the crack tip can be embedded within a largely elastic body of the test specimen. 

It is well documented that the plastic zone size is proportional to the square of the ratio between the 

stress intensity factor and the yield strength (K/σy)2.

56, 57 Using an elastic-perfect-plastic model, Irwin 

estimated that the size of plastic zone as 58

2

p

y

1 K

r

  =  

π σ   

(19)

Comparing Eq. (3) and Eq. (19) indicates that the required specimen dimensions (thickness and remaining 

ligament) are at least a factor of ≈8 greater than the plastic zone size. For a given specimen, the upper 

limit of K to ensure a K-dominated test condition is

1/2 K 0.63 [B, or (W a)] limit y ≤ ⋅σ ⋅ − (20)

For strain-hardening materials, such as austenitic SSs, more stress can be carried by the plastically 

deformed region than that in an elastic-perfect-plastic material. Thus, the actual plastic zone size is 

smaller than that given by Eq. (19), if σy is interpreted as the initial yield stress. For this reason, the flow 

stress, which is defined as the average of yield strength and ultimate tensile stress, can be used in Eq. (20)

instead of σy without compromising the K-dominated test condition.

The picture is more complex in irradiated materials. While neutron irradiation increases yield 

strength significantly, it affects the strain hardening behavior as well. For specimens irradiated to lowdoses, strain hardening remains to some extent. This finding suggests that higher far-field applied

stresses should be allowed for testing of low-dose irradiated specimens compared with nonirradiated 

specimens. In contrast, strain softening often occurs in high-dose specimens due to dislocation 

channeling. Because the plastic zone size can be affected by strain softening, the allowable K for CGR 

tests could be lower for strain-softened materials. 
56

Several authors 59,60 have contended that the K/size criterion specified in Eq. (3) or Eq. (20), where 

σy is taken as the yield strength of the irradiated material, is inadequate due to irradiation-induced strain

softening. Numerous SCC experiments have indeed shown a K-size effect on CGRs when strain 

hardening is absent.61, 62 To account for the strain softening, an effective yield strength (σeff) has been 

proposed to replace σy in Eq. (3). The use of the average or one-third the irradiated and nonirradiated 

yield strengths seems to work well for some experimental data. 59,60 However, the use of the yield 

strength of the nonirradiated material in determining the allowable plastic zone size in an irradiated 

specimen is questionable from a purely mechanics perspective. 

A simple model, shown in Fig. 59, is proposed in this study to account for the strain-softening 

effect in irradiated materials. If an irradiated material that exhibits work-softening behavior has a

monotonic yield strength σy-irr and a reduced stress σs-irr beyond yield, then, according to the elastic

theory, the stress released in the yield region ahead of the crack tip is

y

K

2 r

σ =

π ⋅

(21)

where ry is the size of the yielding zone. The released stress must be redistributed in a large area that is 

defined by rp. Thus,

y y irr s irr r

p 0

K

r dr

2 2 r

σ +σ − − ⋅= ⋅ ∫

π ⋅

(22)

Integrating Eq. (22) and solving for rp give

2

p

y irr y irr s irr

2 K

r

( ) −− −

= ⋅ π σ σ +σ

(23)

Crack

σy-irr

σs-irr

ry rp

σ

X

K

2 r π ⋅

Crack

σy-irr

σs-irr

ry rp

σ

X

K

2 r π ⋅

Figure 59. Stress distribution ahead of a two-dimensional crack in a work-softened material.

The reduced far-field stress beyond the yield point σs-irr should be distinguished from the final 

fracture stress (σf ) beyond necking. Without the normal plastic instability (necking) exhibited in a 

monotonic tensile test, the far-field stress due to strain-softening (σy-irr - σs-irr) is not greatly reduced in 

an irradiated material. Assuming a 20% reduction in this far-field stress, that is σs-irr = 0.8σy-irr, the 
57

plastic zone size is only ≈11% larger than that predicted by Eq. (3) using σy-irr. In most cases, the strain 

softening exhibited in irradiated materials is not more than a few percent of their yield strength (σy-irr). 

Thus, the plastic zone size with and without consideration of strain softening should be similar. Several 

experimental and finite element method (FEM) studies have confirmed that the plastic zone size is not 

strongly affected by strain-softening behavior in irradiated materials.63,64 This observation agrees with 

the simple analysis presented above. Whether or not a K-dominated condition is satisfied in a CGR test

should be determined by Eq. (3) or Eq. (20) using the irradiated yield strength (σy-irr). As pointed out 

previously, this criterion is a very strict requirement to ensure a small-scale yielding condition ahead of a 

crack.47 The specimen-size effects observed in SCC experiments could be attributed to non-mechanical 

factors (e.g., crack tip chemistry and deformation behavior).

4.2 Effects of Weld Process on Cracking Behavior

4.2.1 Cyclic CGR for Nonirradiated Weld HAZs

The present study tested HAZ specimens made from Type 304L SA weld as well as Type 304 SMA 

weld (Table 4). The cracking behavior of nonirradiated specimens of these materials needs to be 

evaluated to establish a baseline for studying IASCC susceptibility. For CGR tests performed on 1/4TCT specimens, the K-dominated test condition is likely to be violated in nonirradiated tests because of the 

low yield strength exhibited in nonirradiated materials. In fact, none of the cyclic CGR tests performed 

on the nonirradiated weld HAZ specimens met the K/size criterion specified by Eq. (3). However, 

loading conditions for most of these data points did satisfy the criterion for fatigue crack propagation tests 

that was detailed in ASTM E-647. A predominantly elastic behavior of specimens in fatigue tests was 

ensured by:

 max 2

y

4 K Wa ( ) − ≥

π σ

(24) 

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-13 10-12 10-11 10-10 10-9 10-8 10-7 10-6

304L SA HAZ, as-welded

304L SA HAZ, 24 h@ 500 oC

304 SMA HAZ, as-welded

304 SMA HAZ, 24 h @ 500 oC

CGRenv (m/s)

CGRair (m/s)

Austenitic SS Weld HAZ

nonIrradiated

289°C, 300–500 ppb DO

Shack/Kassner model 

for 8 ppm DO

Figure 60. Cyclic CGRs for nonirradiated SA and SMA weld HAZ 

specimens under the as-welded and 

thermally treated conditions.19
58

Figure 60 shows cyclic data from nonirradiated SA and SMA weld HAZ specimens. 

Environmental effects can be seen for all specimens. The curve for environmentally enhanced cracking 

represented by the Shack and Kassner model48 for the case of 8 ppm DO water is also shown in the 

figure. A slightly higher environmental enhancement can be seen in the HAZ specimens with post-weld 

heat treatment. 

4.2.2 Cyclic CGR for Irradiated Weld HAZ Materials

Cyclic CGRs for irradiated SA and SMA HAZ specimens were measured at two dose levels, 

≈0.75 dpa and ≈2 dpa. 19,20 Thanks to irradiation hardening, all CGR tests performed at these dose levels 

can meet even the K-size criterion specified by Eq. (3). Figure 61a shows cyclic CGR results for aswelded SA and SMA HAZ specimens at 0.75 dpa, a proposed threshold dose level for IASCC in BWR 

internal components. Environmentally enhanced cracking was achieved easily at this dose level for both 

types of weld HAZ materials under the as-welded condition. The susceptibility to SCC was elevated 

considerably by irradiation. No difference in cyclic CGRs can be seen between SA and SMA weld HAZ 

materials. The rate of cracking generally settled at around 5 to 7×10-10 m/s in simulated BWR NWC

water at ≈290°C. At ≈2 dpa, the cyclic CGRs were also similar for both weld HAZ materials (as shown 

in Fig. 61b). The effect of environmental enhancement was slightly greater at the higher dose, and the 

observed cyclic CGRs were close to 10-9 m/s.

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-13 10-12 10-11 10-10 10-9 10-8 10-7 10-6

304L SA HAZ (Spec. GG5TA)

304L SA HAZ (Spec. GG5TB)

304 SMA HAZ (Spec. 85-7A)

304 SMA HAZ (Spec. 85-7B)

CGRenv (m/s)

CGRair (m/s)

As-welded SA and SMA HAZ

0.75 dpa

289°C, 300–600 ppb DO

Open symbols: 304L SA weld

Close symbols: 304 SMA weld

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-13 10-12 10-11 10-10 10-9 10-8 10-7 10-6

304L SA HAZ (Spec. GG6TA)

304 SMA HAZ (Spec. 85-XA)

CGRenv (m/s)

CGRair (m/s)

As-welded SA and SMA HAZ

~2 dpa

289°C, 300–600 ppb DO

Open symbols: 304L SA weld

Close symbols: 304 SMA weld

(a) (b)

Figure 61. Cyclic CGRs for SA and SMA weld HAZ specimens under as-welded condition irradiated to 

(a) 0.75 dpa, and (b) ≈2 dpa. 

To simulate low-temperature sensitization, both SA and SMA welds were thermally treated at 

500°C for 24 h prior to irradiation. Figure 62 shows the cyclic CGRs for weld HAZ specimens under the 

as-welded and thermally treated conditions. Comparisons are made at 0.75 dpa for 304 SMA weld and at 

≈2 dpa for 304L SA weld. No obvious difference was detected between as-welded and thermally treated 

CGR results for both types of welds. This observation suggests that, at the present dose level, the low-
59

temperature sensitization does not have a significant effect on IASCC susceptibility of SA and SMA weld 

HAZ specimens.

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-13 10-12 10-11 10-10 10-9 10-8 10-7 10-6

As-welded (Spec. 85-7A)

As-welded (Spec. 85-7B)

500oC @ 24 h (Spec. 85-1ATT)

CGRenv (m/s)

CGRair (m/s)

304 SMA weld HAZ

0.75 dpa

289°C, 300–600 ppb DO

Open symbols: as-welded

Close symbols: 24h @ 500oC

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-13 10-12 10-11 10-10 10-9 10-8 10-7 10-6

As-welded (Spec. GG6TA)

500oC @ 24h (Spec. GG3TA-TT)

CGRenv (m/s)

CGRair (m/s)

304L SA weld HAZ

~2 dpa

289°C, 300–600 ppb DO

Open symbols: as-welded

Close symbols: 24h @ 500oC

(a) (b)

Figure 62. Cyclic CGRs for thermally treated and as-welded (a) SMA HAZ specimens at 0.75 dpa and 

(b) SA HAZ specimens at 2 dpa. 

4.2.3 Constant-Load CGR for Nonirradiated and Irradiated Weld HAZ Specimens

Figure 63 shows CGR results for nonirradiated SA and SMA HAZ specimens with and without 

heat treatment under constant-load conditions with periodical unloading.19 The green symbols are data 

points for 304L SA welds, and the brown symbols are for 304 SMA welds. For nonirradiated 1/4T-CT 

specimens, the K-size criterion was not satisfied for any of the data points shown in Fig. 63. Therefore,

caution must be exercised when comparing these data with other CGR results. 

The constant-load CGRs of nonirradiated SA and SMA HAZ specimens (except one data point) is 

bounded by the NUREG-0313 curve that was developed based on sensitized SSs in oxygenate water (8 

ppm DO).50 For the 304L SA weld HAZ specimens (diamond symbols), SCC CGRs in high-DO water 

are about two times lower than the NUREG-0313 curve. Heat treatment at 500°C for 24 h did not affect 

the cracking behavior under constant load. While SCC CGR data for the as-welded 304 SMA HAZ 

specimen are unavailable in NWC, CGRs of the as-welded HAZ specimen in HWC are about 2-5 times 

lower than the NUREG-0313 curve, similar to the values exhibited by the 304L SA HAZ specimen under 

NWC. 

Figure 64 shows the constant-load CGR results for irradiated SA and SMA HAZ specimens. With 

irradiated yield strength, all data reported in Fig. 64 are validated for the K/size criterion specified in 

Eq. (3). A significant increase in constant-load CGR can be seen for all weld HAZ specimens with NWC 

after irradiation. At 0.75 dpa, constant-load CGRs for the HAZ materials of the SA and SMA welds in 

NWC are 8-10 times higher than the NUREG-0313 curve (Fig. 64a). The cracking behavior of the 304L 

SA HAZ and 304 SMA HAZ are very similar under the as-welded condition. The constant-load CGRs of 

thermally treated 304 SMA HAZ is lower but still significantly higher than the NUREG-0313 curve. At 

≈2 dpa, constant-load CGRs for as-welded SA and SMA HAZ in NWC are similar and are about eight
60

times higher than the NUREG-0313 curve. Thermally treatment, at 500°C for 24 h, had no effect on the 

IASCC susceptibility of 304L SA HAZ. For both neutron doses, the beneficial effect of HWC is evident. 

Except for thermally treated 304 SMA HAZ, CGRs measured in HWC are more than one order of 

magnitude lower than that in NWC. For both neutron doses, the constant-load CGRs in HWC are all 

below the NUREG-0313 curve, regardless of weld type and post-weld heat treatment. 

 

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30 35 40

As-welded, NWC, (Spec. GG5G-A)

24 h @ 500oC, NWC, (Spec. GG3B-A-TT)

24 h @ 500oC, NWC, (Spec. 85-3A-TT)

24 h @ 500

Experimental CGR (m/s)

oC, HWC, (Spec. 85-3A-TT)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

SA and SMA weld HAZ

nonIrradiated

~290°C

304L SA HAZ

304 SMA HAZ

Figure 63. Constant-load CGR versus stress intensity for nonirradiated SA and SMA weld HAZ 

specimens with and without heat treatment.

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30 35 40

304L SA HAZ (Spec. GG5T-B)

304 SMA HAZ (Spec. 85-7A)

304 SMA HAZ (Spec. 85-7B)

304 SMA HAZ (Spec. 85-1ATT)

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

SA and SMA weld HAZ

0.75 dpa 

~290°C

8 x NUREG-0313

Curve

As-welded

Thermally-treated for 24 h @ 500oC

Open Symbols: NWC BWR Env.

Closed Symbols: HWC BWR Env.

(a)

Figure 64. Constant-load CGR results versus stress intensity for SA and SMA HAZ specimens irradiated 

to (a) 0.75 dpa and (b) ≈2 dpa
61

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30 35 40

304L SA HAZ (Spec. GG6T-A)

304 SMA HAZ (Spec. 85-XA)

304L SA HAZ (Spec. GG3TA-TT)

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

SA and SMA weld HAZ

2 dpa 

~290°C

8 x NUREG-0313

Curve

As-welded

Thermally-treated for 24 h @ 500oC

Open Symbols: NWC BWR Env.

Closed Symbols: HWC BWR Env.

(b)

Figure 64. (Contd.)

4.2.4 Fracture Toughness of Weld HAZ Specimens

At around 300°C, fracture toughness for nonirradiated austenitic SSs is reported to be in the range 

of 200-600 kJ/m2. 65-67 Weld process and neutron irradiation both result in significant reductions in 

fracture toughness by introducing new embrittlement mechanisms. In the present study, fracture 

toughness measured in NWC BWR water was found to be similar between the SA and SMA HAZ 

specimens under the as-welded condition (Fig. 65). The JIC for the Type 304 SMA HAZ is about 94

kJ/m2 (specimen 85-7B) and 98 kJ/m2 (specimen 85-XA) at 0.75 and 2 dpa respectively, and the JIC for 

the Type 304L SA HAZ is about 108 kJ/m2 (specimen GG-6TA) at ≈2 dpa. Post-weld heat treatment at 

500°C for 24 h reduces the fracture toughness of the 304L SA HAZ from 108 kJ/m2 to 73 kJ/m2

(specimen GG3TA-TT). For the 304 SMA HAZ, the JIC of the heat-treated specimen is significantly 

lower, approximately 34 kJ/m2 (specimen 85-2ATT). The fracture toughness obtained in these HAZ 

specimens is comparable or slightly lower than that of other austenitic SSs irradiated to the same dose 

level. 
62

0

50

100

150

200

0 2 4

J-Integral (kJ/m2)

Neutron Exposure (dpa)

SA and SMA weld HAZ

Tested at ~290oC in NWC BWR env.

304 SMA HAZ, thermally-treated 

at 500OC for 24 h 

304L SA HAZ, as welded

304 SMA HAZ, 

As-welded

304L SA HAZ, thermally-treated 

at 500OC for 24 h

Figure 65. Fracture toughness of irradiated 

Type 304L SA and 304 SMA weld HAZ 

specimens tested in NWC BWR water.

4.3 Effect of Sensitization 

Type 304 SS specimens thermally sensitized at 600°C for 10.5 and 24 h were included in the 

present study. The specimens were irradiated to two dose levels, 0.75 and 2 dpa, prior to being tested in 

the simulated BWR water environment. Figure 66 shows the cyclic CGR results for these sensitized 

specimens in NWC. In general, environmental enhancement of CGRs was readily achieved in the

sensitized specimens at both dose levels. Elevated growth rates were observed at around ≈10 -9 m/s for 

the 0.75-dpa specimens. Slightly higher cracking rates were observed in the 2-dpa specimen. For the two

sensitization conditions (600°C for 10.5 h and 24 h), environmental enhancement was nearly identical in 

NWC at 0.75 dpa. The cyclic cracking behavior seems to be unaffected by the degree of sensitization in 

irradiated Type 304 SS. This observation implies that thermal sensitization may not be the key factor for 

the IASCC process. With respect to the contribution of sensitization, irradiation effects may play a more 

important role in elevating SCC susceptibility in irradiated Type 304 SS.

Constant-load CGRs for the irradiated sensitized Type 304 SS are shown in Fig. 67. For both 

sensitization conditions (e.g. 10.5 h or 24 h at 600°C), the constant-load CGRs with NWC at 0.75 dpa are 

less than two times higher than the NUREG-0313 curve. At this dose level, the constant-load CGRs for 

sensitized SS are much lower than those of the weld HAZ specimens (represented by the “8 x NUREG0313” line). The beneficial effect of HWC chemistry is again evident for the sensitized SS. A 

considerable reduction in CGR can be seen in HWC for Type 304 SS sensitized for 24 h. Meanwhile, the 

available data at 2 dpa do not show a significant increase in the growth rate. This finding may imply a 

weak dose dependence for constant-load CGR in sensitized Type 304 SS. The constant-load CGRs 

obtained at slightly different K levels are fairly consistent for the two sensitization conditions. This 

observation again confirms that the degree of sensitization does not play a significant role in promoting 

SCC in irradiated materials. The effects of neutron irradiation must be a dominant factor contributing to 

the IASCC susceptibility. 
63

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-13 10-12 10-11 10-10 10-9 10-8 10-7 10-6

Sensitized for 10.5 h, 0.75 dpa

(Spec. 85-1TT)

Sensitized for 24 h, 0.75 dpa

(Spec. 85-5A1TT)

Sensitized for 10.5 h, 2 dpa

(Spec. 85-3TT)

CGRenv (m/s)

CGRair (m/s)

Type 304 SS sensitized @ 600oC

289°C, 300–600 ppb DO

Figure 66. Cyclic CGRs for sensitized Type 304 SS irradiated to 0.75 and 2 dpa.

 

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30

Sensitized for 10.5 h, 0.75 dpa 

(Spec. 85-1TT)

Sensitized for 24 h, 0.75 dpa

(Spec. 85-5A1TT)

Sensitized for 10.5 h, 2 dpa

(Spec. 85-3TT)

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

Type 304 SS, sensitized @ 600oC

Tested in BWR water, ~ 290oC

8 x NUREG-0313

Curve

Open Symbols: NWC BWR Env.

Closed Symbols: HWC BWR Env.

Figure 67. Constant-load CGRs versus stress intensity of sensitized Type 304 SS irradiated to 0.75 and 

2 dpa.
64

The fracture toughness of sensitized Type 304 SS is somewhat higher than that of the as-welded 

HAZ specimens. The JIC values of the specimens sensitized for 10.5 h, irradiated to 0.75 dpa (85-1TT) 

and 2 dpa (85-3TT) are 182 kJ/m2 and 145 kJ/m2 respectively. The 2-dpa specimen sensitized for 24 h

(85-5A2TT) shows a slightly higher JIC value (223 kJ/ m2) when tested in air.

4.4 Effect of Material Types

The CGR/J-R test results from the Halden-I study are used to illustrate the influence of material 

type and material chemistry on IASCC behavior. Three austenitic SSs (Type 304L, 316, and 316L SSs)

irradiated to three dose levels (0.45, 1.35, and 3 dpa) are included in the discussion. Table 14 shows the 

composition of these alloys, and experimental details can be found in references [19], [20], and [68]. 

Table 14. Composition of austenitic stainless steels (wt%)

Steel Type Heat ID Ni Si P S Mn C N Cr Mo O

304L C3 9.10 0.45 0.020 0.003 1.86 0.024 0.074 18.93 0.12 0.014

316L C16 12.32 0.42 0.026 0.003 1.65 0.029 0.011 16.91 2.18 0.016

316 C21 10.45 0.61 0.035 0.002 1.23 0.060 0.016 16.27 2.10 0.014

Figure 68 shows the cyclic CGR data at different dose levels for comparison. At 0.45 dpa (Fig. 68

a), no environmental enhancement was detected for Type 304L SS (specimen C3-A), while moderate 

enhancement was observed for Type 316 SS (specimen C21-A). All data points meet the K/size 

requirement in Eq. (3). With increasing dose, environmentally enhanced cracking became evident, and 

the difference in cyclic CGRs for different austenitic SSs was also diminished. At 1.35 and 3 dpa (Fig. 68

b and c), cyclic CGRs for Type 304L and 316 SSs were nearly identical. The low-carbon Type 316 SS 

(specimen C16-B) showed slightly lower cyclic CGRs than the normal-carbon-content Type 316 SS 

(specimen C21-C) at 3 dpa. Note that the nickel content in the Type 316L SS is also higher than that in

the Type 316 SS. So, the slightly lower cyclic CGRs observed in Type 316L SS may be due to its high 

nickel content. 

Constant-load CGRs of austenitic SSs at the three dose levels are plotted in Fig. 69 along with the 

NUREG-0313 and the 8-times-NUREG-0313 lines. At 0.45 dpa (Fig. 69a), two data points for Type 

304L SS are invalid for the K/size criterion by Eq. (3). The remaining CGRs satisfy the K/size criterion. 

The CGRs at 0.45 dpa in simulated BWR NWC are very close to the NUREG-0313 curve that was 

developed based on the CGR data for nonirradiated and sensitized SSs in high-DO water. The CGRs for 

Type 316 SS (specimen C21-A) are slightly higher than those of Type 304L SS (specimen C3-A). In 

general, IASCC susceptibility for both austenitic SSs at this dose level is comparable to or slightly higher 

than that of nonirradiated sensitized SSs. 

With the increase of irradiation dose, IASCC susceptibility becomes evident at 1.35 dpa. Constantload CGRs are elevated considerably at this dose level as shown in Fig. 69b. The cracking behavior 

between Type 304L (specimen C3-B) and 316 SSs (specimen C21-B) is very similar, and the measured 

constant-load CGRs are between 5 and 10 times higher than the NUREG-0313 curve. The beneficial 

effect of HWC is clearly demonstrated, and constant-load CGRs in HWC are at least one order of 

magnitude lower than those in NWC. An additional increase in irradiation dose to 3 dpa did not elevate 

the CGRs further in NWC, as shown in Fig. 69c (open symbols). The constant-load CGRs in NWC are 

still scattered around the 8-times-NUREG-0313 curve. Type 316 SS (specimen C21-C) exhibits slightly 

higher CGRs than those of Type 304L (specimen C3-C) and 316L SSs (specimen C16-B). Compared to 

the 1.35-dpa data, the most noticeable change at 3 dpa is the CGRs in HWC (closed symbols in Fig. 69c). 

The beneficial effect of HWC becomes less pronounced, especially for the Type 304L and 316 SSs under 

high applied load. This observation will be discussed further in the following section. 
65

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-13 10-12 10-11 10-10 10-9 10-8 10-7 10-6

304L SS (Spec. C3-A)

316 SS (Spec. C21-A)

CGRenv (m/s)

CGRair (m/s)

Austenitic SS

0.45 dpa

289°C, 300–350 ppb DO

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-13 10-12 10-11 10-10 10-9 10-8 10-7 10-6

304L SS (Spec. C3-B)

316 SS (Spec. C21-B)

CGRenv (m/s)

CGRair (m/s)

Austenitic SS

1.35 dpa

289°C, 300–350 ppb DO

(a) (b)

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-13 10-12 10-11 10-10 10-9 10-8 10-7 10-6

304L SS (Spec. C3-C)

316 SS (Spec. C21-C)

316L SS (Spec. C16-B)

CGRenv (m/s)

CGRair (m/s)

Austenitic SS

3 dpa

289°C, 300–350 ppb DO

Figure 68. Cyclic CGRs for austenitic SSs 

irradiated to (a) 0.45 dpa, (b) 1.35 dpa, and (c) 

3 dpa.

(c)

Additional information on the influence of materials chemistry on IASCC susceptibility was 

obtained from slow-strain-rate-tensile (SSRT) tests.24,69-72 In the Halden-I and -II studies, numerous 

SSRT tests were performed to screen a wide range of materials with different composition. A higher 

IASCC susceptibility was found for low-carbon and high-sulfur austenitic SSs. A sulfur content below 

0.002 wt.% was found to be critical for IASCC resistance in austenitic SSs in high-DO BWR water. With 

a similar sulfur composition, a low-carbon austenitic SS is more susceptible to IASCC than normalcarbon SSs. A minimum carbon content is needed to suppress IASCC susceptibility in the low-sulfur 

steels.70 Thus, there may be a trade-off in minimizing susceptibility to IASCC and minimizing 

susceptibility to IG cracking resulting from conventional sensitization. Type 316 SSs were found to be 

generally more resistant to IASCC than Type 304 SSs with the same sulfur and carbon composition. The 

presence of molybdenum or the higher nickel concentration in Type 316 SSs may be responsible for the 
66

difference in IASCC susceptibility between Type 304 and 316 SSs.74,75 The oxygen content is also 

critical for IASCC susceptibility because high oxygen content may increase the possibility of IG crack 

initiation.76 

Elevated cracking susceptibility has been reported in nonirradiated austenitic SSs with high silicon 

content.77,78 The high CGRs observed in these high-silicon alloys were also insensitive to corrosion 

potential.77 This observation suggests that radiation-induced segregation (RIS) of silicon to grain 

boundaries could also have a profound effect on IASCC susceptibility. It is well known that silicon 

concentration at the grain boundaries can be elevated significantly under neutron irradiation. At about 3 

dpa, silicon concentration as high as 4-5wt.% has been detected at grain boundaries in SSs.

79 With this 

level of silicon concentration, cracking resistance could be affected. However, despite the possible 

detrimental effect imposed by the RIS of silicon, the influence of silicon on IASCC susceptibility remains 

to be verified in irradiated SSs. 

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30 35 40

304L SS (Spec. C3-A)

316 SS (Spec. C21-A)

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

Austenitic SSs

0.45 dpa

289°C, BWR water

Symbols with a + indicate 

invalid K/size criterion

(a)

Figure 69. Constant-load CGRs versus stress intensity for austenitic SSs irradiated to (a) 0.45 dpa, (b) 

1.35 dpa, and (c) 3 dpa.
67

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30 35 40

304L SS (Spec. C3-B)

316 SS (Spec. C21-B)

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

8 x NUREG-0313

Curve

Open Symbols: NWC BWR Env.

Closed Symbols: HWC BWR Env.

Austenitic SSs

1.35 dpa

289°C, BWR water

(b)

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30 35 40

304L SS (Spec. C3-C)

316 SS (Spec. C21-C)

316L SS (Spec. C16-B)

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

Halden I, 3 dpa

289°C

8 x NUREG-0313

Curve

Open Symbols: NWC BWR Env.

Closed Symbols: HWC BWR Env.

Austenitic SSs

3 dpa

289°C, BWR water

(c)

Figure 69. (Contd.)
68

4.5 Effect of Water Chemistry

It is accepted that SCC susceptibility is a strong function of corrosion potential. A low corrosionpotential system is believed to suppress the crack-tip dissolution rate and thus reduce the rate of crack 

growth substantially.

29,80,81 In the above sections, we have demonstrated repeatedly that HWC can 

effectively reduce the severity of cracking in the BWR water environment. In this section, the CGR 

results from the Halden-I and -II studies concerning HWC will be presented again in a collective way to 

further illustrate the effect of water chemistry.

Figure 70 shows two cyclic CGR tests performed on Type 304L and 316L SSs in NWC and HWC. 

The Type 304L SS was irradiated to 1.35 dpa, and the Type 316L SS was irradiated to 3 dpa. Both 

materials exhibit a significant amount of environmental enhancement under cyclic loading in NWC (open 

symbols in Fig. 70). When the same materials were tested under cyclic loading in HWC, much weaker 

environmental enhancement was observed. The CGR data for NWC and HWC are best represented by

the Shack/Kassner model48 :

m CGR CGR CGR env air = +α⋅ air , (25)

where the coefficient α and the exponent m are derived by fitting of the experimental data. As shown by 

Fig. 70, the CGR in HWC is more than one order of magnitude lower than that in NWC under a similar 

cyclic loading condition. 

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-13 10-12 10-11 10-10 10-9 10-8 10-7 10-6

304L SS

316L SS

CGRenv (m/s)

CGRair (m/s)

Austenitic SSs

1.35 dpa, and 3 dpa

289°C, BWR water

Green Symbols: 1.35 dpa

Red Symbols: 3 dpa

Open Symbols: NWC BWR Env.

Closed Symbols: HWC BWR Env.

NWC

HWC

Figure 70. Cyclic CGRs for Type 

304L and 316L SSs in NWC and 

HWC tests.

All constant-load CGR data in the Halden-I and -II studies concerning the water chemistry effect 

are presented in three groups in Fig.71: sensitized and non-sensitized SSs and weld HAZ. Data for NWC 

and HWC are indicated by open and closed symbols, respectively. The NUREG-0313 curve is included

as a reference. Despite large scatter of the data, the CGRs for NWC (open symbols) and HWC (closed 

symbols) separate into two groups. The CGRs are much higher in NWC than in HWC for all materials 
69

and most irradiation doses included in the present study. The only exception is the data points for the 3-

dpa specimens tested in the high-K range. 

Relatively high CGRs under HWC were also reported by other authors for high-dose specimens at

high applied load.60,82,83 This abnormal cracking behavior in HWC was discussed in the context of 

K/size criterion. It has been argued that loss of constraint under a high applied load may compromise the 

test results.59 However, no compelling explanation has been given to correlate the constraint condition 

with the growth rate. Nonetheless, the diminishing HWC beneficial effect in high-dose specimens under 

high applied load may be more complex than simple mechanics (loss of constraint) can explain. A 

comprehensive investigation on the possible interactions between the crack tip plasticity and local water 

chemistry may be required to fully understand this phenomenon. 

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30

GG5T-B

85-7A

85-7B

85-1ATT

GG6T-A

GG3TA-TT

85-XA

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

SA and SMA weld HAZ

0.75 and 2 dpa 

~290°C, BWR water

Open Symbols: NWC

Closed Symbols: HWC

Specimen IDs:

(a)

Figure 71. Constant-load CGRs versus stress intensity in NWC and HWC for (a) weld HAZ, (b) 

sensitized 304 SS, and (c) wrought SSs.
70

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30

85-1TT

85-5A1TT

85-3TT

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

Sensitized 304 SS

0.75 and 2 dpa 

~290°C, BWR water

Open Symbols: NWC

Closed Symbols: HWC

Specimen IDs:

(b)

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30

304L SS

316 SS

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

Halden I, 3 dpa

289°C

Green Symbols: 1.35 dpa

Red Symbols: 3 dpa

Open Symbols: NWC

Closed Symbols: HWC

Austenitic SSs

1.35 and 3 dpa

289°C, BWR water

(c)

Figure 71. (Contd.)
71

4.6 Effect of Neutron Dose

Previous Halden reports70,71 have addressed the influence of neutron dose on IASCC in SSRT

tests. They contend that IASCC susceptibility is affected by irradiation hardening, embrittlement, and 

radiation-induced microstructural and microchemistry changes. All these irradiation effects are dosedependent, and therefore, IASCC behavior must vary with neutron dose as well. The SSRT tests

indicated that the percentage of IG cracking on the fracture surface is minimal below 0.45 dpa in high-DO 

water. With an increase of dose beyond 1.35 dpa, the fraction of IG cracking increases rapidly for the

fracture surface.

71,72 At ≈3 dpa, some austenitic SSs exhibit nearly 100% IG fracture, while others still 

retain fairly ductile fracture morphology. This observation indicates that the effect of dose on IASCC

behavior is different from one material to another. 

Since IASCC is a complex function of material and environmental variables, it is not easy to single 

out the dose effect from the CGR test results. To minimize the influence of other factors, the Halden-I 

CGR results in NWC on Type 304L, 316 and 316L SSs at 0.45, 1.35 and 3 dpa are used for the analysis. 

The mechanical properties of these specimens are similar for each of the three dose levels. Figure 72

shows the cyclic CGRs of these austenitic SSs. With increasing dose from 0.45 to 3 dpa, cyclic CGRs in

the test environment start to deviate from the diagonal line at an increasingly higher growth rate. The 

rising environmental enhancement under the cyclic loading condition indicates an increased IASCC 

susceptibility. 

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-13 10-12 10-11 10-10 10-9 10-8 10-7 10-6

0.45 dpa

1.35 dpa

3 dpa

CGRenv (m/s)

CGRair (m/s)

Austenitic SSs (304L, 316)

289°C, 300–350 ppb DO

0.45 dpa

1.35 dpa

3 dpa

Figure 72. Dose dependence of cyclic CGRs for Type 304L and 316 SSs.

The constant-load CGRs are plotted in Fig. 73 for the same group of specimens in NWC. The data 

points at the same dose level are fitted with the power law in Eq. (10). The coefficient A is determined to 

be 2.3×10-13, 1.25×10-12, and 1.15×10-12 m/s for 0.45, 1.35, and 3 dpa, respectively. An increasing trend 
72

can be seen for the constant-load CGR between 0.45 and 1.35 dpa, but no change is evident between 1.35 

and 3 dpa for CGR. This trend may imply a saturation effect of constant-load CGR in the high dose

region. 

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30

CGR-0.45dpa

CGR-1.35dpa

CGR-3dpa

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

Types 304L, 316, 316L SSs

289°C, BWR NWC env.

NUREG-0313

Curve

A=2.3E-13

A=1.2E-12

Figure 73. Constant-load CGRs versus stress intensity for Type 304L, 316 and 316L SSs at different 

doses.

The effect of neutron irradiation on fracture toughness has been discussed in detail in a previous 

Halden report.20 Irradiation embrittlement is mainly responsible for the reduction in fracture toughness in 

irradiated materials. The results of BWR irradiated materials fall within the scatter band but are close to 

the lower bound of the data obtained from fast reactor irradiations. The fracture toughness of various 

austenitic SSs decreases rapidly with irradiation dose from nonirradiated values of 400-600 kJ/m2 to less 

than 100 kJ/m2. At 3-5 dpa, the JIC of some austenitic SS can be as low as 15 kJ/m2. Limited data 

showed little environmental effect on the measured fracture toughness in BWR water. The J-R tests 

conducted in air and in water environment yielded approximately the same results. Figure 74 summarizes 

the fracture toughness data obtained in the Halden study. These data are also compared with data from 

similar studies. In the previous Halden study, the threshold neutron dose for a sharp decrease in fracture 

toughness was found to be about 0.3 dpa for austenitic SSs. Above 5 dpa, little plasticity is anticipated in 

austenitic SSs prior to fracture. 
73

0

200

400

600

800

0 4 8

JAPEIC BB

Japeic CT

JAPEIC SR

GE CT

VTT CT

CF-8M_air

C19

C16

SA HAZ

SMA HAZ

304 Sensitized

IT+GBE

J-Integral (kJ/m2)

Neutron Exposure (dpa)

Austenitic SSs

ANL data Figure 74. Dose dependence of fracture 

toughness for austenitic SSs. Dashed lines 

represent the scatter band for the fast reactor 

data at 350-450°C. 20

(GE = General Electric Nuclear Energy, 

JAPEIC = Japan Power Engineering and 

Inspection Corporation, CT = compaction 

tension, BB = bend bar, and SR = short rod.)

4.7 Effect of GBE Treatment

4.7.1 Crack Growth Rates for GBE and Non-GBE Type 304 and 316 SSs

Grain boundary engineering (GBE) is a thermomechanical processing technique that systematically 

increases the population of low Σ coincident site lattice (CSL)*

Figure 75 shows the cyclic CGR results for non-GBE and GBE Type 304 and 316 SSs tested in

NWC. All specimens were irradiated to 2 dpa prior to the CGR tests. The open and closed symbols 

represent the non-GBE and GBE materials, respectively, for all figures shown in this section. At 2 dpa, 

environmentally enhanced cracking was quite evident for Type 304 SSs with or without GBE treatment. 

The GBE Type 304 SS showed somewhat higher susceptibility to environmental cracking than the nonGBE Type 304 SS under cyclic loading. For Type 316 SSs, environmental enhancement was less 

pronounced compared to Type 304 SSs under a similar test condition. The GBE Type 316 SS showed 

some degree of environmental enhancement, while the non-GBE Type 316 SS displayed essentially no 

environmental effect under cyclic loading. In general, the cyclic CGRs indicate that the present GBE 

treatment did not improve IASCC behavior. In contrast, IASCC susceptibility may have been elevated, 

and higher cracking growth rates can be seen for GBE Type 304 and 316 SSs. 

grain boundaries in materials.73 Because 

the low Σ (≤29) grain boundaries usually exhibit good resistant to IG cracking in nonirradiated 

materials,37-40 GBE has also been considered to be a potential mitigation method for IASCC. Type 304 

and 316 SSs with and without GBE were included in the present study to investigate the effect of GBE 

treatment on IASCC susceptibility under conditions relevant to BWRs. Fractions for CSL boundaries are

given in Table 2 for the non-GBE (baseline) and GBE Type 304 and 316 SSs.

 

* An approach to describe atomic coherency of grain boundaries. A CSL grain boundary is defined by Σ, the reciprocal of the 

density of coincident sites. 
74

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-13 10-12 10-11 10-10 10-9 10-8 10-7 10-6

IT304-01, nonGBE 304

GB304-01, GBE 304

CGRenv (m/s)

CGRair (m/s)

GBE vs nonGBE 304 SS

2 dpa

289°C, 300–400 ppb DO

Open symbols: nonGBE

Close symbols: GBE

10-13

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-13 10-12 10-11 10-10 10-9 10-8 10-7 10-6

IT316-01, nonGBE 316

GB316-01, GBE 316

CGRenv (m/s)

CGRair (m/s)

GBE vs nonGBE 316 SS

2 dpa

289°C, 600–700 ppb DO

Open symbols: nonGBE

Close symbols: GBE

(a) (b)

Figure 75. Cyclic CGRs for irradiated non-GBE and GBE treated (a) Type 304 SSs and (b) Type 316 

SSs tested in BWR NWC. 

10-12

10-11

10-10

10-9

10-8

10-7

5 10 15 20 25 30

nonGBE 304 SS

GBE 304 SS

nonGBE 316 SS

GBE 316 SS

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

NUREG-0313

Curve

nonGEB and GBE SSs

2 dpa

289°C, BWR NWC

Figure 76. Constant-load CGRs versus stress 

intensity for non-GBE and GBE treated Type 

304 and 316 SSs. 

The constant-load CGRs for all non-GBE and GBE materials are shown in Fig. 76. The CGRs for 

both baseline materials (Type 304 and 316 SSs) are very close to or below the NUREG-0313 curve, 

which represents the CGR values for nonirradiated, sensitized austenitic SSs. However, the GBE-treated 

materials show considerably higher CGRs under the same test condition. The high constant-load CGRs 

for GBE materials verified the unfavorable outcome from the present GBE treatment. Both cyclic and 
75

constant-load CGR tests demonstrated that the GBE treatment employed in the present study did not 

improve the cracking resistance of the baseline materials. The GBE-treated Type 304 and 316 SS 

specimens exhibit a significantly higher susceptibility to IASCC. 

4.7.2 Fracture Toughness of GBE and Non-GBE Type 304 and 316 SSs

Fracture toughness J-R tests were conducted on all non-GBE and GBE specimens after the CGR 

tests. The fracture toughness data for these alloys are shown in Fig. 77. Both baseline materials show 

relatively poor fracture toughness at 2 dpa. The JIC values are 83 and 89 kJ/m2 for non-GBE Type 304 

and 316 SSs, respectively. The fracture toughness of the two GBE materials is higher than that of the 

non-GBE materials. Apparently, the applied GBE treatment improves the irradiated fracture toughness 

for the baseline alloys. The different impact of GBE treatment on the CGR and fracture toughness results 

may indicate that fracture modes are different in the CGR and fracture toughness tests for these materials. 

0

100

200

300

400

500

600

700

800

0 0.5 1 1.5 2 2.5 3

GBE 304 SS

nonGBE 304 SS

J (kJ/m2)

Crack Extension (mm)

Type 304 SS 

2 dpa

J = 392∆a0.74

JIC = 138 kJ/m2

J = 132∆a0.3

JIC = 83 kJ/m2

0

100

200

300

400

500

600

700

800

0 0.5 1 1.5 2 2.5 3

GBE 316 SS

nonGBE316 SS

J (kJ/m2)

Crack Extension (mm)

Type 316 SS 

2 dpa

J = 290∆a0.36

JQ = 178 kJ/m2

J = 163∆a0.42

JIC = 89 kJ/m2

(a) (b)

Figure 77. Fracture toughness J-R curves for non-GBE and GBE (a) Type 304 SSs and (b) 

Type 316 SSs. 

4.7.3 Hardness and Microstructure of Nonirradiated GBE and Non-GBE Type 304 and 316 SSs

Hardness and microstructure of nonirradiated GBE and non-GBE Type 304 and 316 SSs were

examined. Figure 78 shows the hardness measurement performed on the baseline and GBE materials 

prior to irradiation. The applied GBE treatment produced inconsistent results for the two types of 

austenitic SSs. For Type 316 SS, the GBE-treated material was harder than its baseline material. For 

Type 304 SS, the GBE-treated material was softer. 
76

Figure 78. Hardness measurements of nonGBE and GBE Type 304 and 316 SSs.

(a) non-GBE Type 304 SS (b) GBE Type 304 SS

(c) non-GBE Type 316 SS (d) GBE Type 316 SS

Figure 79. Grain morphology revealed by electron backscattered diffraction.

All four alloys were examined by electron backscattered diffraction to reveal their grain boundary 

configurations. Figure 79 shows the grain morphology for each alloy with and without GBE treatment. 

The quantitative measurements of grain structure are given in Fig. 80. The average grain size for GBE 

Type 304 SS (ASTM grain size number, G 6.5) is much larger than that of its baseline material (ASTM G 

8.0). This increase in grain size by GBE processing may be responsible for the low hardness of GBE 

Type 304 SS shown in Fig. 78. For Type 316 SS, the grain size change due to GBE processing is 

minimal. The ASTM grain size numbers for GBE and non-GBE 316 SSs are 6.0 and 6.5 respectively. 

The CSL fraction is quantified for each alloy according to the Brandon criterion, and the results are given 

in Fig. 80. The increase in CSL fraction is moderate for Type 304 SS, and much less for Type 316 SS. 

The population for low-Σ boundaries is 67% for GBE Type 304 SS and about 58% for GBE Type 316 SS. 

Transmission electron microscopy (TEM) images are shown in Fig. 81 for GBE and non-GBE

Type 316 SS. While the non-GBE Type 316 SS is clean and free from precipitates, the GBE Type 316 
77

SS shows many precipitates scattered along grain boundaries and inside grains. The size of the 

precipitates is around 100-200 nm. The precipitates were determined to be M6

The TEM examination of GBE and non-GBE 304 SSs revealed a larger grain size but fewer

precipitates in the GBE material (as shown in Fig. 82). Stress concentration resulting from dislocation 

pile-ups could be seen at grain boundaries. Non-uniform dislocation networks were also observed in 

some areas in the GBE 304 SS. The larger grain size exhibited in the GBE material may be the primary 

reason for the high cracking rate demonstrated in the CGR test.

C (η-carbide), M23C6, or 

M7C. The increasing amount of precipitates found in the GBE-treated 316 SS may be responsible for the 

elevated IASCC susceptibility after irradiation. 

304 SS 304 SS

316 SS 316 SS

Figure 80. Grain size distribution (left) and CSL length fraction (right) for GBE and non-GBE Type 304 

and 316 SSs 
78

(a) non-GBE Type 316 SS

(b) GBE Type 316 SS

Figure 81. Transmission electron microscopy images for (a) non-GBE and (b) GBE Type 316 SSs. 

Both GBE 316 and 304 SSs showed unfavorable microstructure (e.g. precipitation, or large grain 

size) that is known to be critical to IASCC susceptibility. This vulnerable microstructure may indicate a 

poorly handled GBE process. The thermomechanical treatment employed may also have introduced

sensitization in the alloys. It appears that the potential benefit resulting from the increased CSL 

boundaries to ≈60% is insufficient to offset the detrimental effects caused by the susceptible 

microstructure (e.g. large grain size, precipitation) in Types 304 and 316 SSs. The ineffectiveness of the 

applied GBE treatment suggests that the beneficial effect warranted by a higher fraction of CSL 

boundaries may be secondary for limiting crack propagation in austenitic SSs. Other detrimental effects 
79

introduced by large grain size or brittle precipitates may be more important in determining the rate of 

crack growth. The beneficial effects of GBE treatment on CGRs have to be evaluated with higher CSL 

fractions in austenitic SSs. Any practical use of GBE SSs will require the development of procedures to 

better control GBE treatment.

(a) (b)

Figure 82. TEM images for (a) non-GBE and (b) GBE 304 SSs.
80

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81

5 Summary

Crack growth rate tests have been conducted in simulated BWR environments on 1/4T-CT 

specimens of Type 304L SA and 304 SMA weld HAZ, sensitized Type 304 SS, and Type 304 and 316 

SSs with and without GBE treatment. The specimens were irradiated in the Halden reactor to 0.75 and ≈2 

dpa. Based on the cyclic and constant-load CGR results, IASCC susceptibility was evaluated for the 

tested materials as a function of neutron dose, processing conditions, and water chemistry. Fracture

toughness tests were also carried out in the BWR environment after the CGR tests. Combining the 

present results with those from Halden-I and -II tests, we analyzed the effects of neutron dose, water 

chemistry, alloy compositions, and welding and processing conditions on IASCC. The effect of neutron 

irradiation on the fracture toughness of the tested materials was also evaluated at dose levels relevant to 

BWR internals. The major findings of this investigation are as follow:

(1) Crack growth tests require a K-dominated loading condition to ensure the test results are 

independent of specimen size and geometry. A K/size criterion is, therefore, required to maintain a

minimum constraint in front of a crack tip. In the present study, the effect of strain softening,

exhibited by high-dose specimens, on the plastic zone size was analyzed. Assuming 20% strainsoftening, the plastic zone size was determined to be only 11% larger than that required by the 

ASTM standard E-399 for fracture toughness tests. This analysis confirmed that the K/size 

criterion is satisfied in most of the tests conducted in the present study.

(2) Neutron irradiation elevated the cyclic CGRs significantly for all HAZ specimens at all dose levels

regardless of weld type and as-welded or thermally treated conditions. The constant-load CGRs for 

as-welded SA and SMA HAZ were about 8-10 times higher than the NUREG-0313 curve. A postweld heat treatment at 500°C for 24 hours did not affect the constant-load CGRs of the irradiated 

HAZ specimens significantly. Fracture toughness of thermally treated HAZ specimens was slightly 

lower than that of as-welded HAZ materials. In general, the effect of low-temperature sensitization 

appears to be insignificant for the IASCC susceptibility of HAZ materials for both SA and SMA 

welds. 

(3) Type 304 SS specimens were thermally sensitized at 600°C for 10.5 h and 24 h to investigate the 

effect of degree of sensitization on IASCC. The specimens sensitized at the two conditions showed

similar cracking behavior under both cyclic and constant-load test conditions. The degree of 

sensitization appears to play a less significant role in the cracking susceptibility of irradiated 

compared with nonirradiated materials. Irradiation effects appear to be the dominating factor for

IASCC behavior.

(4) The influence of material type and material chemistry on IASCC susceptibility was evaluated by 

analysis of CGR and SSRT results obtained from previous Halden studies. For CGR tests, cyclic 

and constant-load CGRs of three austenitic SSs were compared at three dose levels in NWC and 

HWC. Within the scatter of the data, IASCC susceptibility appears to be insensitive to the material 

type in NWC. A comparison of CGRs in HWC was inconclusive because of the concern over 

invalidation of the K/size criterion. Significantly higher load was required to achieve some 

measurable CGRs in tests performed in HWC. For SSRT tests, the fraction of intergranular 

cracking was used to characterize the IASCC susceptibility of different alloys. The SSRT results 

indicated that minor alloying elements such as carbon, sulfur, oxygen, and silicon may play an 

important role in the IASCC process. Low-carbon and high-sulfur SSs were found to be prone to 

IASCC. Sulfur content below 0.002 wt.% was found to be critical for IASCC resistance in 
82

austenitic SSs in high-DO BWR water. The effect of alloying element must be addressed in the 

context of radiation-induced segregation. 

(5) Stress corrosion cracking is a strong function of corrosion potential. It has been repeatedly 

demonstrated that a low corrosion potential can effectively mitigate IASCC. In our analysis, both 

cyclic and constant-load CGR results showed at least one order of magnitude reduction in HWC

compared with NWC. However, a few tests conducted in the high-K range showed that the 

beneficial effect of HWC is absent in some high-dose specimens (>3 dpa). This unusual cracking 

behavior in HWC has often been explained with respect to the K/size criterion. Nonetheless, the 

insensitiveness to low corrosion potential demonstrated by some high-dose specimens should not be 

overlooked based solely on the loading condition. Several tests conducted under the validated Kconstraint condition have shown the HWC beneficial effect to be lost with Type 304L SS at doses

as low as 3 dpa. Since the dose dependence of the HWC effect is poorly understood, additional 

data are needed for dose greater than 3 dpa.

(6) Halden-I crack growth results in NWC at three doses were used to illustrate the dose dependence of 

IASCC behavior. A trend of increased environmental enhancement with dose could be seen with 

the cyclic CGRs for this group of specimens. The constant-load CGRs of these specimens also 

increased sharply between 0.45 and 1.35 dpa. Above 1.35 dpa, the increasing trend of the constantload CGR became slower. Fracture toughness of austenitic SSs also decreased rapidly with neutron 

dose. The threshold dose for a sharp decrease in fracture toughness is about 0.3 dpa, and little 

plasticity is expected prior to fracture for austenitic SSs above 5 dpa. 

(7) Crack growth tests were performed on irradiated GBE and non-GBE Types 304 and 316 SSs in 

simulated BWR water environments. A beneficial effect on IASCC resistance was not readily 

revealed in these alloys with about 60% CSL boundaries. Microstructure examination on the 

nonirradiated GBE and non-GBE materials showed significant precipitation and large grain size in 

the GBE alloys. The ineffectiveness of GBE treatment demonstrated in the present study suggests 

that the beneficial effect provided by a high fraction of CSL boundaries may be overwhelmed by 

detrimental effects introduced by brittle precipitates or large grain size.

Future work for IASCC study in BWR environment should be focused on high-dose specimens 

(>2x1021 n/cm2, >3 dpa) in HWC. At present, crack growth rate data for irradiated SSs at high doses are 

scarce and the dose-dependence of the effect of HWC on IASCC behavior is not clear. While it is clear 

that there is a beneficial effect of HWC on CGR for nonirradiated and low-dose (< 3 dpa) SSs, it is not 

clear that the same degree of mitigation occurs for materials at higher doses. Cracking and deformation 

mechanisms must remain the same with increasing dose for HWC to be as effective. The assumption that 

HWC is equally effective at higher doses may not necessarily be true because irradiation hardening, loss 

of ductility and work hardening behavior continue to evolve beyond 3 dpa in SSs. Additional work is

thus needed to further characterize the effectiveness of HWC at high doses.
83

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pp. 337-348, 2005.
NUREG/CR

-7018 Irradiation

-Assisted Stress Corrosion Cracking of 

Austenitic Stainless Steels in BWR Environments May 2010 
Irradiation-Assisted Stress 

Corrosion Cracking of

Austenitic Stainless Steels in 

BWR Environments

Office of Nuclear Regulatory Research 

NUREG/CR-7018

ANL-09/17