NUREG/CR-6964 

ANL-07/12

Crack Growth Rates and 

Metallographic Examinations of 

Alloy 600 and Alloy 82/182 from 

Field Components and 

Laboratory Materials Tested in 

PWR Environments

Office of Nuclear Regulatory Research 
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NUREG/CR-6964 

ANL-07/12 

Crack Growth Rates and 

Metallographic Examinations of 

Alloy 600 and Alloy 82/182 from 

Field Components and 

Laboratory Materials Tested in 

PWR Environments 

Manuscript Completed: February 2008 

Date Published: May 2008 

Prepared by 

B. Alexandreanu, O.K. Chopra, and W. J. Shack 

Argonne National Laboratory 

Argonne, IL 60439 

S. Crane and H.J. Gonzalez, NRC Project Managers 

NRC Job Code Y6388 

Office of Nuclear Regulatory Research 

 

iii 

Abstract 

In light water reactors, components made of nickel-base alloys are susceptible to environmentally 

assisted cracking. This report summarizes the crack growth rate results and related metallography for 

field and laboratory-procured Alloy 600 and its weld alloys tested in pressurized water reactor (PWR) 

environments. The report also presents crack growth rate (CGR) results for a shielded-metal-arc weld of 

Alloy 182 in a simulated PWR environment as a function of temperature between 290°C and 350°C. 

These data were used to determine the activation energy for crack growth in Alloy 182 welds. The tests 

were performed by measuring the changes in the stress corrosion CGR as the temperatures were varied 

during the test. The difference in electrochemical potential between the specimen and the Ni/NiO line 

was maintained constant at each temperature by adjusting the hydrogen overpressure on the water supply 

tank. The CGR data as a function of temperature yielded activation energies of 252 kJ/mol for a double-J 

weld and 189 kJ/mol for a deep-groove weld. These values are in good agreement with the data reported 

in the literature. The data reported here and those in the literature suggest that the average activation 

energy for Alloy 182 welds is on the order of 220-230 kJ/mol, higher than the 130 kJ/mol commonly used 

for Alloy 600. The consequences of using a larger value of activation energy for SCC CGR data analysis 

are discussed. 

Foreword

v 

This report presents crack growth rate (CGR) data and the results of the corresponding fracture surface 

and metallographic examinations from cyclic loading and primary water stress-corrosion cracking 

(PWSCC) tests of two nickel-base Alloy 182 (A182) weldments. These weldments are typical of those 

used in vessel penetrations and piping butt welds in nuclear power plants. The effect of crack orientation 

with respect to dendrite orientation is the most significant variable investigated in this study. However, 

this report also compiles data from other laboratories that describe the effects of material composition, 

loading characteristics, and chemistry of the aqueous environment. The PWSCC growth rates described 

for the A182 specimens in the report are comparable to the CGR that characterize the performance of 

Alloy 600 (A600). 

This report is one of a series of reports documenting the results of CGR testing in vessel head penetration 

materials. This report focuses on the Alloy 82 (A82) and A182 weld metals. This report also presents 

results from tests of the A600 base metal. The researchers tested (1) a laboratory-fabricated, shieldedmetal-arc deposit of A182; (2) a weldment sample from the J-groove weld of a control rod drive 

mechanism nozzle from the Davis–Besse Nuclear Power Plant; and (3) a hot leg nozzle-to-safe end weld 

from the Virgil C. Summer Nuclear Station. 

The objective of the research was to generate PWSCC crack growth rates in A82 and A182 assist NRC in 

determining the validity and acceptability of licensee flaw analyses Title 10 of the Code of Federal 

Regulations Part 50.55a (10 CFR 50.55a). In addition, the results will also assist NRC in its review of 

licensees’ requests for relief to some in-service inspection requirements of 10 CFR 50.55a, if submitted. 

 

vii 

Contents 

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

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

Executive Summary.................................................................................................................................... xiii

Acknowledgments ...................................................................................................................................... xv

Acronyms and Abbreviations .................................................................................................................... xvii

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

2 Experimental ........................................................................................................................................ 3

2.1 Material and Specimen Design .............................................................................................. 3

2.2 Test Facilities.......................................................................................................................... 5

2.2.1 Facility with 6-Liter Autoclave .......................................................................... 5

2.2.2 Facility with 1-Liter Autoclave............................................................................ 6

2.2.3 Facility with a 2-Liter Autoclave......................................................................... 8

2.2.4 Primary Water Environment ................................................................................ 10

2.3 Test Procedure ........................................................................................................................ 13

2.4 Analysis of Crack Growth Rate Data .................................................................................... 14

3 Microstructural Characterization ........................................................................................................ 17

3.1 Laboratory-Prepared Alloys................................................................................................... 17

3.2 Field Alloys............................................................................................................................. 19

4 Crack Growth Test Results ................................................................................................................. 23

4.1 Double-J Weld Specimen A182-1......................................................................................... 23

4.2 Deep-Groove Weld Specimen CT933H-1 ............................................................................ 26

4.3 Deep-Groove Weld Specimen CT933H-2 ............................................................................ 28

5. Discussion ............................................................................................................................................ 31
 

viii 

5.1. SCC Growth Rates.................................................................................................................... 31

5.1.1 Alloy 600 ............................................................................................................. 31

5.1.2 Ni-Alloy Welds...................................................................................................... 32

5.2. Activation Energy for SCC Crack Growth.............................................................................. 36

5.3. Consequence of Using a Larger Activation Energy............................................................ 39

5.4 Cyclic Crack Growth Rates in PWR Environment .............................................................. 40

5.4.1 Alloy 600................................................................................................................ 40

5.4.2 Ni-Alloy Welds...................................................................................................... 43

6 Summary............................................................................................................................................... 47

References ................................................................................................................................................... 49

Appendix A: Crack Growth Rate Data for Davis-Besse and V. C. Summer Ni-base alloys................ 55
 

ix 

Figures 

1. Configuration of the 1-T and 1/2-T CT specimens used for this study............................................ 3

2. Schematic of the weld joint design and weld passes for Alloy 182 SMA double-J weld and the 

deep-groove weld. ............................................................................................................................ 4

3. Orientation of the CT specimens from the Alloy 182 SMA double-J weld and the deep-groove 

weld. .................................................................................................................................................. 4

4. Photograph of the specimen load train. .............................................................................................. 5

5. Schematic diagram of the recirculating 6-liter autoclave system. .................................................... 6

6. Photograph of the specimen load train for the 1-liter autoclave....................................................... 7

7. Schematic diagram of the recirculating 1-liter autoclave system. .................................................... 8

8. Layout of the 2-liter autoclave system. .............................................................................................. 9

9. Photograph of the specimen load train for the 2-liter autoclave....................................................... 9

10. Schematic diagram of the recirculating 2-liter autoclave system. .................................................... 10

11. Hydrogen concentration at Ni/NiO phase transition as a function of temperature.......................... 12

12. Dendritic microstructure on a laboratory-preapred Alloy 182 weld specimen................................ 17

13. Micrograph showing one of the matrix precipitates observed on the surface of sample 3A and 

EDX spectra resulting from the bulk and the precipitate............................................................... 17

14. OIM maps from a plane parallel to the direction of dendrites showing the grain orientations. ..... 18

15. Grain boundary character distribution for laboratory-prepared welds. ............................................ 19

16. Microstructures observed on planes 03-1 and 03-2 of the CRDM nozzle alloy.............................. 20

17. Microstructure observed on the Alloy 182 J-groove weld specimen from Davis-Besse. ............... 20

18. Microstructure of the butter Alloy 182............................................................................................... 21

19. Microstructure of the weld Alloy 82 from the V.C. Summer plant.................................................. 21

20. Grain boundary sliding at the weld-pipe interface............................................................................. 21

21. Fracture surface of specimen A182-1: optical image and SEM image. ........................................... 24
 

x 

22. Crack length vs. time for Alloy 182 weld specimen A182-1 in simulated PWR environment 

during test periods precracking, 1-2, 3, and 4-7............................................................................. 25

23. Fracture surface of Alloy 182 weld specimen CT933H-1 tested in simulated PWR environment 

at temperatures of 290-310°C.......................................................................................................... 26

24. Crack length vs. time for Alloy 182 weld specimen CT933H-1 in simulated PWR environment 

during test periods precracking-2, 3-4, 5-6, and 7-8. ..................................................................... 27

25. Fracture surface of Alloy 182 weld specimen CT933H-2 tested in simulated PWR environment 

at temperatures of 290-350°C.......................................................................................................... 29

26. Crack length vs. time for Alloy 182 weld specimen CT933H-2 in simulated PWR environment 

during test periods 1-5, 6-12, and 13-18. ........................................................................................ 30

27. CGR data for Alloy 600 from Davis-Besse CRDM nozzle #3 in PWR water at 316°C under 

constant load and log-normal distribution of constant ! for 26 heats of Alloy 600.................... 31

28. Transition from TG to IG fracture mode in specimen N3CL-1........................................................ 32

29. SCC crack growth data for the Argonne Alloy 182 welds plotted as a function of K and 

compared with the available data for Alloy 182 and 82 welds in simulated PWR 

environment. ..................................................................................................................................... 33

30. CGR data for Alloy 182 from the Davis-Besse CRDM nozzle J-groove weld in PWR 

environment at 316°C under constant load..................................................................................... 34

31. Comparison of the CGR data for the V.C. Summer weld and butter alloys with the data 

obtained on laboratory-prepared welds in the present study and V.C. Summer weld and 

butter alloys by Jacko et al............................................................................................................... 35

32. Log-normal distribution of constant ! for several heats of Alloy 82 and 182. ............................... 35

33. Micrographs from the fracture surface of weld specimens: Davis-Besse J11CC-1, and V.C. 

Summer BCR-01. ............................................................................................................................. 36

34. Temperature dependence of the CGR data for Alloy 82 obtained at Bechtel Bettis, ETH, and 

Lockheed Martin at temperatures between 290°C and 360°C. ..................................................... 37

35. Temperature dependence of the CGR data for Alloy 182 obtained at Westinghouse and CEA at 

temperatures between 290°C and 340°C. ....................................................................................... 37

36. Temperature dependence of the CGR data for Alloy 182 obtained at ANL at temperatures 

between 290°C and 350°C............................................................................................................... 38

37. Comparison of observed activation energies for Alloy 182 with a normal distribution. ................ 38

38. Cumulative distribution of the parameter ! in the CGR relationship for Ni-alloy welds with the 

data normalized to 325°C using an activation energy of 130 kJ/mol and 220 kJ/mol................. 39
 

xi 

39. CGR data for the V.C. Summer weld and Davis-Besse weld alloys................................................ 40

40. Cyclic CGR data for several Alloy 600 heats in deaerated water at 320°C under cyclic loading.. 41

41. CGR data for Davis-Besse CRDM nozzle #3 Alloy 600 in PWR water at 316°C under cyclic 

loading............................................................................................................................................... 42

42. CGR data in PWR environment for Alloy 182 SMA weld at 320°C, and Alloy 82 and Alloy 52 

welds at 315°C as a function of the growth rates for Alloy 182 weld in air. ............................... 43

43. Transition from TG to IG fracture and higher magnification micrograph of the boxed area 

showing the tip of a few secondary cracks. .................................................................................... 44

44 CGR data for Alloy 182 SMA weld-metal specimens as a function of growth rates for 

Ni-weld alloy in air. ......................................................................................................................... 44

45. Cyclic CGR data for Alloy 182 and Alloy 82 from the Davis-Besse CRDM nozzle J-groove 

weld and V.C. Summer reactor vessel nozzle-to-pipe weld.......................................................... 45

46. Micrographs of the fracture surface of specimen WCR-01 and WLR-01. ...................................... 46
 

xii 

Tables 

1. Welding process and conditions for various weld passes. ................................................................ 4

2. Chemical composition (wt.%) of Alloy 182 weld metals. ................................................................ 4

3. Test conditions for various temperature and hydrogen tank overpressures..................................... 12

4. Test conditions and crack growth data for specimen A182-1 of Alloy 182 SMA weld in 

PWR water........................................................................................................................................ 23

5. Test conditions and crack growth data for specimen CT933H-1 of Alloy 182 SMA weld in 

PWR water........................................................................................................................................ 27

6. Test conditions and crack growth data for specimen CT933H-2 of Alloy 182 SMA weld in 

PWR water........................................................................................................................................ 29
 

xiii 

Executive Summary 

Ni-base alloys used as construction material in light water reactors (LWRs) have experienced stress 

corrosion cracking (SCC). Such cracking was first observed in steam generator tubes, but it has also 

occurred in Ni alloys used in applications such as instrument nozzles and heater thermal sleeves in the 

pressurizer and penetrations for control-rod drive mechanisms in the reactor-vessel closure heads. In 

operating plants, weld metals composed of Alloys 82 and 182 are used with Alloy 600. Less cracking has 

been observed in the weld materials than in the Alloy 600 wrought material. However, laboratory tests 

indicate that in PWR coolant environments, the SCC susceptibility of Alloy 182 may be greater than that 

of Alloy 600, while the susceptibility of Alloy 82 may be comparable to that of Alloy 600. 

A program is being conducted at Argonne National Laboratory (ANL) to evaluate the resistance of 

Ni alloys and their welds to environmentally assisted cracking in simulated LWR coolant environments. 

The report summarizes the crack growth rate (CGR) results and related metallography for field and 

laboratory-prepared Alloy 600 and Ni-base alloy welds obtained in the program. This report also presents 

CGR results for Alloy 182 shielded-metal-arc welds in a simulated PWR environment as a function of 

temperature. These results were used to determine the activation energy for crack growth in Alloy 182 

welds. 

Metallographic examinations of the Ni–based alloy welds showed that the weld structure consists 

of vertically aligned columnar grains and dendrites. Orientation imaging microscopy (OIM), a 

diffraction-based technique, was used to determine the orientations of the grains and the type of grain 

boundaries present. The results show that a large proportion (70%) of the grain boundaries are random or 

high-angle boundaries, which are more susceptible to cracking than those in specific orientation 

relationships, also known as coincident site lattice boundaries. In addition, the OIM maps show the 

presence of clusters of grains that share similar orientations. 

Tests used to determine the activation energy for SCC crack growth were performed in a simulated 

PWR environment at temperatures between 290°C and 350°C (554 - 662°F). The tests were performed 

by transitioning a fatigue crack into a stress corrosion crack in a specimen, and by measuring the changes 

in the stress corrosion CGR as the temperatures were varied during the test. The difference in 

electrochemical potential between the specimen and the Ni/NiO line was maintained constant at each 

temperature by adjusting the hydrogen overpressure on the water supply tank. 

Activation energies for SCC crack growth in Alloy 182 SMA weld alloy were calculated from the 

results for SCC CGR as a function of temperature. Activation energies of 252 kJ/mol (59 kcal/mol) were 

obtained for the double-J weld material and 189 kJ/mol (33 kcal/mol) for the deep-groove weld material. 

These values are in agreement with the range of values reported in the literature. The data reported here 

and those in the literature suggest that the average activation energy for Alloy 182 welds is on the order of 

220-230 kJ/mol, which is higher than the 130 kJ/mol (31 kcal/mol) commonly used for Alloy 600. The 

activation energy data for Alloy 82 are more limited. The average value from the literature data is 

!169 kJ/mol, which is somewhat higher than the value for Alloy 600. The consequences of using these 

larger values for the analysis of SCC CGR data were analyzed. The analysis showed that when an 

activation energy of 220 kJ/mol is used for calculation of the temperature-corrected CGRs for Alloy 182, 

the 75th percentile value of ln(!), which has been the value used for a disposition curve, becomes -27.54, 

yielding a value of 1.1 x 10-12 at 325°C (617°F) for the ! parameter. This results in a disposition curve 

for Ni-alloy welds that is !36% lower than the curve based on data normalized with 130 kJ/mol; thus, the 

use of the 130 kJ/mol value leads to somewhat more conservative results. 
 

xiv 

In addition to the tests on laboratory prepared welds, CGR tests were performed on weldment 

samples from a Davis–Besse CRDM J-groove nozzle weld and from a V.C. Summer hot–leg nozzle to 

safe-end weld. For the Davis-Besse Alloy 182 weld specimens, the SCC CGRs under constant load are 

an order of magnitude lower than the disposition curve proposed for Alloy 182 weld metals. For the V.C. 

Summer weld specimens, the CGR data under constant load were lower than data obtained by others on 

the same alloys and lower than the proposed disposition curve. The data from the field welds correspond 

to values in the lower quartile of the distribution of crack growth rate data obtained on laboratory 

specimens. These results are not so statistically unlikely that they can be said to prove that there is a 

difference between field and laboratory welds, but such a difference would be consistent with the 

observations that operating experience indicates that PWSCC appears to occur more frequently in 

wrought Ni-base Alloy 600 components than in the weld metal Alloys 82 and 182 used with Alloy 600, 

despite the fact that in laboratory tests in PWR coolant environments, the stress corrosion cracking (SCC) 

susceptibility of Alloy 182 is usually found to be greater than that of Alloy 600, while that of Alloy 82 is 

comparable to that of Alloy 600. 

Such a difference is also supported by the metallographic observations of extensive secondary 

cracking along the crack faces of specimens from field welds that is not observed in tests on laboratory 

welds. The reasons for these differences are unclear. The most obvious difference between field and 

laboratory welds is the degree of structural constraint imposed on the welds during the welding process, 

which is typically much greater for the field welds on actual components. However, metallographic 

studies on the welds away from the crack planes do not show hot cracks so that the secondary cracks 

seem to be part of the corrosion process. In any case, the use of the current database developed on 

laboratory welds would appear to be conservative. 

The report also discusses the crack growth rate results for actual Alloy 600 component material and 

the environment enhancement of CGRs under cyclic loading. The SCC CGRs of the Alloy 600 nozzle 

from Davis-Besse are a factor of 4–8 higher than those of the median curve for Alloy 600. The growth 

rates correspond to the !95th percentile of the various data sets used in developing the median curve, i.e., 

the nozzle material exhibits very high susceptibility to SCC. The material exhibits predominantly IG 

fracture, even during precracking. 

The environmental enhancement of CGRs under cyclic loading was determined relative to the 

CGRs that would be expected under the same loading conditions for Ni-weld alloys in air. The cyclic 

CGRs were analyzed by a superposition model to establish the individual contributions of mechanical 

fatigue, corrosion fatigue, and SCC. 
 

xv 

Acknowledgments 

The authors thank E. Listwan, T. Galvin, and R. Clark for their contributions to the experimental 

effort. The authors are also grateful to W. H. Cullen, Jr., for many helpful discussions. This work is 

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

Job Code Y6388; Program Managers: S. Crane and H. J. Gonzalez. 

 

xvii 

Acronyms and Abbreviations 

ANL Argonne National Laboratory 

ASTM American Society for Testing and Materials 

BWR Boiling Water Reactor 

C Circumferential 

CF Corrosion Fatigue 

CGR Crack Growth Rate 

CRDM Control Rod Drive Mechanism 

CSL Coincident Site Lattice 

CT Compact Tension 

DO Dissolved Oxygen 

ECP Electrochemical Potential 

EDX Energy Dispersive X-ray 

GBC Grain Boundary Coverage 

GBCD Grain Boundary Character Distribution 

GTA Gas Tungsten Arc 

HAB High Angle Boundary 

HAZ Heat Affected Zone 

IG Intergranular 

L Longitudinal 

LAB Low Angle Boundary 

LWR Light Water Reactor 

MA Mill Annealed 

NRC Nuclear Regulatory Commission 

NWC Normal Water Chemistry 

OIM Orientation Imaging Microscopy 

PWSCC Primary Water Stress Corrosion Cracking 

PWR Pressurized Water Reactor 

R Radial 

RTZ Rolled Transition Zone 

S Side 

SA Solution Annealed 

SCC Stress Corrosion Cracking 

SEM Scanning Electron Microscopy 

SHE Standard Hydrogen Electrode 

SMA Shielded Metal Arc 

SS Stainless Steel 
 

xviii 

T Transverse 

TG Transgranular 

TT Thermally Treated 
 

1 

1 Introduction

Nickel-base alloys used as construction material in pressurized water reactors (PWRs) have 

experienced primary water stress corrosion cracking (PWSCC). Such cracking has occurred in Alloy 600 

and Alloy 182/82 components such as steam generator tubes, heater thermal sleeves and penetrations in 

the pressurizer, penetrations for the control rod drive mechanisms (CRDMs) in the reactor pressure vessel 

head, bottom-mounted instrumentation nozzles, and dissimilar metal welds between the low-alloy steel 

reactor vessel nozzle and stainless steel (SS) safe-end or pipe. Initiation of PWSCC involves a significant 

incubation period that depends primarily on temperature and material susceptibility. In general, cracking 

has occurred much earlier in the Alloy 600 base metal than in the Alloy 182 and Alloy 82 welds and at 

higher temperature locations within the reactor coolant system.1 Also, small-bore pipes and tubes have 

cracked earlier than larger components.1

The PWSCC of Alloy 600 steam generator tubes in PWRs has been studied intensively.2-4 In 

general, cracking occurs in regions of high residual stress due to cold work, such as the tube roll transition 

zone (RTZ), U-bends, tube denting locations, and plugs and sleeves.5-7 The RTZ at the top of the tube 

sheet is the most common location for such cracking, and the cracks are most frequently axial, although 

circumferential cracks also occur. The PWSCC at U-bends has been associated with high residual 

stresses from the bending process. The PWSCC at dented tube support plates has also been related to the 

presence of stresses caused by the denting, and this problem has largely been resolved. Tube plugs, 

which typically have high residual stresses, are themselves prone to PWSCC. Heats of Alloy 600 show a 

wide variability in their susceptibility to PWSCC.8-10 The use of improved thermal processing and a 

more resistant material, Alloy 690, has greatly reduced the incidence of this cracking in steam generator 

tubes. 

The earliest location of PWSCC in components other than the steam generator tubes has been in the 

base metal of instrument nozzles and heater thermal sleeves in the pressurizer.11 In these cases, the 

cracks were always axial and occurred at locations at the highest temperature and in materials with yield 

strength as low as !240 MPa (!35 ksi), which is the minimum value allowed. 

The other major locations of PWSCC of Alloy 600 components are at the nozzle penetrations for 

the CRDM in the reactor vessel closure heads.12 In the fall of 1991, during an overpressurization test, a 

leak was discovered in the CRDM nozzle at the Bugey 3 plant in France. Metallurgical evaluations 

indicated that the leak was caused by PWSCC.13 The main crack had initiated in Alloy 600 base metal 

and propagated into the Alloy 182 weld metal. Subsequent inspections of CRDM penetrations in the 

early 1990s in foreign PWRs indicated that !6.5% of the nozzles in French plants had axial cracks on the 

nozzle inner surface, while only !1.25% of the nozzles that were inspected in other plants had axial 

cracks. Inspection of the CRDM nozzles in seven plants in the United States (Point Beach 1, Oconee 2, 

Cook 2, Palisades, North Anna 1, Millstone 2, and Ginna) at this time suggested that the cracking was 

much less frequent than in the French plants. None of the cracks found in U.S. plants was through-wall, 

and until late 2000, no more leaks were found in pressure-vessel head penetrations. 

In November 2000, leaks from axial through-wall cracks were identified at Oconee Unit 1, and in 

February 2001, at Arkansas Nuclear One Unit 1.14 During the next 15 months, inspections at Oconee 

Units 2 and 3 and follow-up inspection at Unit 1 identified both axial and OD-initiated circumferential 

cracks in reactor vessel head penetrations.15 The presence of circumferential cracks, in particular, raised 

concerns regarding structural integrity.16,17 Also, in October 2000, significant boron deposits were 

discovered near the Loop “A” reactor vessel nozzle to the hot-leg reactor coolant pipe weld at the 
 

2 

V. C. Summer plant.18 Ultrasonic inspection of the pipe revealed an axial crack and a short, intersecting 

circumferential crack at the top of the pipe in the dissimilar metal butt weld. The failed region was 

repaired with a new SS spool piece and Alloy 152 weld filler metal. Earlier in 2000, two shallow axial 

flaws were found in the outlet nozzle-to-safe-end weld of Ringhals Unit 3, and four axial indications were 

found in the same region of Ringhals Unit 4, in Sweden.19 Cracks were also found in pressure-vessel 

head penetrations at North Anna Unit 2,20 the Davis-Besse nuclear power plant,21 and more recently, in 

the bottom-mounted instrumentation nozzles at South Texas Unit 1.22,23 

Other instances of PWSCC in components fabricated from Ni alloys include cracking of the nozzlecoupling-to-vessel weld at the steam-generator lower head bowl drain, observed at Catawba Unit 2 in 

2001,24 and cracking in the heat-affected zone of an Alloy 600 safe end in a power-operated relief valve, 

near the connection to the pressurizer at the Palisades nuclear plant in 1993.25 

In some incidents of PWSCC (e.g., cracking of the CRDM nozzle at the Davis-Besse plant),21

leakage of the primary coolant resulted in significant degradation of ferritic steel components. At 

Davis-Besse, downstream of nozzle #3, a triangular cavity, about 127 mm (50 in.) wide and 178 mm 

(70 in.) long, had penetrated completely through the thickness of the low-alloy steel used for the reactor 

pressure-vessel head, leaving only a layer of SS cladding. 

Plant operating experience indicates that PWSCC appears to occur more frequently in wrought 

Ni-base Alloy 600 components than in the weld metal Alloys 82 and 182 used with Alloy 600.1 

However, in laboratory tests in PWR coolant environments, the stress corrosion cracking (SCC) 

susceptibility of Alloy 182 is usually found to be greater than that of Alloy 600, while that of Alloy 82 is 

comparable to that of Alloy 600. This apparent inconsistency between field and laboratory experience is 

an issue that needs to be better understood. 

This report summarizes the technical data obtained to date in the ANL study on Ni-alloy cracking 

in the PWR environment. New data are presented on CGRs in Alloy 182 welds in PWR environments at 

temperatures between 290°C and 350°C (554 - 662°F). As with the previous SCC tests conducted at 

ANL, precracking was carried out in the PWR environment and continued with loading cycles with 

increasing load ratios and increasing rise times prior to setting those samples at constant load to determine 

the SCC CGRs. This approach assured a complete SCC engagement and a uniform crack front. Each test 

was complemented by a detailed fractographic examination. The CGR data were analyzed to determine 

the activation energy for SCC growth rates in Alloy 182 welds in the PWR environment. The results 

from the present study and data available in the literature have been reviewed to determine the effects of 

material and environmental parameters on the cyclic and constant-load CGRs in these alloys. The alloys 

being investigated include laboratory-procured Alloys 600 and 690,26-31 laboratory-prepared deep-groove 

and double-J welds of Alloy 182,32 Alloy 600 from Davis-Besse CRDM nozzle #3, Alloy 182 from a 

J-groove weld in Davis-Besse nozzle #11, and Alloys 182 and 82 from a hot-leg nozzle-to-pipe weld in 

the V.C. Summer reactor coolant system.33 The results have been compared with the existing CGR data 

to determine the relative susceptibility of these alloys to PWSCC and to determine whether the crack 

growth data in the field materials are consistent with our general understanding of CGRs in Ni-base 

alloys. 
 

3 

2 Experimental 

2.1 Material and Specimen Design 

All crack growth rate tests on Ni alloys and Ni-base welds were conducted in simulated PWR 

environments at temperatures of 290-350°C in accordance with American Society for Testing and 

Materials (ASTM) Designation E 647, “Standard Test Method for Measurement of Fatigue Crack Growth 

Rates,” and ASTM E-1681, “Standard Test Method for Determining a Threshold Stress Intensity Factor 

for Environment-Assisted Cracking of Metallic Materials under Constant Load.” The tests were 

performed on both 1-T and 1/2-T compact tension (CT) specimens; the configurations of the CT 

specimens are shown in Fig. 1. 

LAY MARKS TO FALL

PERPENDICULAR TO CENTER

LINE OF RADIUS "M SURFACE" 

DETAIL 'C' DETAIL 'M'

40°

.025

.076 R

±2°

1.19 2.39

1.19 R

1.27

9.53

±0.25

±0.05

12.7

C

25.4

2.39

CENTERED

C

30.5 ±0.25

14.0 ±0.25

30.5 'C'

XXXXXX

SPECIMEN 

 I. D.

12.7

'M'

2 THRU HOLES

12.7 DIA.

+0.25

- 0.00

63.5±0.51

±0.25

12.7

2.26 DIA. DRILL 9.53 DP.

4-40 UNC-2B TAP 6.35 ± 0.51 DP.

ONE EA. ON TOP & BOTTOM FACE 

TOP FACE

±0.25

50.8

61.0±0.51 14.0 ±0.25

±0.25

(a) 

14.00

14.00

6.60

6.60

1.60

CENTERED

6.00 DIA.

2 THRU HOLES

+.05

- .00 30.00

28.00

'M'

A .02

A .02

B

B .02

A .02

A .02

12.00

24.00

4.00

1.53 DIA

2 THRU HOLES

4.00

4.00

2.00

4.5

2.00

1.53 DIA. 2 HOLES 8.0 DEEP.

XXX-X

SPECIMEN ID

C

C .02

C .02

9.0 A

.90 R

.90

DETAIL 'M'

(b) 

Figure 1. Configuration of the (a) 1-T and (b) 1/2-T CT specimens used for this study. 

For the tests focused on determining the activation energy for SCC crack growth in Alloy 182 weld, 

one 1-T CT specimen (A182-1) was machined from a laboratory-prepared double-J weld (Fig. 2a), and 

two 1/2-T specimens (CT933H-1 and CT933H-2) were machined from a deep-groove filled weld 

(Fig. 2b). The welds were prepared following ASME Boiler and Pressure Vessel Code, Section IX. The 

double-J weld was produced by joining two Alloy 600 plates and was prepared by 48 weld passes. Root 

passes 1 to 5 involved gas tungsten arc (GTA) welding with Alloy 82 filler/electrode, and the other 

passes, shielded-metal-arc (SMA) welding with Alloy 182 filler. A schematic of the weld design and 

various passes is shown in Fig. 2a. The conditions for each weld pass are listed in Table 1. During 

welding, the maximum inter-pass temperature was !120°C (250°F), and the weld surfaces were cleaned 

by wire brushing and grinding and were rinsed with de-mineralized water or alcohol. The deep-groove 

filled weld was prepared by using a 51-mm (2.0 in.) thick Alloy 600 plate with a deep groove that was 
 

4 

filled by several passes of SMA welding with Alloy 182 filler/electrode that was either 1/8 or 5/32 in. in 

size (Fig. 2b). The chemical compositions of the weld metals are given in Table 2. 

Table 1. Welding process and conditions for various weld passes. 

Weld 

Pass Process 

Filler 

Metal 

Filler/Electrode Size 

(in.) 

Current 

(A) 

Voltage 

(V) 

Travel Speed 

(in./min) 

1 - 5 GTA Alloy 82 3/32 185 - 215 21 - 22 2 - 4 

6 - 10 SMA Alloy 182 3/32 140 - 155 24 - 26 6 - 7 

11 -27 SMA Alloy 182 1/8 155 - 170 25 - 27 6 - 7 

28 - 48 SMA Alloy 182 5/32 170 - 180 26 - 28 6 - 7 

(a) (b) 

Figure 2. Schematic of the weld joint design and weld passes for (a) Alloy 182 SMA double-J weld and 

(b) the deep-groove weld (dimensions are in inches). 

Table 2. Chemical composition (wt.%) of Alloy 182 weld metals. 

Alloy ID (Heat) Analysis C Mn Fe S P Si Cu Ni Cr Ti Nb Co 

A 182 Spec. 0.10* 5.0-9.5 6.0-10.0 0.015* – 1.0* 0.5* Bal 13.0-17.0 1.0* 1.0-2.5 0.12*

A 182 Double-J ANL 0.0415 7.095 6.005 0.008 0.06 0.53 0.03 65 14.35 0.43 1.585 0.03 

A 182 Deep Groove ANL 0.04 7.08 6.82 0.005 0.025 0.35 0.03 70.44 13.81 0.30 1.06 0.02 *Maximum. 

Specimen A182-1 was cut from a double-J Alloy 182 SMA weld in the TS orientation,* as shown 

schematically in Fig. 3a. The two 1/2-T CT specimens, both in the TS orientation, were cut from the 

deep-groove Alloy 182 weld (Fig. 3b). In order to accommodate the 1-T CT dimensions, Alloy 600 plate 

was EB-welded as shown schematically in the figures. 

(a) (b) 

Figure 3. Orientation of the CT specimens from (a) the Alloy 182 SMA double-J weld and (b) the 

deep-groove weld. 

 

*The first letter represents the direction normal to the fracture plane and the second represents the direction of crack advance. As 

shown in Fig. 3, the three directions are: T = transverse, L = longitudinal, and S = side. 
 

5 

2.2 Test Facilities 

The CGR tests in simulated primary water environments were conducted in the following test 

facilities. Specimen A182-1 was tested in a 6-liter SS autoclave configured for 1-T CT specimens. One 

1/2-T CT specimen, CT933H-1, was tested in a 1-liter SS autoclave, while the companion 1/2-T CT 

specimen, CT933H-2, was tested in a 2-liter SS autoclave. This section provides a description of these 

testing facilities. 

2.2.1 Facility with 6-Liter Autoclave 

The 6-liter autoclave facility is typically used for CGR tests at temperatures not exceeding 320°C 

(608°F). The facility consists of a closed-loop electro-hydraulic material test system equipped with an 

extended column load frame rated at 89 kN (20,000 lb) maximum. The 5.7-liter Type 316 SS autoclave 

has a 175-mm (6.875-in.) outside diameter and is rated for a working pressure of 5050 psig (35 MPa) at 

343°C (650°F). The load frame is accessorized with an Instron Model 8800 control console, and a 

hydraulic pump. The autoclave is connected to a recirculating water system. The test facility also 

consists of a temperature control unit, a DC potential control console, a DC potential measurement unit, 

and a data acquisition and recording system. The autoclave, mounted within the load-frame, has been 

modified to permit a !19-mm (0.75-in.) dia shaft to load the test specimen through a “Bal-Seal” gland in 

the top of the autoclave cover. Up to three 1-T or 1/2-T CT specimens can be tested in series inside the 

autoclave. 

The test facility is designed for easy access to the specimens during assembly of the test train. The 

actuator assembly, consisting of the hydraulic actuator, load cell, autoclave plug, and the internal 

specimen load train, may be raised and lowered hydraulically to position the specimens at a convenient 

height. A photograph of the specimen load train is shown in Fig. 4. A CT specimen may be substituted 

for any or all of the three central in-line blocks. 

The autoclave is continuously supplied with simulated primary water solution from a feedwater 

tank. Figure 5 shows a schematic diagram of the water system. The water system consists of a feedwater 

storage tank, high-pressure pump, regenerative heat exchanger, autoclave preheater, test autoclave, 

electrochemical potential (ECP) cell, regenerative heat exchanger, back-pressure regulator, and return line 

to the feedwater tank. The storage tank has a hydrogen cover gas to maintain a desired dissolved 

hydrogen concentration in the water. In the once-through mode, the return line is connected to the drain. 

Water is circulated at relatively low flow rates (15-25 mL/min). 

Figure 4. 

Photograph of the specimen load train. 
 

6 

1

32

4

9

8

7

6

V2

V1

5

12 V3

29

38

28

12 12

13

14

15

24

26

22

17

23

18

19

20

25

V6

V4

V5

V8

V7

12

33

32

30

33 V15 V10

V11

V9

V10

V11

36 35

31

37

34

15

15

V15

V13

V14 V14

16

10

11

11

27

V16

V2

V19

V18

V17

V20

12

21

 1. COVER GASS SUPPLY TANK 20. HEAT EXCHANGER (HX) 

 2. TWO-STAGE HIGH-PRESSURE REGULATOR 21. DRAIN 

 3. LOW-PRESSURE REGULATOR 22. HX FEEDWATER OUTLET TC 

 4. FLOW METER 23. AUTOCLAVE & ECP CELL PREHEATER 

 5. COMPOUND VACUUM & PRESSURE GAUGE 24. HEAT EXCHANGER 

 6. PRESSURE RELIEF VALVE 25. COMMERCIAL AUTOCLAVE 

 7. VENT TO AIR & FLASH ARRESTOR 26. THERMOWELL 

 8. FEEDWATER STORAGE TANK 27. “BAL SEAL” RETAINER 

 9. SPARGE TUBE 28. ECP CELL 

10. FEEDWATER FILL PORT 29. HX EFFLUENT INLET TC 

11. WATER SAMPLE PORT 30. ECP CELL INLET TC 

12. SYSTEM BLEED PORT 31. AIR-COOLED COIL 

13. SOLENOID VALVE 32. BPR INLET TC 

14. HIGH-PRESSURE PUMP 33. BACK-PRESSURE REGULATOR (BPR) 

15. CHECK VALVE 34. PRESSURE RELIEF VALVE 

16. PRESSURE TRANSDUCER 35. CONDUCTIVITY METER 

17. RUPTURE DISC 36. PH METER 

18. ACCUMULATOR 37. ION EXCHANGE BED 

19. HIGH-PRESSURE GAUGE 38. ION EXCHANGE BED BYPASS LINE 

Figure 5. Schematic diagram of the recirculating 6-liter autoclave system. 

2.2.2 Facility with 1-Liter Autoclave 

The 1-liter autoclave facility is a modified version of an in-cell facility for conducting CGR tests on 

irradiated materials and is capable of test temperatures up to 310°C (590°F). The actuator assembly 

consisting of a hydraulic actuator, a 22-kN (5-kip) load cell, an autoclave cover plate, an internal 

specimen load train, and a furnace that is mounted on top of a portable wheeled cart that can be easily 

relocated. A 1-liter SS autoclave is installed inside the furnace for conducting tests in simulated PWR 

environments. The furnace and the autoclave body are mounted on a pneumatic cylinder and can be 

raised to enclose the internal specimen load train and the specimen during the test. Water is circulated 

through a port in the autoclave cover plate that serves both as inlet and outlet. 
 

7 

The 1/2-T CT specimen is mounted in the clevises with Zircaloy pins (Fig. 6). The two sets of 

clevises are at 90o opposition, forming a pair of universal joints. The specimen and clevises are kept 

electrically insulated from the load train by using oxidized Zircaloy pins and mica washers to connect the 

clevises to the rest of the load train. The crack length is monitored by reversing-current DC potential 

drop measurements. Platinum wires are used for the current and potential leads. The current leads are 

attached to SS split pins that are inserted into the holes at the top and bottom of the specimen. The 

potential leads are attached by screwing short SS pins into threaded holes in the specimen and attaching 

the platinum wires with in-line SS crimps. This system does not operate with a reference bar for 

correcting the DC potential. An Instron Model 8500+ Dynamic Materials Testing System is used to load 

the specimen. 

Figure 6. 

Photograph of the specimen load train for the 

1-liter autoclave. 

A schematic diagram of the recirculating water system is shown in Fig. 7. The recirculating water 

system consists of a storage tank, high pressure pump, an accumulator, regenerative heat exchanger, 

autoclave preheater, autoclave, ECP cell preheater, ECP cell, air-cooled coil, water-cooled chill block, 

Mity MiteTM back-pressure regulator, and return line to the storage tank. In the recirculating mode, the 

ECP cell is bypassed, and water from the autoclave is returned to the storage tank through the 

regenerative heat exchanger, water-cooled chill block, and back-pressure regulator. With the ECP cell in 

use, to avoid any possible contamination of chlorides from the reference electrode, the system is operated 

only in the once-through mode; water from the autoclave is circulated through the ECP cell preheater, 

ECP cell, air-cooled coil, water-cooled chill block, back-pressure regulator, and drain to the sump. Water 

is circulated at flow rates of 45-160 mL/min. 

The feedwater storage tank, manufactured by Filpaco Industries, has 130-L capacity and is 

constructed of stainless steel. The tank is designed for vacuum and overpressure to 414 kPa (60 psig). 

The storage tank has a hydrogen cover gas to maintain a desired dissolved hydrogen concentration in the 

water. 
 

8 

V9

1

3

2

4

7

6

V2

V1

5 V2

10

24

9

11

8 3

V3

V20

V4

V5 V6

V8

V7

12

13

15 14

23

20 19

18

17

16

6

6

22

25

26

27

28

29

30

31

32

21

8

33

34

33 V10

V11 V12

V13 V14

V15

V16

12

12 35

V18 V19

V17

 1. COVER GASS SUPPLY TANK 19. RUPTURE DISC 

 2. TWO-STAGE HIGH-PRESSURE REGULATOR 20. ACCUMULATOR 

 3. FLASH ARRESTOR 21. HEAT EXCHANGER (HX) 

 4. LOW-PRESSURE REGULATOR 22. AUTOCLAVE PREHEATER 

 5. FLOW METER 23. PREHEATER OUTLET TC 

 6. CHECK VALVE 24. AUTOCLAVE INLET TC 

 7. COMPOUND VACUUM & PRESSURE GAUGE 25. COMMERCIAL AUTOCLAVE 

 8. PRESSURE RELIEF VALVE 26. THERMOWELL 

 9. VENT TO AIR & FLASH ARRESTOR 27. “BAL SEAL” RETAINER 

10. FEEDWATER STORAGE TANK 28. ECP CELL PREHEATER 

11. SPARGE TUBE 29. ECP CELL 

12. WATER SAMPLE PORT 30. AIR-COOLED COIL 

13. FEEDWATER FILL PORT 31. WATER COOLED HEAT EXCHANGER 

14. FEEDWATER TANK RECIRCULATION PUMP 32. BPR INLET TC 

15. SOLENOID VALVE 33. BACK-PRESSURE REGULATOR (BPR) 

16. HIGH-PRESSURE PUMP 34. PH METER 

17. PRESSURE TRANSDUCER 35. CONDUCTIVITY METER 

18. HIGH-PRESSURE GAUGE 

Figure 7. Schematic diagram of the recirculating 1-liter autoclave system. 

2.2.3 Facility with a 2-Liter Autoclave 

The 2-liter autoclave test facility was built by Cortest Inc. and allows test temperatures of up to 

350°C (662°F). The servohydraulic test frame consists of a load cell, load train, autoclave, and furnace. 

The hydraulic actuator is mounted on bottom of the test frame, with the test train components suspended 

above it. The load cell is located at the bottom of the pull rod. An Instron Model 8800 control console is 

used to load the specimen. A 2-liter autoclave is installed inside the furnace for conducting tests in 

simulated LWR environments. The heater bands are mounted on the autoclave vessel. Figure 8 is a 

photograph showing the entire test system. 
 

9 

Figure 8. Layout of the 2-liter autoclave system. 

The load cage that contains the test specimen consists of the cover plate of the autoclave and a thick 

bottom plate separated by four compression rods (Fig. 9). The lower two-piece clevis assembly is 

fastened to the bottom plate of the cage, and the upper piece is connected to the pull rod. A 1/2T CT or 

dog-bone tensile specimen can be mounted between the clevises. The specimen and clevises are kept 

electrically insulated from the load train by using oxidized Zircaloy pins and mica washers to connect the 

clevises to the rest of the load train. The crack length is monitored by the same method as for the 

previous two facilities. Water is circulated through a port in the autoclave body, which serves both as 

inlet and outlet. A schematic diagram of the recirculating water system is shown in Fig. 10. 

Figure 9. 

Photograph of the specimen load train for the 

2-liter autoclave. 
 

10 

V9

1

3

2

4

7

6

V2

V1

5 V2

10

9

11

8 3

V3

V22

V4

V5 V6

V8

V7

12

13

15 14

25

20

19

18

17

16

6

6

22

27

28

29

30

31

32

33

21

8

34

35

34 V10

V11 V12

V13 V14

V15

V16

12

12 36

V18 V19

V17

26

23

24

V20

22

V21

 1. COVER GASS SUPPLY TANK 19. ACCUMULATOR 

 2. TWO-STAGE HIGH-PRESSURE REGULATOR 20. RUPTURE DISC 

 3. FLASH ARRESTOR 21. HEAT EXCHANGER (HX) 

 4. LOW-PRESSURE REGULATOR 22. DRAIN 

 5. FLOW METER 23. SYSTEM BLEED PORT 

 6. CHECK VALVE 24. HEAT EXCHANGER OUTLET TC 

 7. COMPOUND VACUUM & PRESSURE GAUGE 25. AUTOCLAVE PREHEATER 

 8. PRESSURE RELIEF VALVE 26. PREHEATER OUTLET TC 

 9. VENT TO AIR & FLASH ARRESTOR 27. COMMERCIAL AUTOCLAVE 

10. FEEDWATER STORAGE TANK 28. THERMOWELL 

11. SPARGE TUBE 29. “BAL SEAL” RETAINER 

12. WATER SAMPLE PORT 30. ECP CELL 

13. FEEDWATER FILL PORT 31. AIR-COOLED COIL 

14. FEEDWATER TANK RECIRCULATION PUMP 32. WATER COOLED HEAT EXCHANGER 

15. SOLENOID VALVE 33. BPR INLET TC 

16. HIGH-PRESSURE PUMP 34. BACK-PRESSURE REGULATOR (BPR) 

17. PRESSURE TRANSDUCER 35. PH METER 

18. HIGH-PRESSURE GAUGE 36. CONDUCTIVITY METER 

 

Figure 10. Schematic diagram of the recirculating 2-liter autoclave system. 

2.2.4 Primary Water Environment 

The simulated PWR feedwater contains 2 ppm Li as LiOH, 1000 ppm B as HBO3, !2 ppm 

dissolved hydrogen (!23 cm3/kg), and less than 10 ppb dissolved oxygen (DO). It is prepared from the 

laboratory supply of deionized water by first passing this water through a local filtration system that 
 

11 

includes a carbon filter, an Organex-Q filter, two ion exchangers, and a 0.2-mm capsule filter. The DO in 

the deionized water is reduced to <10 ppb by bubbling/sparging pure H2 through the water. To speed 

deoxygenation of a fresh tank of water, a vacuum may be applied to the feedwater tank at the vent port. 

The PWR water is prepared by dissolving boric acid and lithium hydroxide in 20 L of deionized water 

before adding the solution to the supply tank. The dissolved hydrogen in the water is calculated from the 

supply tank hydrogen pressure and temperature. 

Water samples are taken periodically to measure pH, resistivity, and DO concentration both 

upstream and downstream from the autoclave. An Orbisphere meter and CHEMetricsTM ampoules are 

used to measure the DO concentrations in the supply and effluent water. The redox and open-circuit 

corrosion potentials are monitored at the autoclave outlet by measuring the ECPs of platinum and an 

Alloy 600 electrode, respectively, against a 0.1 M KCl/AgCl/Ag external (cold) reference electrode. The 

measured ECPs, E(meas) (mV), were converted to the standard hydrogen electrode (SHE) scale, 

E(SHE) (mV), by the polynomial expression34 

-4 2 -6 3 E + 286.637 - 1.0032 ( T) + 1.7447 10 ( T) - 3.03004 10 ( T) (meas) ! " # ! " # ! " , (1)

where !T(°C) is the temperature difference of the salt bridge in a 0.1 M KCl/AgCl/Ag external reference 

electrode (i.e., the test temperature minus ambient temperature). 

The effect of the proximity of the corrosion potential for the SCC test to the potential at the Ni/NiO 

phase transition on crack growth rates is well-established. While the mechanistic origin is unclear, there 

are consistent data showing that CGRs of Ni-base alloys peak near the Ni/NiO phase boundary.35-37 

Thus, to isolate the effect of temperature, the tests are conducted so that the difference between the 

corrosion potential of the specimen and that of the Ni/NiO phase transition was maintained approximately 

constant. The Ni/NiO phase transition is very close to the H2/H2O phase transition, and the latter controls 

the corrosion potential of the simulated primary water environment. As such, the corrosion potential 

difference reduces to the difference of two hydrogen electrodes:37

2 test

test Ni / NiO

2 Ni / NiO

RT [H ] E E E ln

2F [H ]

! " # = $ = $ % & ' (

(2)

where: Etest = potential at test temperature T 

ENi/NiO = potential of the Ni/NiO transition at test temperature T 

R = universal gas constant 

 = 8.314 x 10-3 kJ/mol K (1.103 x 10-3 kcal/mol°R) 

F = Faraday constant 

 = 96485 C/mol 

[H2]test = hydrogen concentration at test conditions 

[H2]Ni/NiO = hydrogen concentration at the Ni/NiO transition 

 For two tests, conducted at temperatures T1 and T2 and satisfying the constant corrosion potential 

difference to the Ni/NiO line, Eq. 2 transforms into the following relationship between hydrogen 

concentrations and test temperatures: 

2 1 2 test2 2 test1

2 Ni / NiO test2 2 Ni / NiO test1

RT [H ] RT [H ] ln ln

2F [H ] 2F [H ] ! !

" # " # $ % $ = % & ' & '

(3)
 

12 

where: [H2]test2 = hydrogen concentration at test temperature T2

[H2]test1 = hydrogen concentration at test temperature T1 

[H2]Ni/NiO-test2 = hydrogen concentration at Ni/NiO transition for test 2 

[H2]Ni/NiO-test1 = hydrogen concentration at Ni/NiO transition for test 1 

 The hydrogen concentration at the Ni/NiO transition as a function of temperature was determined 

by fitting the data by Attanasio et al.36 with a fourth-order polynomial (Fig. 11). The best fit was found 

to be given by the following expression: 

2 3 -8 4 [H ] (T) = -8783.1+53.803 T- 0.11961 T + 0.00011238 T - 3.628899 10 T 2 Ni/NiO ! ! ! " ! (4)

0.0

5.0

10.0

15.0

20.0

25.0

30.0

520 540 560 580 600 620 640

H concentration at the Ni/NiO transition

(cm3/kg)

T (K)

y = - 8783.1 + 53.803x - 0.11961x2 +

 + 0.00011238x3 - 3.6288E-8x4

Figure 11. 

Hydrogen concentration at Ni/NiO phase transition 

as a function of temperature (data taken from 

Ref. 36). 

Equation 3 can be solved for the hydrogen concentration at test temperature T2. This concentration 

maintains the potential difference to the Ni/NiO line similar to that of a test conducted at temperature T1. 

For the purpose of this calculation, test conditions “1” are considered those typical of tests conducted in a 

PWR environment at 320°C (T1 = 593 K, [H2]test1 = 23 cm3/kg). The results for several temperatures of 

interest are shown in Table 3. For each temperature, the hydrogen concentration at the Ni/NiO transition 

is calculated by Eq. 4, and then the concentration needed to maintain the corrosion potential difference 

constant is calculated by Eq. 3. 

Table 3. Test conditions for various temperature and hydrogen tank overpressures. 

Test Temperature 

(°C) (K)

[H2]Ni/NiO 

(cm3/kg) 

[H2]test 

(cm3/kg) 

!E 

(mV) 

290 563 3.0 11.01 31.5 

300 573 3.9 13.81 31.3 

305 578 4.5 15.60 31.1 

310 583 5.2 17.71 30.8 

320 593 7.1 23.00 30.1 

350 623 17.8 47.59 26.4 
 

13 

2.3 Test Procedure 

The CGR tests were conducted in the load-control mode using a triangular, sawtooth, or trapezoidal 

waveform with load ratio R of 0.3-0.7. The CT specimens were fatigue precracked with a triangular 

waveform in the test environment at load ratio R = 0.3, frequency !1 Hz, and maximum stress intensity 

factor Kmax of 20-25 MPa m1/2 (18.2-22.8 ksi in1/2). After !0.5-mm (20 mils) extension, R was 

increased incrementally to 0.5-0.7, and the loading waveform changed to a slow/fast sawtooth with rise 

times of 30-1000 s, and unload time of 12 s This loading sequence is considered to result in reproducible 

CGRs.38 For sawtooth loading, the CGRs were calculated on the basis of the rise time. During 

individual test periods, Kmax was maintained approximately constant by periodic load shedding (less than 

2% decrease in load at any given time). 

Crack extensions were monitored by the reversing-DC potential difference method. The current 

leads were attached to the holes on the top and bottom surfaces of the specimen (Fig. 1), and potential 

leads were either welded on the front face of the specimen across the machined notch but on diagonal 

ends (1T-CT specimens, Fig. 1a) or attached to pre-machined taps (1/2T-CT specimens, Fig. 1b). Also, 

to compensate for the effects of changes in resistivity of the material with time, an Alloy 182 internal 

reference bar was installed near the test specimen. The CT specimen and reference bar were connected in 

series, and the DC potential across the specimen as well as the reference bar was monitored continuously 

during the test. The results for the reference bar were used to normalize potential drop measurements for 

the CT test specimen. 

During crack growth tests in high-temperature water, environmental enhancement of CGRs does not 

occur from the start of the test. Under more rapid cyclic loading, the crack growth is dominated by 

mechanical fatigue. The CGRs during precracking and initial periods of cyclic loading were primarily 

due to mechanical fatigue. In general, environmental enhancement is typically observed under loading 

conditions that would lead to CGRs between 10-10 and 10-9 m/s in air. 

The stress intensity factor range !K was calculated in accordance with ASTM E 1681 and E 647 as 

follows: 

( )1/ 2 3/ 2

N

a 2

P a W K f

BB W a W

1

W

! " # $ +

% & ' ! " % = # $ ! " & ' # $ ( & '

(5)

!P = Pmax " Pmin for R > 0 (6) 

2 3 4

a a a a a f 0.886 4.64 13.32 14.72 5.6

W W W W W

! " ! " ! " ! " ! " $ % = + $ % $ % $ % $ % # + # & ' & ' & ' & ' & ' , (7) 

where Pmax and Pmin are maximum and minimum applied load, a is crack length, W is the specimen 

width, and effective thickness Beff = (B BN)0.5. The applied K and crack size limits for the tests met the 

ASTM E 1681 and E 647 criteria. These criteria are intended to ensure applicability and transferability of 

the cracking behavior of a component or specimen of a given thickness under a specific loading condition 

to a crack associated with a different geometry, thickness, and loading condition. The K/size criteria 
 

14 

require that the plastic zone at the tip of a crack is small relative to the specimen geometry. For constant 

load tests, ASTM E 1681 requires that 

Beff and (W – a) !2.5 (K/!ys)2, (8) 

and for cyclic loading ASTM 647 requires that 

(W – a) !(4/") (K/!ys)2, (9) 

where K is the applied stress intensity factor, and !ys is the yield stress of the material. In 

high-temperature water, because the primary mechanism for crack growth during continuous cycling is 

not mechanical fatigue, Eq. 8 is probably the more appropriate criterion, but Eq. 9 may give acceptable 

results. For high-strain hardening materials (i.e., ultimate-to-yield stress ratio !ult/!ys !1.3), both criteria 

allow the use of the flow stress defined as !f = (!ult + !ys)/2 rather than the yield stress. 

In an earlier report,33 experimental CGR data were obtained at ANL on field Alloy 600 and weld 

Alloys 82 and 182. In that report, K values were calculated using the correlations for a disc-shaped 

specimen instead of a standard CT specimen. The earlier data have been corrected using Eqs. 5-7; the 

corrected data are given in Appendix A of this report. The difference between the K values based on the 

correlations for a disc-shaped specimen and standard CT specimen is <5%. 

After the test the specimen was fractured in liquid nitrogen, and the fracture surfaces were 

examined by optical or electron microscopy to measure the final crack length using the 9/8 averaging 

technique; that is, the two near-surface measurements were averaged, and the resultant value was 

averaged with the remaining seven measurements. 

2.4 Analysis of Crack Growth Rate Data 

Under cyclic loading, the CGR (m/s) can be expressed as the superposition of the rate in air 

(i.e., mechanical fatigue) and the rates due to corrosion fatigue and SCC ( CF a! and aSCC ! , respectively), 

given as 

a a a a ! ! ! ! env air CF SCC = + + . (10) 

The cyclic CGRs for Ni alloys and welds in air were determined from the correlations developed earlier at 

ANL:31,33 

( ) ( ) 2.2 4.1

air r r

da a / t C 1 0.82 R K / t

dN

" # $ % ! = = ( ) & ! & & ' * + , - . / ! , (11) 

where da/dN is the growth rate per cycle, tr is the rise time for the loading cycle, R is the load ratio (i.e., 

ratio of the minimum and maximum stress intensity factors Kmin/Kmax), #K is Kmax - Kmin in MPa m1/2, 

and the constant C depends on the material and temperature. For Alloy 600 the constant (CA600) is a 

third-order polynomial with respect to temperature T (°C), 

-14 -17 -18 2 -21 3 C = 4.835 10 - (1.622 10 )T + (1.490 10 )T - (4.355 10 )T A600 ! ! ! ! . (12) 
 

15 

For Ni-alloy welds, the constant (CNiweld) is given by a fourth-order polynomial with respect to 

temperature T (°C), expressed as 

-14 -17 -18 2

Ni-weld

-20 3 -23 4

C = 8.659 10 - (5.272 10 )T + (2.129 10 )T 

 - (1.965 10 )T + (6.038 10 )T

! ! !

! !

(13)

In LWR coolant environments, the CGRs of Alloy 600 show frequency-dependent enhancement 

under cyclic loading conditions. In high-DO water [i.e., normal water chemistry (NWC) boiling water 

reactor (BWR) water], the environmental enhancement of the growth rates does not appear to depend on 

the material composition (e.g., C content) or material heat treatment. In contrast, environmental 

enhancement of CGRs of Alloy 600 in low-DO water does seem to be strongly dependent on material 

conditions. In the literature,39-43 such variability has been attributed to thermomechanically controlled 

parameters, such as yield strength and grain boundary coverage of carbides, although the evidence for this 

dependence is more substantial for steam generator tubing than for structural components. 

In earlier ANL work, correlations were developed to estimate the enhancement of cyclic CGRs in 

LWR environments relative to the CGRs in air under the same loading conditions. In the absence of any 

significant contribution of SCC to growth rate, the cyclic CGRs for Alloy 600, either in the solution 

annealed (SA) condition or the SA plus thermally treated condition, in !300 ppb DO water at 289°C is 

given by the expression28

( )0.33 7

env,A600 air,A600 air,A600 a a 4.4 10 a ! ! ! = + " # ! . (14)

In low-DO environments (e.g., hydrogen water chemistry BWR or PWR environment) at 320°C (608°F), 

some alloys show little enhancement, while others show enhancement comparable to that predicted by 

Eq. 14. 

Similarly, correlations describing the cyclic CGRs of Ni-alloy welds (e.g., Alloys 182 and 82) have 

been developed.33 Under similar loading conditions, the CGRs of Ni-alloy welds are a factor of 2-3 

higher than those of Alloy 600. The analysis33 indicated that the cyclic CGRs of Ni-alloy welds in PWR 

water may either be bounded by 5 ! CGRair or be represented by the expression 

( )0.78

env,Ni weld air,Ni weld air,Ni weld a a 0.018 a ! = + ! ! ! ! " ! . (15)

The SCC growth rate data for Alloy 600 and its weld metals have been reviewed by White and 

Hickling44,45 to determine the effects of critical parameters such as stress intensity factor, temperature, 

material heat treatment, cold work, and water chemistry on growth rates. For Alloy 600, the CGR (m/s) 

under SCC conditions is represented by the expression,44

A600 th

ref

Q 1 1 a exp (K K ) R T T

! " # $ % = & ' ' ' ( ) * + . / , -

! , (16) 

where: Q = activation energy for crack growth, 

 = 130 kJ/mol (31.1 kcal/mol) for Alloy 600, 

R = universal gas constant, 

 = 8.314 " 10-3 kJ/mol K (1.103 " 10-3 kcal/mol°R), 
 

16 

T = absolute operating temperature in K (or °R), 

Tref = absolute reference temperature used to normalize the CGR data, 

 = 598 K (1076.67°R), 

! = crack growth amplitude (2.67 ! 10-12 at 325°C), 

K = crack tip stress intensity factor (MPa m1/2), 

Kth = crack tip stress intensity factor threshold (9 MPa m1/2), and 

" = exponent 1.16. 

The effect of K on the SCC growth rate for Ni-alloy welds in PWR environments has been 

represented by a modified45 version of the above relationship. Unlike the CGR relationship for 

Alloy 600, the relationship for Ni-alloy welds has no threshold value for the stress intensity factor K, 

Ni weld

ref

Q 1 1 a exp K

R T T

! "

# $ % & = ' " " ( ) * + . / , -

! , (17) 

where Q, R, T, and Tref are the same as in Eq. 16; crack growth amplitude ! is 1.5 ! 10-12 at 325°C; and 

exponent " is 1.6. Also, unlike Alloy 600 for which a reliable value for the activation energy for crack 

growth has been determined, no such number is available for the Ni-weld alloys. Thus, for Ni-weld 

alloys, the activation energy is assumed to be the same as that for Alloy 600. 
 

17 

3 Microstructural Characterization 

The alloys used in the present project have been subject to an extensive metallograpic analysis. 

The results were presented in detail in previous reports.32,33 This section summarizes the most important 

findings for laboratory-prepared alloys and for field components from the Davis-Besse and V. C. Summer 

reactors. 

3.1 Laboratory-Prepared Alloys 

A metallographic examination was performed to characterize the microstructure of the laboratoryprepared welds by scanning electron microscopy (SEM), energy dispersive X-ray (EDX) spectroscopy, 

and OIM. For this purpose, three rectangular pieces (1 cm ! 2 cm ! 0.5 cm) were cut from the double-J 

weld. All samples were mechanically polished through 1-μm diamond paste, followed by electrolytic 

etching for 10-20 s in a 70% H3PO4 and water solution at 5 V at room temperature to reveal the dendritic 

structure and carbides. 

Scanning electron microscopy revealed a microstructure with elongated grains and dendritic 

features typical of weld alloys (Fig. 12). A large variation in grain size between the different weld passes 

was also observed. In addition, high angle boundaries were observed to extend from the weld into the 

weld heat-affected zone (HAZ). The implication of this observation is that once a crack initiates in the 

HAZ, it can then readily extend into the weld or vice-versa via such continuous, cracking-susceptible 

boundaries. 

Figure 12. 

Dendritic microstructure on a laboratory-preapred 

Alloy 182 weld specimen. 

Both matrix and grain boundary precipitates were also investigated. Figure 13a shows an example 

of one such matrix precipitate, and Fig. 13b shows the EDX spectra resulting from the matrix and the 

precipitate. The comparison of the two spectra indicates that the Ti and O peaks are higher in the 

spectrum resulting from the precipitate, suggesting that the particle composition is TiO2. 

(a) (b) 

Figure 13. (a) Micrograph showing one of the matrix precipitates observed on the surface of sample 3A 

and (b) EDX spectra resulting from the bulk and the precipitate shown in (a). 
 

18 

The deep-groove weld was also examined by OIM to determine primarily the grain-boundary 

character distribution and to check for any texture in the microstructure. Both these parameters are 

known to influence the SCC behavior of austenitic alloys in high-temperature water environments.46-52 

The OIM analysis allows a classification of boundaries according to the coincident site lattice model as 

either coincident site lattice boundaries (CSLBs) or high-angle boundaries (HABs). The CSLBs are 

formed when the neighboring grains are in specific orientation relationships and have been shown to 

possess an increased resistance to SCC over HABs.46-52 In addition, texture affects the high-temperature 

deformation behavior of a polycrystalline material and thus is expected to play a role in the SCC behavior 

as well. 

The OIM maps obtained from a plane parallel to the direction of dendrites are shown in Figs. 14a 

and b. The microstructure consists of columnar grains, typical of those observed on this plane. Grains 

with similar orientations tend to cluster (the different orientations are shown with different colors). The 

boundary character distributions resulting from the two scans (Figs. 14b and d) show CSL fractions of 

27.5% and 22.7%, much lower numbers than those typically found for solution-annealed Alloy 600. 

(a) (b) 

Figure 14. OIM maps from a plane parallel to the direction of dendrites showing the grain orientations 

(a, c). 

Figure 15 shows the grain-boundary character distribution (GBCD) of the weld alloys. This 

distribution is dominated by low-angle boundaries (LABs, !1") and is consistent with the clustering of 

grains of similar orientations. The higher the proportion of neighboring grains that share some particular 

orientation, the higher becomes the probability that these grains have parallel crystal structures. The fact 

that the GBCD is dominated by !1 boundaries also makes the weld metal significantly different from 

Alloy 600, in which the GBCD is dominated by twin-related !3, !9, and !27 boundaries (of which 

approximately half are coherent).52

 

"In the CSL model the ! number is the reciprocal of the density of coincident sites. 
 

19 

0

0.05

0.1

0.15

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

Scan 1 Deep Groove Weld

Scan 2 Deep Groove Weld

Scan 3 Double J Weld

Fraction of CSL boundaries

CSL ! number

Figure 15. 

Grain boundary character 

distribution for laboratoryprepared welds. 

The OIM analysis presented in this section shows that the intragranular dendrites are coherent and 

are, therefore, expected to be resistant to cracking. In consequence, it appears that the grain boundary 

character distribution (along with residual deformation) will determine the cracking behavior of the 

welds. The proportion of cracking-resistant CSL boundaries in the weld was found to be relatively small; 

however, the grains having similar orientations were observed to cluster. 

3.2 Field Alloys 

Microstructural characterization of material from the Alloy 600 nozzle and J-groove weld from 

Davis-Besse was conducted on the planes of interest for crack formation and growth in cylindrical-tensile 

and CT specimens.33 The specimens were mechanically polished with 1-μm diamond paste. The Alloy 

600 specimens were etched with Marble's reagent (hydrochloric acid, copper sulfate, and distilled water). 

The weld specimens were electrochemically etched in a 5% nital solution at 5 V at room temperature to 

show grain boundaries, followed by etching in 70% phosphoric acid solution at 5 V at room temperature 

to show the interdendritic microstructure. 

Figure 16 shows the microstructure observed on the Alloy 600 nozzle specimens. The grain size 

varies significantly, ranging from 30 to 200 μm, with an average size of !75 μm (ASTM grain size 4). 

Microhardness measurements indicated that the hardness was uniform across the nozzle wall. There was 

extensive grain boundary coverage (GBC) by Cr-rich carbides, and a few carbides were randomly 

distributed in the matrix. Although in most cases the boundaries were decorated with carbides, in some 

cases the carbides appeared to be present only on one side of the boundary (e.g., the boundary going from 

top left to bottom right in Fig. 16b). The average size of the grain-boundary carbides was !0.3 μm, and 

the GBC was estimated to be in the range 50-60%. The specimens were also examined by EDX 

spectroscopy to investigate the precipitates in the matrix and along the grain boundaries. The large 

precipitates in the grain matrix and boundaries were Ti-rich. The effect of Ti-rich precipitates on the SCC 

behavior is unknown. The precipitation of Ti carbides can reduce the concentration of free carbon 

retained in the matrix to a level which limits the precipitation of the desirable grain-boundary chromium 

carbides. If Ti carbides precipitate at grain boundaries, they may oxidize when exposed to the high 

temperature water environment, possibly leading to grain boundary embrittlement. However, crack tip 

examinations by TEM have found no evidence of Ti oxides,53 suggesting that these particles do not 

participate directly in the SCC process. 
 

20 

(a) (b) 

Figure 16. Microstructures observed on planes (a) 03-1 and (b) 03-2 of the CRDM nozzle alloy. 

Figure 17 shows the microstructures observed on the specimens cut from the Davis-Besse 

Alloy 182 J-groove weld. These micrographs show the interdendritic microstructure typical of welds and 

are consistent with the microstructure expected based the plane of observation. 

Figure 17. 

Microstructure observed on the Alloy 182 J-groove 

weld specimen from Davis-Besse. 

As with the material from Davis-Besse, the metallography for the weld alloys from the 

V. C. Summer power plant was also focused on the planes of interest for crack formation and growth in 

the cylindrical-tensile and CT specimens. Both the Alloy 182 butter and the Alloy 82 weld were 

examined. The samples were mechanically polished with 1-μm diamond paste, and electrochemically 

etched in a 70% phosphoric acid solution at 5 V. 

Figure 18 shows a micrograph of the Alloy 182 butter microstructure. The micrograph shows 

interdendritic microstructure consistent with the plane of observation: the elongated grains (and dendrites) 

are oriented primarily in a horizontal direction. 
 

21 

Figure 18. 

Microstructure of the butter Alloy 182. 

Figure 19 shows the microstructure of the Alloy 82 weld. This typical weld microstructure consists 

of dendrites (and columnar grains) oriented primarily in the vertical direction in the micrograph. 

Figure 19. 

Microstructure of the weld Alloy 82 from the 

V.C. Summer plant. 

Of possible significance to the SCC behavior of this weld was the observation of grain boundary 

sliding at the interface between the Alloy 82 weld metal and the SS pipe. Several examples are indicated 

with white arrows in Fig. 20. Sliding occurred exclusively at grain boundaries (most likely high angle) 

extending from the pipe material to the weld material. Most likely, the grain boundary sliding occurred 

during the repair of this weld due to high residual stresses and local heating. Because grain boundary 

sliding is a precursor to intergranular SCC initiation,54 these observations suggest that the interface may 

be quite susceptible to SCC. Nevertheless, the picture shown in Fig. 20 was taken from a location away 

from where the samples for SCC testing were cut; thus, the observed sliding is unlikely to have any 

bearing on the SCC results presented in this report. 

Figure 20. 

Grain boundary sliding at the weld-pipe interface. 

 

23 

4 Crack Growth Test Results 

This section focuses on the crack growth data obtained from the three Alloy 182 specimens 

described in Section 2.1. The objective of these tests was to determine the activation energy in Alloy 182 

weld. The first specimen (A182-1) was a 1-T CT specimen cut from a laboratory-prepared double-J weld, 

while the other two (CT933H-1 and CT933H-2) were companion 1/2-T CT specimens machined from a 

laboratory-prepared deep-groove weld. 

4.1 Double-J Weld Specimen A182-1 

This test was conducted with the A182-1 specimen in the temperature range 290-320°C 

(554-608°F). The test was started at 320°C. After a steady-state CGR was obtained at 320°C, the 

temperature was reduced to 305°C and then to 290°C. 

The specimen was fatigue precracked at Kmax = 20 MPa m1/2 (18.2 ksi in1/2) R = 0.3, and a 

triangular waveform, and then cycled with a sawtooth waveform at R = 0.5 with increasing rise times up 

to 1000 s. The specimen was next set to constant load. After approximately 600 h and 0.270 mm (10 

mils) of growth, the temperature was lowered to 305°C. The hydrogen gas concentration was lowered 

from 23 to 15 cc/kg by adjusting the hydrogen gas overpressure on the water supply tank. After 

approximately 1300 h and 0.1 mm (4 mils) of growth, the temperature was lowered to 290°C, and the 

hydrogen concentration was lowered to 11 cc/kg. After another approximately 1200 h and 0.06 mm (2.3 

mils) of growth, the temperature was increased back to 320°C. The hydrogen concentration was also 

increased to 23 cc/kg. 

The entire fracture surface of specimen A182-1 is shown in Fig. 21. Figure 21a shows the optical 

image, while Fig. 21b shows the SEM image. Both images show a relatively uniform crack surface, with 

an intergranular fracture mode in the second half of the test run. The fracture surface images were used 

for correcting the DC potential data. The correction factors were 1.29 for the transgranular region and 2.6 

for the intergranular region. The resulting cyclic and constant load data obtained on this specimen are 

summarized in Table 4, and the changes in crack length and Kmax with time are shown in Fig. 22. 

Table 4. Test conditions and crack growth data for specimen A182-1 of Alloy 182 SMA weld in PWR 

water.a

Test 

Test 

Time Temp. 

H 

Conc. 

O2 

Conc.b

Load 

Ratio 

Rise 

Time 

Down 

Time Kmax !K CGRenv

Estimated 

CGRair

Crack 

Length 

Period (h) (°C) (cc/kg) (ppb) R (s) (s) (MPa m1/2) (MPa m1/2) (m/s) (m/s) (mm) 

Pre a 103 321.8 23 <10 0.30 0.5 0.5 20.8 14.5 1.48E-07 6.05E-08 12.820 

Pre b 126 321.3 23 <10 0.30 5 5 22.4 15.6 3.54E-08 8.15E-09 14.253 

Pre c 142 321.2 23 <10 0.30 100 100 22.4 15.7 3.93E-09 4.14E-10 14.354 

Pre d 148 321.8 23 <10 0.3 5 5 23.2 16.3 4.60E-08 9.54E-09 15.027 

Pre e 166 321.6 23 <10 0.3 100 100 23.3 16.3 3.95E-09 4.79E-10 15.073 

Pre f 173 322.2 23 <10 0.3 5 5 24.1 16.9 5.37E-08 1.11E-08 15.828 

Pre g 266 321.1 23 <10 0.3 100 100 25.2 17.6 4.24E-09 6.64E-10 16.574 

1 365 321.8 23 <10 0.5 300 12 29.0 14.5 2.59E-09 1.71E-10 17.571 

2 599 320.9 23 <10 0.5 600 12 30.7 15.3 1.43E-09 1.07E-10 18.772 

3 867 320.8 23 <10 0.5 1000 12 32.1 16.0 1.04E-09 7.79E-11 19.705 

4 1464 322.1 23 <10 1.0 - - 33.7 0.0 5.88E-10 - 20.696 

5 2776 307.3 15 <10 1.0 - - 34.5 0.0 1.23E-10 - 21.196 

6 3,982 291.1 11 <10 1.0 - - 35.2 0.0 3.63E-11 - 21.552 

7 4,819 321.4 23 <10 1.0 - - 37.8 0.0 4.95E-10 - 23.057 aSimulated PWR water with 2 ppm Li, 1100 ppm B, and 2 ppm. Conductivity was 21 ± 3 μS/cm, and pH was 6.4. bRepresents values in the effluent. 
 

24 

(a) 

(b) 

Figure 21. Fracture surface of specimen A182-1: (a) optical image and (b) SEM image. 
 

25 

11.2

12.0

12.8

13.6

14.4

15.2

16.0

16.8

17.6

10

15

20

25

30

35

40

45

50

120 160 200 240 280

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

Kmax

Crack Length

Alloy 182 SMA Double–J Weld

Specimen # A182–1

320°C, Simulated PWR Water

4.24 x 10–09 m/s

25.2 MPa m0.5

0.30, Rise time = 100 s

Pre

a

Pre

b

Pre

c

Pre

d

Pre

e

Pre

f

Pre

g

(a) 

13.0

14.0

15.0

16.0

17.0

18.0

19.0

20.0

10

15

20

25

30

35

40

45

50

300 350 400 450 500 550 600 650

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

Kmax

Crack Length

Period 1

2.59 x 10–9 m/s

29.0 MPa m0.5

0.50, Rise time = 300 s

Period 2

1.43 x 10–9 m/s

30.7 MPa m0.5

0.50, Rise time = 600 s

Alloy 182 SMA Double–J Weld

Specimen # A182–1

320°C, Simulated PWR Water

(b) 

15.00

16.00

17.00

18.00

19.00

20.00

21.00

22.00

10

15

20

25

30

35

40

45

50

600 650 700 750 800 850

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

Kmax

Crack Length

Period 3

1.04 x 10–9 m/s

32.1 MPa m0.5

0.50, Rise time = 1000 s

Alloy 182 SMA Double–J Weld

Specimen # A182–1

320°C, Simulated PWR Water

(c) 

Figure 22. Crack length vs. time for Alloy 182 weld specimen A182-1 in simulated PWR environment 

during test periods (a) precracking, (b) 1-2, (c) 3, and (d) 4-7. 
 

26 

16.00

17.00

18.00

19.00

20.00

21.00

22.00

23.00

24.00

10

20

30

40

50

60

70

1000 1500 2000 2500 3000 3500 4000 4500 5000

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

Kmax

Alloy 182 SMA Double–J Weld

Specimen # A182–1

Simulated PWR Water

322°C

5.9 x 10–10 m/s

33.7 MPa m0.5

Constant load temperature 305°C,

re-initialized

temperature 290°C,

re-initialized

temperature 320°C,

re-initialized

307°C

1.2 x 10–10 m/s

34.5 MPa m0.5

Constant load

291°C

3.6 x 10–11 m/s

35.2 MPa m0.5

Constant load

Crack Length

321°C

4.9 x 10–10 m/s

37.8 MPa m0.5

Constant load

(d) 

Figure 22. (Cont.) 

4.2 Deep-Groove Weld Specimen CT933H-1 

This test was conducted with the CT933H-1 specimen in the temperature range 290-310°C 

(554-590°F). The test was started at 310°C. After a steady-state CGR was obtained at 310°C, the 

temperature was reduced to 300°C and then to 290°C. At the end of the test, the temperature was 

increased back to 310°C. 

The CT933H-1 specimen was fatigue precracked in the environment at 310°C with 

Kmax = 20 MPa m1/2 (18.2 ksi in1/2), R = 0.3, and a triangular waveform, and then cycled with a sawtooth 

waveform at R = 0.5 with increasing rise times up to 500 s. The specimen was then set at constant load. 

After approximately 1900 h, the temperature was decreased to 300°C, and the supply tank hydrogen 

concentration was lowered from 18 to 14 cc/kg by adjusting the hydrogen gas overpressure on the water 

supply tank. After another 1600 h the temperature was further decreased to 290°C, and the water 

hydrogen concentration was lowered to 11 cc/kg. After approximately 700 h at 290°C the temperature 

was increased back to 310°C (hydrogen concentration 18 cc/kg) for the final period. 

The entire fracture surface of the CT933H-1 specimen is shown in Fig. 23. This picture was used for 

data correction. The correction factors were 1.15 for the transgranular region and 4.3 for the intergranular 

region. The testing conditions are shown in Table 5, and the changes in crack length and Kmax with time 

are shown in Fig. 24. 

Figure 23. 

Fracture surface of Alloy 182 

weld specimen CT933H-1 tested 

in simulated PWR environment 

at temperatures of 290-310°C 

(554-590°F). 
 

27 

Table 5. Test conditions and crack growth data for specimen CT933H-1 of Alloy 182 SMA weld in 

PWR water.a

Test 

Test 

Time Temp. 

H 

Conc. 

O2 

Conc.b

Load 

Ratio 

Rise 

Time 

Down 

Time Kmax !K CGRenv

Estimated 

CGRair

Crack 

Length 

Period (h) (°C) (cc/kg) (ppb) R (s) (s) (MPa m1/2) (MPa m1/2) (m/s) (m/s) (mm) 

Pre a 51 310.8 18 <10 0.29 1.00 1.00 19.6 13.9 2.60E-08 2.22E-08 12.404 

Pre b 66 311.1 18 <10 0.30 10 10.0 12.3 8.6 1.49E-09 3.20E-10 12.412 

Pre c 69 311.1 18 <10 0.32 1.0 1.0 18.2 12.4 2.36E-08 1.47E-08 12.581 

1 96 311.3 18 <10 0.16 50.0 12.0 18.7 15.7 1.85E-09 5.51E-10 12.636 

2 146 310.5 18 <10 0.15 500.0 12.0 23.6 20.1 5.21E-10 1.45E-10 12.720 

3 718 310.4 18 <10 0.50 500.0 12.0 24.8 12.4 1.27E-10 4.84E-11 13.084 

4 2685 310.2 18 <10 1.00 0.0 - 26.0 0.0 5.88E-11 - 13.418 

5 4294 301.2 14 <10 1.00 0.0 - 26.7 0.0 1.25E-11 - 13.493 

6 4,968 292.1 11 <10 1.00 0.0 - 26.8 0.0 2.68E-11 - 13.541 

7 6,870 310.7 18 <10 1.00 0.0 - 27.1 0.0 3.62E-11 - 13.689 

8 6,958 310.7 18 <10 0.50 500.0 12.0 27.2 13.6 1.30E-10 7.02E-11 13.773 aSimulated PWR water with 2 ppm Li, 1100 ppm B, and 2 ppm. Conductivity was 21 ± 3 μS/cm, and pH was 6.4. 

bRepresents values in the effluent. 

11.50

12.00

12.50

13.00

13.50

10

20

30

40

50

60

60 80 100 120 140 160

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

Kmax

Crack Length

Alloy 182 Weld

Specimen # CT933H-1

310°C, Simulated PWR Water

Pre

a

Period 2

5.21 x 10–10 m/s

23.6 MPa m0.5

R = 0.16, Rise 500 s

Pre

b

Pre

c

Period 1

(a) 

12.40

12.80

13.20

13.60

14.00

10

20

30

40

50

60

500 1000 1500 2000 2500

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

Kmax

Crack Length

temperature 300°C,

re-initialized

Alloy 182 Weld

Specimen # CT933H-1

Simulated PWR Water

Period 3

1.27 x 10–10 m/s

24.8 MPa m0.5

R = 0.50, Rise 500 s

Period 4

310°C

5.88 x 10–11 m/s

26.0 MPa m0.5

Constant load

(b) 

Figure 24. Crack length vs. time for Alloy 182 weld specimen CT933H-1 in simulated PWR environment 

during test periods (a) precracking-2, (b) 3-4, (c) 5-6, and (d) 7-8. 
 

28 

12.60

12.80

13.00

13.20

13.40

13.60

13.80

14.00

14.20

10

20

30

40

50

60

70

80

3000 3500 4000 4500 5000

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

Kmax

Crack Length

temperature 300°C,

re-initialized

Period 5

300°C

1.25x 10–11 m/s

26.7 MPa m0.5

Constant load

temperature 310°C,

re-initialized

temperature 290°C,

re-initialized

Alloy 182 Weld

Specimen # CT933H-1

Simulated PWR Water Period 6

290°C

2.68 x 10–11 m/s

26.8 MPa m0.5

Constant load

(c) 

12.60

12.80

13.00

13.20

13.40

13.60

13.80

14.00

14.20

10

20

30

40

50

60

70

80

5000 5500 6000 6500 7000

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

Kmax

Crack Length

temperature 310°C,

re-initialized

Period 8 Alloy 182 Weld

Specimen # CT933H-1

Simulated PWR Water

Period 7

310°C

3.62 x 10–11 m/s

27.1 MPa m0.5

Constant load

(d) 

Figure 24. (Cont.) 

4.3 Deep-Groove Weld Specimen CT933H-2 

This test was conducted with the CT933H-2 specimen in the temperature range 290-350°C 

(554-662°F). The test on CT933H-2 (companion of CT933H-1) was started at 350°C. After a steadystate CGR was obtained at 350°C, the temperature was reduced to 320°C and then to 290°C. Finally, the 

temperature was increased back to 350°C. 

The specimen CT933H-2 was fatigue precracked in the environment at 350°C with 

Kmax = 25 MPa m1/2 (22.8 ksi in1/2), R = 0.45, and a triangular waveform, and then cycled at R = 0.60 

with increasing rise times up to 500 s. The specimen was next set at constant load. The specimen was 

further cycled with increasing rise times, 500 and 1000 s, then set at constant load again. After the 

sequence was repeated three times, yielding three CGR measurements at 350°C, the temperature was 

decreased to 320°C. The hydrogen concentration in water was decreased from 48 to 23 cc/kg by lowering 

the hydrogen gas overpressure on the supply tank. Next, the temperature was further reduced to 290°C, 

and the hydrogen concentration was further reduced to 11 cc/kg. After approximately 800 h at 290°C the 

temperature was increased back to 350°C (hydrogen concentration of 48 cc/kg) for the final period. 
 

29 

The entire fracture surface of the CT933H-2 specimen is shown in Fig. 25. The image of Fig. 25 

was used for data correction. The correction factors were 1.17 for the transgranular region and 3 for the 

intergranular region. The testing conditions are shown in Table 6, and the changes in crack length and 

Kmax with time are shown in Fig. 26. 

Figure 25. Fracture surface of Alloy 182 weld specimen CT933H-2 tested in 

simulated PWR environment at temperatures of 290-350°C. 

Table 6. Test conditions and crack growth data for specimen CT933H-2 of Alloy 182 SMA weld in 

PWR water.a

Test 

Test 

Time Temp. 

H 

Conc. 

O2 

Conc.b

Load 

Ratio 

Rise 

Time 

Down 

Time Kmax !K CGRenv

Estimated 

CGRair

Crack 

Length 

Period (h) (°C) (cc/kg) (ppb) R (s) (s) (MPa m1/2) (MPa m1/2) (m/s) (m/s) (mm) 

1 49 351 47 <10 0.45 0.50 0 18.8 13.5 5.10E-08 6.06E-08 12.059 

2 62 351 47 <10 0.46 100 100 18.4 13.2 2.05E-09 5.57E-10 12.101 

3 69 351 47 <10 0.45 1.0 1 21.9 15.8 1.24E-07 1.15E-07 13.211 

4 97 351 47 <10 0.61 50.0 12 22.3 8.7 2.12E-09 2.60E-10 13.288 

5 269 351 47 <10 0.61 500.0 12 22.2 8.7 2.84E-10 2.55E-11 13.594 

6 929 351 47 <10 1.00 - - 23.4 0.0 6.49E-11 - 13.764 

7 936 351 47 <10 0.50 500.0 12 25.8 12.9 1.02E-09 9.00E-11 13.794 

8 1029 351 47 <10 0.50 1000.0 12 25.9 13.0 5.04E-10 4.60E-11 13.963 

9 1096 351 47 <10 1 - - 25.9 0.0 1.06E-10 - 14.009 

10 1200 351 47 <10 0.50 500.0 12 26.8 13.4 2.34E-10 1.05E-10 14.111 

11 1342 350 47 <10 0.50 1000.0 12 27.2 13.6 9.21E-11 5.58E-11 14.143 

12 1717 351 47 <10 1 - - 26.8 0.0 4.65E-11 - 14.177 

13 2428 320 23 <10 1 - - 27.0 0.0 1.36E-10 - 14.601 

14 3240 291 11 <10 1 - - 28.8 0.0 2.72E-11 - 14.683 

15 3402 352 47 <10 1 - - 29.3 0.0 4.53E-10 - 14.821 

16 3434 352 47 <10 0.50 500.0 12 30.0 15.0 6.50E-09 1.68E-10 15.727 

17 4,294 352 47 <10 1 - - 35.2 0.0 6.59E-11 - 15.771 

18 4,382 351 47 <10 0.50 500.0 12 35.5 17.7 3.84E-09 3.35E-10 16.036 aSimulated PWR water with 2 ppm Li, 1100 ppm B, and 2 ppm. Conductivity was 22 ± 3 μS/cm, and pH was 6.4. 

bRepresents values in the effluent. 
 

30 

12.00

12.50

13.00

13.50

14.00

10

15

20

25

30

35

50 100 150 200 250

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

Kmax

Crack Length Alloy 182

Specimen # CT933H-2

350°C, Simulated PWR Water

5

2.84 x 10–10 m/s

22.2 MPa m0.5

0.60, Rise time = 500 s

1 2

3 4

(a) 

13.40

13.60

13.80

14.00

14.20

14.40

10

20

30

40

50

60

1000 1200 1400 1600

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

Kmax

Crack Length

Alloy 182

Specimen # CT933H-2

350°C, Simulated PWR Water

12

4.65 x 10–11 m/s

26.8 MPa m0.5

Constant load

6

7

9

11

temperature 320°C,

re-initialized

8

10

(b) 

13.50

14.00

14.50

15.00

15.50

16.00

16.50

17.00

17.50

10

20

30

40

50

60

70

80

90

2000 2400 2800 3200 3600 4000 4400

Crack Length (mm)

Kmax (MPa m0.5)

Time (h)

Kmax

Crack Length

Alloy 182

Specimen # CT933H-2

Simulated PWR Water

13

1.36x 10–10 m/s

26.9 MPa m0.5

Constant load

temperature 300°C,

re-initialized

14

2.72 x 10–11 m/s

28.8 MPa m0.5

Constant load

15

16

17

6.69 x 10–11 m/s

35.2 MPa m0.5

Constant load

18

temperature 350°C,

re-initialized

(c) 

Figure 26. Crack length vs. time for Alloy 182 weld specimen CT933H-2 in simulated PWR environment 

during test periods (a) 1-5, (b) 6-12, and (c) 13-18. 
 

31 

5. Discussion 

The section summarizes the crack growth rate results for Alloy 600 (obtained previously in the 

program) and Ni-base alloy welds (obtained from present and past efforts). The materials tested include 

both laboratory-prepared alloys and field components from Davis-Besse and V. C. Summer reactors. The 

most recent SCC CGR data on Alloy 182 are analyzed to extract the activation energy for crack growth in 

this alloy. In addition, the cyclic CGRs are analyzed with a superposition model for the purpose of 

establishing the individual contributions of mechanical fatigue, corrosion fatigue, and SCC. 

5.1. SCC Growth Rates 

5.1.1 Alloy 600 

The experimental CGRs for Alloy 600 (Heat M3935) from the Davis-Besse CRDM nozzle under 

constant load, with or without partial unloading, are plotted as a function of applied Kmax in Fig. 27a. The 

cumulative distribution of the log-normal fit to the ordered median ranking of the constant ! (Eq. 16) for 

various data sets of Alloy 600 is shown in Fig. 27b. The experimental CGRs for Alloy 600 from the 

CRDM nozzle #3 correspond to the !95th percentile of the distribution fitted to the various data sets (i.e., 

Heat M3935 of Alloy 600 from the Davis-Besse CRDM nozzle #3 exhibits very high susceptibility to 

PWSCC). The median and 75th percentile CGR curves developed by the MRP for thick-wall Alloy 600 

are also included in the figure.44 The growth rates for the Davis-Besse nozzle Alloy 600 are a factor of 

! 4-8 higher than the median curve and are a factor of ! 2-4 higher than the 75th percentile curve, which 

has been proposed as a disposition curve. 

10-12

10-11

10-10

10-9

10 15 20 25 30 35 40 45 50

N3CC-02

N3CC-03

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

Alloy 600 (Heat M3935)

(Davis Besse CRDM Nozzle #3)

316°C PWR Water 

Median 

Alloy 600 at 316°C

Proposed Disposition Curve

75th Percentile

Alloy 600 at 316°C

1/4–T CT Specimen

1/2–T CT Specimen 0.0

0.2

0.4

0.6

0.8

1.0

10-13 10-12 10-11 10-10

Alloy 600 Heats

D–B Nozzle #3

Cumulative Distribution F

Constant ! at 325°C

75th Percentile

50th Percentile

25th Percentile

! = 1.44 x 10-12

! = 2.98 x 10-12

! = 6.96 x 10-13

(a) (b) 

Figure 27. (a) CGR data for Alloy 600 from Davis-Besse CRDM nozzle #3 in PWR water at 316°C 

(601°F) under constant load and (b) log-normal distribution of constant ! for 26 heats of 

Alloy 600. 
 

32 

Based on the observed microstructure and tensile strength of the nozzle Alloy 600,33 the 

experimental CGRs are somewhat surprising. Materials with good grain boundary carbide coverage 

(Fig. 16) and relatively low or average tensile yield stress are generally considered to have low 

susceptibility to SCC in PWR water.4 A unique feature of the fracture surface for the Davis-Besse nozzle 

alloy is the predominantly IG fracture even during mechanical fatigue loading (Fig. 28). 

The high susceptibility for the material to PWSCC cannot be reconciled in terms of the tensile 

strength or microstructure of the alloy. Additional CGR data and metallographic evaluation of Alloy 600 

from service components are needed to understand better the environmentally assisted cracking behavior 

of Alloy 600 in PWR environments. 

Figure 28. 

Transition from TG to IG fracture mode in 

specimen N3CL-1. The sample notch is 

at the top of the figure, and crack advance 

is from top to bottom. 

5.1.2 Ni-Alloy Welds 

Experimental SCC CGRs obtained from the ANL tests on laboratory-prepared Alloy 182 weld 

specimens32 are plotted as a function of stress intensity factor K in Fig. 29a and are compared with 

available CGR data for Ni-alloy welds59-68 in Fig. 29b. The double-J and deep groove welds are different 

heats of material as well as different weld geometries. The disposition CGR curve proposed by the White 

and Hickling45 for Alloy 182 in the PWR environment is also plotted in the figures. The disposition 

curve was developed for a temperature of 325°C (617°F), and the version of the disposition curve in 

Fig. 29a was normalized to 320°C (608°F) using Eq. 17 and an activation energy of 130 kJ/mol 

(31 kcal/mol). In Fig. 29b for comparison with the available SCC CGR data, the ANL test results were 

normalized to 325°C (640°F). Most of the data for Alloys 182 and 82 welds are between the median 

Alloy 600 curve and a factor of 10 above this curve, suggesting that the median curve for weld materials 

is about a factor of 3 greater than that for Alloy 600. 

In the ANL tests, CGRs were measured along the plane of the columnar grains/dendrites in a 

direction parallel to the columnar grains (TS or LS orientation) and perpendicular to the columnar grains 

(TL orientation). Note that the crack planes for TS and LS orientations are along the columnar dendrites, 

although they are perpendicular to each other. The results indicate that at high K values the CGRs for 

cracks propagating along the direction of dendrites (orientations TS and LS) are about a factor of two 
 

33 

greater than the CGRs for crack propagation across the direction of dendrites (orientation TL) (Fig. 29a). 

Similar results have also been obtained in other studies.59,60,63,64

10-12

10-11

10-10

10-9

10 20 30 40 50

CT31-W01 TS

CT31-W02 TS

CT933-01 TS

CT933-TL1

CT933-LS

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

Ni-Alloy Weld Metal

320°C

Proposed Disposition

Curve for Alloy 182 Weld

10-12

10-11

10-10

10-9

10 20 30 40 50

D545

PP751

33644

26B2

VCS182

VCS82

LM182

LM82

CT933-TS1

CT933-TL1

CT933-LS

CT31-W01

CT31-W02

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

Ni–Alloy Weld Metal

325°C

Open Symbols: Alloy 182

Closed Symbols: Alloy 82

Median 

Alloy 600

Proposed Disposition

Curve for Alloy 182 Weld

(a) (b) 

Figure 29. SCC crack growth data for the Argonne Alloy 182 welds (a) plotted as a function of K and 

(b) compared with the available data for Alloy 182 and 82 welds in simulated PWR 

environment. Data from Refs. 59-68. 

The experimental SCC growth rates for Alloy 182 from the Davis-Besse nozzle #11 J-groove weld 

under constant loading conditions (with or without periodic partial unloading) are shown in Fig. 30. The 

proposed disposition curve for Alloy 182 welds in a PWR environment45 is also included in the figure. 

The disposition curve was normalized to 316°C (600°F) using Eq. 16 and an activation energy of 

130 kJ/mol (31 kcal/mol). The CGRs for the Davis-Besse J-groove weld (in red in the figure) are an 

order of magnitude lower than the disposition curve proposed for Alloy 182 weld metals.45 

The experimental CGRs under constant load for the two V.C. Summer weld and butter alloys are 

compared with similar data obtained on laboratory-prepared welds in Fig. 31a. The CGRs for the 

V.C. Summer alloys are comparable to those determined for the Davis-Besse CRDM nozzle J-groove 

weld. The proposed disposition curve for Alloy 182 welds in a PWR environment45 is also included in 

the figure. The results indicate that the CGRs of Alloy 82 weld material (specimens WLR-01 and 

WCR-01) with growth direction along the dendrites are comparable to those of the Alloy 182 butter 

material (specimen BCR-01) with growth direction transverse to the dendrites. Note that the crack planes 

for LR and CR orientations are along the columnar dendrites, although they are perpendicular to each 

other (where L = longitudinal, C = circumferential, and R = radial). In general, CGRs in Alloy 182 are a 

factor of 2.6 higher than in Alloy 82; however, the rates along a direction transverse to the dendrites are a 

factor of !2 lower than those parallel to the dendrites.45 The effects of alloy type and specimen 

orientation seem to have cancelled each other, yielding approximately the same growth rates for the weld 

and butter alloys. 
 

34 

Figure 31b shows the V.C. Summer CGR data plotted along with data obtained by Jacko et al.67 on 

similar V.C. Summer weld and butter alloys in the PWR environment at 325°C (640°F); the ANL results 

were normalized to 325°C for comparison (Eq. 17). The CGRs obtained at ANL are a factor of !4 lower 

than those obtained by Jacko et al. The reasons for the differences in the results between the two studies 

are not clear. The crack fronts in the tests reported here are relatively straight. The crack fronts in the 

tests in Jacko et al.67 do not appear to be as straight. The CGRs in isolated “fingers” may be more rapid 

than the overall growth of a straight crack front. 

10-12

10-11

10-10

10-9

10 15 20 25 30 35 40 45 50

J11CC-3 (1/2-T CT)

J11CC-1 (1/4-T CT)

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

Alloy 182 J-Groove Weld

Davis Besse Nozzle #11

316°C

Proposed Disposition

Curve for Alloy 182 Weld

Specimen #

Figure 30. 

CGR data for Alloy 182 from the 

Davis-Besse CRDM nozzle J-groove 

weld in PWR environment at 316°C 

(600°F) under constant load. 

Figure 32 shows the cumulative distribution of the log-normal fit to the ordered median ranking of 

the constant ! in Eq. 17 for various data sets of Ni-alloy welds available in the literature; it essentially 

represents a means of ranking the various heats with respect to their susceptibility to SCC. However, the 

experimental CGRs were first adjusted for incomplete initiation of SCC, temperature, orientation, and 

alloy type, as follows: (a) divide by the engagement fraction to allow for incomplete initiation, (b) use 

Eq. 17 and an activation energy of 130 kJ/mol (31 kcal/mol) to normalize the CGR data to a common 

reference temperature of 325°C (640°F), (c) multiply the CGRs for TL orientations by 2 to account for 

orientation effects, and (d) multiply the CGRs for Alloy 82 welds by 2.6 to account for the effects of alloy 

type. Also, studies at CEA65 generally report only the maximum values of CGR; the average CGRs were 

determined from the maximum values by using the correlation proposed by Attanasio et al.62 The ratio 

between the maximum and average CGR was assumed to be the same as the ratio between the maximum 

and average crack extension (i.e., R = "amax/"aave). The ratio R is expressed as 

ln(R – 1) = 2.48 – 0.762 ln("aave/0.051), (18) 

where crack extension is in millimeters. The results of Jacko et al.67 suggest that the V. C. Summer weld 

and butter alloys correspond to the !95th and 80th percentile, respectively, of the distribution of the 

various heats, whereas the results from the present study indicate that they correspond to the !25th and 

10th percentile, respectively. 
 

35 

10-12

10-11

10-10

10-9

10 20 30 40 50

CT31-W01 TS

CT31-W02 TS

CT933-01 TS

CT933-TL1

CT933-LS

WLR-01

WCR-01

BCR-01

Experimental CGR (m/s)

Stress Intensity K (MPa·m1/2)

Ni-Alloy Weld Metal

320°C

Double–J Weld

Deep Groove Weld

V. C. Summer Weld & Butter

Proposed Disposition

Curve for Alloy 182 Weld

10-12

10-11

10-10

10-9

10 20 30 40 50

WCR-01

BCR-01

WLR-01

A 82 weld, Jacko et al.

A 182 butter, Jacko et al.

Experimental CGR (m/s)

Stress Intensity K (MPa·m1/2)

V.C. Summer Ni-Alloy Weld Metal

325°C

Proposed Disposition

Curve for Alloy 182 Weld

(a) (b) 

Figure 31. Comparison of the CGR data for the V.C. Summer weld and butter alloys with the data 

obtained on (a) laboratory-prepared welds in the present study and (b) V. C. Summer weld 

and butter alloys by Jacko et al. (Ref. 67). 

0.0

0.2

0.4

0.6

0.8

1.0

10-13 10-12 10-11

A 182

A 82

A 182 D-B 

A 82 VCS (ANL)

A 182 VCS (ANL)

A 82 (WH)

A 182 VCS (WH)

Cumulative Distribution F

Constant ! at 325°C

75th Percentile

25th Percentile

Median

Ni-Alloy Weld Metal

325°C

Figure 32. 

Log-normal distribution of constant ! for 

several heats of Alloy 82 and 182. 
 

36 

The ANL results indicate that the SCC growth rates for the field weld alloys from the Davis-Besse 

and V. C. Summer reactors are lower than the average CGRs for Alloy 182 welds. The results are not 

inconsistent with the heat-to-variability expected from laboratory tests. However, it is possible that other 

factors are at work. For example, field welds may be under a higher degree of structural constraint than 

laboratory-prepared welds. 

A common feature observed in all test specimens, with the exception of one V.C. Summer 

specimen, is the cracks on the fracture surface where the propagating SCC cracks seem to arrest. Figure 

33 illustrates this behavior with micrographs of the fracture surface from the Davis-Besse specimen 

J11CC-1 (Fig. 33a) and V. C. Summer specimen BCR-01 (Fig. 33b). The cracks occurred most likely 

during the laboratory CGR tests because they were not observed during the metallographic examinations. 

Furthermore, it can be hypothesized that the development of such cracks relieves the stress driving the 

propagation of the main crack. However, the cause for these cracks is unknown, although they could be 

associated with the greater constraint typically present in field welds, and determining their effect on the 

SCC behavior requires further investigation. 

(a) (b) 

Figure 33. Micrographs from the fracture surface of weld specimens: (a) Davis-Besse J11CC-1, and 

(b) V.C. Summer BCR-01. 

5.2. Activation Energy for SCC Crack Growth 

The available CGR vs. temperature data for Alloy 82 are plotted in Fig. 34. The different data sets 

yield activation energies that are either comparable to that observed for Alloy 600 (130 kJ/mol) (e.g., 

Bechtel Bettis59,60 and ETH61) or a factor of !2 higher (e.g., Lockheed Martin62). However, the 

corrosion potential for the Bechtel Bettis tests conducted at 360°C (680°F) with 150 cc/kg hydrogen was 

farther removed from the Ni/NiO phase transition than in the case of the 338°C (640°F) tests with 

40-60 cc/kg hydrogen. As such, since the CGRs are highest near the Ni/NiO phase transition and decrease 

as the potential deviates from this transition,36,37 the CGRs in the 360°C tests may have been reduced 

because of the high dissolved hydrogen levels. Thus, the actual activation energy for the Bechtel Bettis 

data may be somewhat higher than that determined in Fig. 34a. 

Figure 35 shows the Alloy 182 temperature dependence of the CGR data obtained at 

Westinghouse63,64 and CEA65. The resulting activation energies, 228 and 248 kJ/mol, are in very good 

agreement with each other. While the data are limited, they seem to suggest that the activation energy for 

the temperature dependence of CGRs in Alloys 182 welds may be higher than that in Alloy 600. 
 

37 

10-12

10-11

10-10

10-9

1.55 1.60 1.65 1.70 1.75 1.80

40-60 cc/kg H2

150 cc/kg H2

40-60 cc/kg H2

decreased by factor of 2

CGR (m/s)

1000/T (°K)

Bechtel Bettis

Alloy 82 GTA Weld 

PWR Water

333 315 298 283

Temperature (°C)

352

Kmax = 36 MPa m1/2

Q = 125 ± 10 kJ/mole 

 (29.8 ± 2.5 kcal/mole)

10-12

10-11

10-10

10-9

1.55 1.60 1.65 1.70 1.75 1.80

CGR (m/s)

1000/T (°K)

Q = 119 ± 18 kJ/mole 

 (28.5 ± 4.4 kcal/mole)

ETH

Alloy 82 

PWR Water 

333 315 298 283

Temperature (°C)

352

(a) (b) 

10-12

10-11

10-10

10-9

1.55 1.60 1.65 1.70 1.75 1.80

CGR (m/s)

1000/T (°K)

Q = 249 ± 55 kJ/mole 

 (59.5 ± 13.1 kcal/mole)

Lockheed Martin 

Alloy 82 GTA Weld 

PWR Water 30-40 cc/kg H2

333 315 298 283

Temperature (°C)

352

Kmax = 36 MPa m1/2

Figure 34.

Temperature dependence of the CGR data for 

Alloy 82 obtained at (a) Bechtel Bettis, 

(b) ETH, and (c) Lockheed Martin at 

temperatures between 290°C (554°F) and 

360°C (680°F) (Refs. 59-62). 

(c) 

 

10-12

10-11

10-10

10-9

1.55 1.60 1.65 1.70 1.75 1.80

Crack Growth Rate (m/s)

1000/T (°K)

Q = 228 ± 45 kJ/mole 

 (54.5 ± 10.8 kcal/mole)

Westinghouse 

Alloy 182 (TS Orientation)

PWR Water 25 cc/kg H2

333 315 298 283

Temperature (°C)

352

Kmax = 23 MPa m1/2

10-12

10-11

10-10

10-9

1.55 1.60 1.65 1.70 1.75 1.80

CGR (m/s)

1000/T (°K)

Q = 248 ± 45 kJ/mole 

 (59.3 ± 10.7 kcal/mole)

CEA

Alloy 182 

PWR Water 

333 315 298 283

Temperature (°C)

352

(a) (b) 

Figure 35. Temperature dependence of the CGR data for Alloy 182 obtained at 

(a) Westinghouse and (b) CEA at temperatures between 290°C (554°F) and 

340°C (644°F) (Refs. 63-65). 
 

38 

Figure 36 shows the temperature dependence of the CGR data for the three specimens (A182-1, 

CT933H-1, and CT933H-2) tested at ANL. In addition, data generated previously at 320°C (608°F) are 

also included. The colors in the plot identify the type of weld: red for deep-groove and blue for double-J. 

Also, the 1-T specimens are shown with open symbols, and the 1/2-T specimens are shown with solid 

symbols. Note that in addition to the difference in weld geometry, the two welds were prepared from 

different heats of Alloy 182. The CGR data for the double-J welds yields an activation energy of 

252 kJ/mol (59 kcal/mol). This value is in excellent agreement with the literature data shown in Fig. 35. 

Similarly, the activation energy for the deep-groove Alloy 182 weld was calculated to be 189 kJ/mol 

(33 kcal/mol). For this latter calculation, the CGR data points obtained at 350°C (662°F) were excluded 

because the data suggest a plateau in CGR at temperatures beyond 320°C (608°F) (see the dotted line in 

Fig. 36). The average value of the activation energy based on the available data is 229 kJ/mol 

(54.6 kcal/mol). As shown in Fig. 37, the distribution of observed values is at least roughly normal. The 

lower 5th percentile estimate of the activation energy is 182 kJ/mol (43 kcal/mol). Thus, the results 

clearly indicate that the activation energy for Alloy 182 is greater than that commonly used for Alloy 600, 

130 kJ/mol (31 kcal/mol), and that a value of about 220-230 kJ/mol (52.4-54.8 kcal/mol) is a better 

descriptor for the CGR vs. temperature dependence of Alloy 182. 

The average value for Alloy 82 is 164 kJ/mol (39 kcal/mol), which is somewhat greater than that 

assumed for Alloy 600, but the uncertainty in the value makes it more difficult to justify the use of a 

higher activation energy. 

10-11

10-10

10-9

1.55 1.60 1.65 1.70 1.75 1.80

A182-1

CT933H-2

CT933H-1

CT933 1-T CT

CGR (m/s)

1000/T (°K)

Q = 252 kJ/mole 

 (59 kcal/mole)

ANL

Alloy 182 Weld 

PWR Water 10-48 cc/kg H2

333 315 298 283

Temperature (°C)

352

Kmax = 30 MPa m1/2

Q = 189 kJ/mole 

 (33 kcal/mole)

Figure 36. 

Temperature dependence of the CGR data for 

Alloy 182 obtained at ANL at temperatures 

between 290°C (554°F) and 350°C (662°F). 

180 200 220 240 260 280

Cumulative Probability (%)

Activation Energy (kJ/mol)

1

10

20

30

40

50

60

70

80

90

99.9

0.1

99

Figure 37. 

Comparison of observed activation energies for 

Alloy 182 with a normal distribution. 
 

39 

5.3. Consequence of Using a Larger Activation Energy 

Because much of the database on CGRs in Ni-alloy weld materials has been obtained at different 

test temperatures, it is necessary to normalize the data to a single temperature for the purpose of 

comparison.45 This normalization has been previously done by using the activation energy for Alloy 600, 

130 kJ/mol (31 kcal/mol).45 This section analyzes the effect of using the larger value of 220 kJ/mol 

(52.4 kcal/mol) to carry out the normalization. 

Figure 38 shows a comparison of the cumulative distribution of the parameter ! in the CGR 

relationship (Eq. 17) for Ni-alloy welds45 with the data normalized to 325°C (617°F) using activation 

energies of 130 kJ/mol and 220 kJ/mol. The 75th percentile value of ln(!) becomes -27.5396, yielding a 

value of 1.1 x 10-12 at 325°C for the !parameter, which is !36% lower than the corresponding value of 

!, 1.5 x 10-12, obtained using a value of 130 kJ/mol for the normalization. Thus, the current approach 

yields somewhat more conservative values. If 1.1 x 10-12 is used instead of 1.5 x 10-12 in Eq. 17, the 

disposition curve takes the form: 

12 1.6

Ni weld

ref

Q 1 1 a 1.1 10 exp K

R T T

! !

" # $ % = & ! ! ' ( ) * - . + ,

! . (19) 

0.0

0.2

0.4

0.6

0.8

1.0

-30.0 -29.0 -28.0 -27.0 -26.0

Alloy 182

Alloy 82

Alloy 132

Cumulative Distribution F

ln(Parameter ! at 325°C)

75th Percentile

25th Percentile

Median

ln(!) = -27.795

10-13 10-12

Parameter ! at 325°C

Data normalized with

Q = 130 kJ/mol

 

0.0

0.2

0.4

0.6

0.8

1.0

-30.0 -29.0 -28.0 -27.0 -26.0

Alloy 182

Alloy 82

Alloy 132

Cumulative Distribution F

ln(Parameter ! at 325°C)

75th Percentile

25th Percentile

Median

ln(!) = -28.067

10-13 10-12

Parameter ! at 325°C

Data normalized with

Q = 220 kJ/mol

(a) (b) 

Figure 38. Cumulative distribution of the parameter ! in the CGR relationship for Ni-alloy welds with the 

data normalized to 325°C (617°F) using an activation energy of (a) 130 kJ/mol and 

(b) 220 kJ/mol. 

Figure 39 shows the effect of using the larger value of the activation energy on the proposed 

disposition curve for Alloy 182.45 The data in the figure are from field welds from V. C. Summer and 

Davis-Besse plants. The data obtained in this program are shown with red symbols while the data from 
 

40 

the literature67 are shown with blue symbols. Also included for comparison is the proposed disposition 

curve for Alloy 600.45 While less conservative, the disposition curve obtained using an activation energy 

of 220 kJ/mole for normalization bounds the available data for field welds. This curve and the disposition 

curve for Alloy 600 bound most of the data obtained for field welds in the ANL program. The data 

obtained in the ANL program also seems close to the Alloy 600 disposition curve suggesting that the 

difference in cracking behavior between field weld alloys and Alloy 600 may not be as large as the two 

proposed disposition curves might suggest. 

10-12

10-11

10-10

10-9

10 20 30 40 50

A82 VCS WCR-01

A82 VCS WLR-01

A182 VCS BCR-01

A82 VCS (Jacko et al)

A182 VCS (Jacko et al)

J11CC-1 D-B

J11CC-3 D-B

Experimental CGR (m/s)

Stress Intensity K (MPa m1/2)

Davis-Besse and V.C. Summer 

Ni-Alloy Weld Metal

325°C Proposed Disposition

Curve for Alloy 182 Weld

Proposed Disposition

Curve for Alloy 600

Q = 220 kJ/mol

Figure 39. 

CGR data for the V.C. Summer weld and 

Davis-Besse weld alloys. Also included are 

data on V.C. Summer weld alloys by Jacko et 

al. (Ref. 67). 

5.4 Cyclic Crack Growth Rates in PWR Environment 

5.4.1 Alloy 600 

The cyclic CGRs for the Alloys 600 specimens obtained in the ANL program have been evaluated 

to develop correlations for estimating the cyclic CGRs in Alloy 600 as a function of loading and 

environmental conditions. The data were obtained from tests conducted with a triangular or slow/fast 

sawtooth waveform. The load ratio R was in the range of 0.2-0.9, Kmax was in the range of 

30-50 MPa m1/2 (27.3-45.5 ksi in1/2), and the rise time was 12 s for most of the tests.26-28,31 

The experimental CGRs in the PWR environment at 320°C and those predicted in air for the same 

loading conditions for several heats (and heat treatment conditions) of Alloy 600 are shown in Fig. 40. 

Preparation conditions included mill annealing (MA), solution annealing (SA) plus thermal treatment 

(TT), and SA alone. The tests on Heats NX8844J-26 and NX131031 and some of the tests on Heat 

NX8197 were performed in deaerated water with <5 ppb DO. Also, the rise time for Heat NX131031 

was 60-5000 s. In this figure, the data points that lie along the diagonal represent predominantly 

mechanical fatigue, and those away from the diagonal indicate environmentally enhanced crack growth. 

The results indicate that unlike the crack growth behavior in high-DO BWR water where all heats of 
 

41 

Alloy 600 showed environmental enhancement,31 only two heats of Alloy 600 show environmental 

enhancement of CGRs in low-DO PWR water. Environmental enhancement is observed for solutionannealed Heat NX9244G and some specimens of this heat in the thermally treated condition, as well as 

for Heat NX131031; for the latter heat, the load ratio was 0.7, and rise times were 60-5000 s. These 

results indicate that in addition to material conditions such as yield strength and grain-boundary coverage 

of carbides, long rise times and high R values also seem to be important for environmental enhancement 

of growth rates in PWR water. 

10-12

10-11

10-10

10-9

10-8

10-12 10-11 10-10 10-9 10-8

NX8197, MA

NX9244G, SA 1025°C + TT

NX9244G, SA 1025°C

NX9244G, SA 1115°C + TT

NX9244G, SA 1115°C

NX8844J-26, SA 1038°C

NX131031, MA

CGRenv (m/s)

CGRair (m/s)

Alloy 600

320°C , <10 ppb DO

CGRair + 4.4 x 10–7(CGRair)

0.33

Including SCC rate

at Kmax = 27 MPa m1/2

Figure 40. 

Cyclic CGR data for several Alloy 600 

heats in deaerated water at 320°C 

(608°F) under cyclic loading. 

The uppermost solid line in the figure is based on the superposition model (Eq. 10), that is, the 

growth rate in the environment is determined from the superposition of the rate in air (Eqs. 11 and 12) and 

rates due to corrosion fatigue (Eq. 14) and SCC (Eq. 16). For cyclic loading using either a triangular or a 

slow/fast sawtooth waveform, aSCC ! is determined by considering the contribution of SCC during the 

slow rise time of the cycle; an equivalent Kmax is computed to determine the contribution of fatigue 

loading. The average value of Kmax used in calculating the superposition curves is given in the figure. 

The experimental CGRs for Alloy 600 from the Davis-Besse CRDM nozzle #3 and those predicted 

in air for the same loading conditions are shown in Fig. 41. The crack plane represents a circumferential 

crack in the CRDM nozzle for specimen N3CL-01 and an axial crack for specimens N3CC-02 and 

N3CC-03. The tests were conducted with a slow/fast sawtooth waveform, load ratio of 0.3-0.7, and rise 

times of 5-1000 s. As expected, specimen orientation does not seem to have any effect on growth rates 

(e.g., the CGRs for N3CL-1 and N3CC-3 are comparable and those for N3CC-2 are slightly lower).33 

Although the results show considerable scatter, the CGRs for the Davis-Besse CRDM nozzle Alloy 600 

are generally higher than those for the heats of Alloy 600 tested earlier (Fig. 40). 
 

42 

10-12

10-11

10-10

10-9

10-8

10-12 10-11 10-10 10-9 10-8

N3CL-01

N3CC-02

N3CC-03

CGRenv (m/s)

CGRair (m/s)

Alloy 600 (Heat M3935)

(Davis Besse CRDM Nozzle #3)

316°C PWR Water 

CGRair + 6.6 x 10–7(CGRair)

0.33

1/4–T CT Specimen

1/2–T CT Specimen

Including SCC rate

at Kmax = 17 MPa m1/2

Figure 41. 

CGR data for Davis-Besse CRDM 

nozzle #3 Alloy 600 in PWR water at 

316°C (601°F) under cyclic loading. 

The uppermost solid line in Fig. 41 is also based on the superposition model. For the loading 

conditions used in these tests, the average equivalent Kmax is !17 MPa m1/2 (15.5 ksi in1/2) The 

contribution of the SCC growth rate, aSCC ! , is relatively small. Also, for the Davis-Besse nozzle alloy, 

the corrosion fatigue growth rate CF a! is better represented by 

( ) 7 0.33

env air air a a 6.6 10 a ! ! ! = + " # ! . (20) 

In Fig. 40 the cyclic CGRs in the PWR environment are represented by the best-fit curve for 

Alloy 600 in high-temperature water determined previously in our program (Eq. 14).31 Even when the 

expected SCC contribution is added, the prediction underestimates the data. Therefore, a new best-fit 

curve was calculated based on cyclic data for Davis-Besse Alloy 600. This approach was warranted 

because, in all three Davis-Besse Alloy 600 specimens, the fracture mode was IG regardless of the 

loading conditions. The fracture mode appeared to have changed to IG from the very first grain boundary 

encountered (hence, the enhanced cyclic CGRs). The other two specimens exhibited either a predominant 

IG (N3CC-2) or a partial IG (N3CC-3) fracture mode during precracking. 

The Davis-Besse CRDM nozzle Alloy 600 exhibits a predominantly IG fracture even during fatigue 

loading. The fracture surface of specimen N3CL-1 near the machine notch is shown Fig. 28. 

Transgranular growth is observed at the very beginning of the test (i.e., near the machine notch), but 

changes to IG when the first grain boundary is encountered. The fact that IG growth takes place in a 

regime dominated by mechanical fatigue (which normally results in TG cracking) suggests that the grain 

boundaries may have suffered some form of sensitization during fabrication and/or during two decades in 

service. As noted previously, metallographic examination of nozzle #3 revealed what is typically 

considered a “good” microstructure (i.e., extensive grain-boundary coverage of Cr-rich carbides and 

relatively low or average tensile strength). These conditions are typically associated with low 

susceptibility of the material to PWSCC. Differences in the microstructure in terms of extent and nature 

of carbide precipitation and segregation of other impurities on the grain boundaries may be important and 

should be investigated. As such, the chemical compositions of the grain boundary carbides should be 

determined as well as any impurity and/or major element segregation at the boundary. 
 

43 

5.4.2 Ni-Alloy Welds 

The cyclic CGR data on Ni-alloy welds in simulated PWR environments are available from 

!120 tests conducted on Alloys 182, 82, 152, and 52 at 243-345°C (469-653°F).55-58 The loading 

conditions for these tests include R = 0.1-0.75, Kmax = 20-100 MPa m1/2 (18.2-91 ksi in1/2), and rise time 

of 0.5-5000 s. The results indicate very little effect of PWR environment on the cyclic CGRs of these 

welds. However, only about 10% of the data was obtained under conditions for which significant 

environmental enhancement would be expected. 

The cyclic CGR data obtained at ANL on laboratory-prepared Alloy 182 weld specimens32 in the 

PWR environment at 320°C (608°F) are shown in Fig. 42a. For comparison, similar cyclic CGR data 

obtained by Van der Sluys et al.57 on Alloys 82 and 52 in PWR environment at 315°C (600°F) are shown 

in Fig. 42b. In both figures, rates predicted from the superposition model are shown by the solid line, and 

those predicted without the SCC component are shown by the dashed line. For Alloy 182, select loading 

conditions appear to show minimal environmental enhancement; however, the vast majority do not. 

Similarly, the Alloy 82 data do not appear to show environmental enhancement. The conditions that 

resulted in environmental enhancement of CGRs in Alloy 182 are load ratios of R ! 0.7 and rise times of 

1000 s. 

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-12 10-11 10-10 10-9 10-8 10-7 10-6

CT31-W01 TS

CT31-W02 TS

CT933-TS

CT933-LS

CT933-TL

CGRenv Alloy 182 (m/s)

CGRair Ni-alloy Weld (m/s)

Alloy 182

Simulated PWR Water, 320°C

CGRair + 0.018 (CGRair)

0.78

Including SCC rate at

Kmax = 24 MPa m0.5

TS, LS orientations

Including SCC rate at

Kmax = 24 MPa m0.5

TL orientation

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-12 10-11 10-10 10-9 10-8 10-7 10-6

A52 (Van Der Sluys et al.)

A82 (Van Der Sluys et al.)

CGRenv Alloy 182 (m/s)

CGRair Ni-alloy Weld (m/s)

Alloy 82

Simulated PWR Water, 315°C

CGRair + 0.018 (CGRair)

0.78

Including SCC rate at

Kmax = 30 MPa m0.5

(a) (b) 

Figure 42. CGR data in PWR environment for (a) Alloy 182 SMA weld at 320°C (608°F), and (b) Alloy 82 

and Alloy 52 welds at 315°C (600°F) (Ref. 57) as a function of the growth rates for Alloy 182 

weld in air. 

Of the five laboratory-prepared Alloy 182 weld specimens (Fig. 42a), only specimen CT31-W01, 

tested at R = 0.7 and rise time = 1000 s, showed environmental enhancement of cyclic CGRs. For this 

specimen, the transition from TG to IG fracture mode is shown in Fig. 43. The images appear to indicate 

that the transition from TG to IG occurred readily. 
 

44 

(a) (b) 

Figure 43. (a) Transition from TG to IG fracture and (b) higher magnification micrograph of the boxed 

area showing the tip of a few secondary cracks. Crack extension from bottom to top. 

Figure 44 shows the CGR data obtained at ANL on 1/2-T CT specimens of laboratory-prepared 

Alloy 182 weld in the PWR environment at 310-350°C (590-662°F) as a function of the rates for 

Alloy 182 weld in air under the same loading conditions. The solid lines represent rates predicted from 

the superposition model, and the dashed line represents the predicted rates without the SCC contribution. 

The cyclic CGR data for the three specimens are consistent with previous results summarized in Fig. 42. 

The loading conditions for these specimens did not include the conditions that resulted in enhanced 

growth rates in specimen CT31-W01 (R = 0.7 and rise time = 1000 s). 

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-12 10-11 10-10 10-9 10-8 10-7 10-6

CT933H-1

CT933H-2

A182-1

CGRenv Alloy 182 (m/s)

CGRair Ni-alloy Weld (m/s)

CGRair + 0.018 (CGRair)

0.78

Kmax = 21 MPa m0.5, 310°C

Alloy 182 Weld

Simulated PWR Water 

Kmax = 21 MPa m0.5, 320°C

Kmax = 21 MPa m0.5, 350°C

Including SCC rate at:

Figure 44. 

CGR data for Alloy 182 SMA weld-metal 

specimens as a function of growth rates for 

Ni-weld alloy in air. The tests were conducted 

in simulated PWR environment in the 

310-350°C (590-662°F) temperature range. 

 

Figure 45 shows the cyclic CGR data in the PWR environment for Ni-alloy welds from field 

components plotted as a function of the growth rate in air under the same loading conditions. Figure 45a 

shows the data for Alloy 182 from the Davis-Besse CRDM nozzle #11 J-groove weld and V. C. Summer 

reactor vessel nozzle butter, and Fig. 45b shows the data for Alloy 82 from the V. C. Summer reactor 

vessel nozzle-to-pipe weld. For the V. C. Summer weld, the fracture plane in specimens WCR-01 and 

WLR-01 represents axial and circumferential through-wall cracks in the plane of the columnar 
 

45 

grains/dendrites and growing along the dendrites, while specimen BCR-01 represents axial through-wall 

cracks in the plane of the columnar grains/dendrites and growing across the dendrites. 

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-12 10-11 10-10 10-9 10-8 10-7 10-6

VCS BCR-01

D-B J11CC-1

D-B J11CC-3

CGRenv Alloy 182 (m/s)

CGRair Ni-alloy Weld (m/s)

Alloy 182 field welds (D-B and VCS)

Simulated PWR Water

CGRair/4 + 0.018 (CGRair/4)0.78

Including SCC rate at

Kmax = 20 MPa m0.5

CC orientation

Including SCC rate at

Kmax = 25 MPa m0.5

CR orientation

10-12

10-11

10-10

10-9

10-8

10-7

10-6

10-12 10-11 10-10 10-9 10-8 10-7 10-6

WCR-01

WLR-01

CGRenv Alloy 182 (m/s)

CGRair Ni-alloy Weld (m/s)

Including SCC rate at

Kmax = 28 MPa m0.5

V. C. Summer Nozzle-to-Pipe Alloy 82 Weld

Simulated PWR Water at 320°C 

CGRair/4 + 0.018 (CGRair/4)0.78

(a) (b) 

Figure 45. Cyclic CGR data for (a) Alloy 182 and (b) Alloy 82 from the Davis-Besse CRDM nozzle 

J-groove weld and V.C. Summer reactor vessel nozzle-to-pipe weld. 

The results indicate that the cyclic CGRs in air of the field welds are generally lower than those 

predicted by Eqs. 11 and 13 for laboratory-prepared Alloy 182 welds. In Fig. 45, under loading 

conditions that result in essentially mechanical fatigue (i.e., growth rates >1 x 10-9 m/s in air) the CGRs 

of the field welds are a factor of 3-6 lower than the predicted values. When environmental enhancement 

occurs, the CGRs in the PWR environment compared with those in air increase with decreasing CGRs in 

air. As before, the solid and dashed lines are based on the superposition model with and without the SCC 

contribution, respectively. The SCC contribution for the Alloy 82 weld was computed by considering the 

CGRs for Alloy 82 to be a factor of 2.6 lower than those for Alloy 182. After accounting for the 

difference in CGRs in air for the field welds and laboratory welds, the data in Fig. 45 show some 

environmental enhancement of CGRs for Ni-alloy welds in the PWR environment. As with the 

laboratory-prepared welds, the conditions that appear to show the most environmental enhancement are 

loading ratios of R ! 0.7 and rise times of 1000 s. 

Figure 46 shows images from the fracture surfaces of specimens WLR-01 and WCR-01. In 

specimen WLR-01, the transition from a fatigue TG crack to SCC IG crack occurred apparently readily, 

while in specimen WCR-01 a more mixed fracture mode occurred. Nevertheless, the reason for the lower 

growth rates for the field welds relative to those for the laboratory-prepared welds is not clear. 
 

46 

(a) (b) 

Figure 46. Micrographs of the fracture surface of specimen (a) WCR-01 and (b) WLR-01. Crack 

advance is from top to bottom. 
 

47

6 Summary 

Crack growth rate results are presented for Alloy 182 SMA weld metal in the simulated PWR 

environment in the temperature range 290-350°C. The tests were conducted on a 1-T compact tension 

specimen from an Alloy 182 double-J weld and on two companion 1/2-T compact tension specimens 

from a deep-groove Alloy 182 weld. Metallographic examination of the fracture surfaces revealed 

relatively straight crack fronts. The fracture modes correlated well with the test conditions. High rise 

times or long hold periods favor IG SCC. Also, IG cracking apparently advanced more readily along 

some grain orientations than others, resulting in crack fronts with occasional unbroken ligaments and few 

regions of TG cracking. 

In addition, metallographic examinations were performed to characterize the microstructure of the 

weld. The weld structure consisted of vertically columnar grains and dendrites. The weld microstructure 

was also examined by OIM to determine the orientations of the grains and the type of grain boundaries 

present. The results indicate that the intragranular dendrites are coherent and are, therefore, expected to 

be very resistant to cracking. The proportion of cracking-resistant, coincident-site lattice boundaries in 

the weld was found to be relatively small (≈30%) by comparison with that of Alloy 600 in a solutionannealed condition (≈50%). The comparison suggests that Alloy 182 is more susceptible to cracking than 

Alloy 600, and this inference is supported by the CGR findings presented in this report. 

The activation energy for SCC crack growth in Alloy 182 SMA weld alloy was calculated from the 

SCC CGR vs. temperature curves. The curves yield activation energies of 252 kJ/mol for the double-J 

weld material and 189 kJ/mol for the deep-groove weld material. These values are in good agreement 

with the data reported in the literature. The data reported here and those in the literature suggest that the 

average activation energy for Alloy 182 welds is on the order of 220-230 kJ/mol, higher than the 130 

kJ/mol commonly used for Alloy 600. The activation energy data for Alloy 82 are more limited. The 

average value from the literature data is about 169 kJ/mol. The consequences of using these larger values 

for the analysis of SCC CGR data were analyzed. The analysis showed that when an activation energy of 

220 kJ/mol is used, the 75th percentile value of ln(α), which has been the value used for a disposition 

curve, becomes -27.5396, yielding a value of 1.1 x 10-12 at 325°C for the α parameter. Thus, the 

disposition curve for Ni-alloy welds based on data that have been normalized with an activation energy of 

220 kJ/mol is ≈36% lower than the curve based on data normalized with 130 kJ/mol; thus, the use of the 

130 kJ/mol value leads to somewhat more conservative results. 

In addition to the tests on laboratory prepared welds, CGR tests were performed on weldment 

samples from a Davis–Besse CRDM J-groove nozzle weld and from a V.C. Summer hot–leg nozzle to 

safe end weld. For the Davis-Besse Alloy 182 weld specimens, the SCC CGRs under constant load are 

an order of magnitude lower than the disposition curve proposed for Alloy 182 weld metals. For the V.C. 

Summer weld specimens, the CGR data under constant load were lower than data obtained by others on 

the same alloys and lower than the proposed disposition curve. The data from the field welds correspond 

to values in the lower quartile of the distribution of crack growth rate data obtained on laboratory 

specimens. These results are not so statistically unlikely that they can be said to prove that there is a 

difference between field and laboratory welds, but such a difference would be consistent with the 

observations that operating experience indicates that PWSCC appears to occur more frequently in 

wrought Ni-base Alloy 600 components than in the weld metal Alloys 82 and 182 used with Alloy 600, 

despite the fact that in laboratory tests in PWR coolant environments, the stress corrosion cracking (SCC) 

susceptibility of Alloy 182 is usually found to be greater than that of Alloy 600, while that of Alloy 82 is 

comparable to that of Alloy 600. 
 

48 

laboratory welds is the degree of structural constraint imposed on the welds during the welding process, 

which is typically much greater for the field welds on actual components. However, metallographic 

studies on the welds away from the crack planes do not show hot cracks so that the secondary cracks 

seem to be part of the corrosion process. Another factor that could also contribute to a difference between 

CGRs from field and laboratory welds are the adjustments that are often made to laboratory data to 

account for the non-uniformity and less than complete engagement of the crack fronts. In any case, the 

use of the current database developed on laboratory welds would appear to be conservative. 

The report also discusses the crack growth rate results for actual Alloy 600 component material and 

the environment enhancement of CGRs under cyclic loading. The SCC CGRs of the Alloy 600 nozzle 

from Davis-Besse are a factor of 4–8 higher than those of the median curve for Alloy 600. The growth 

rates correspond to the !95th percentile of the various data sets used in developing the median curve, i.e., 

the nozzle material exhibits very high susceptibility to SCC. The material exhibits predominantly IG 

fracture, even during precracking. 

The environmental enhancement of CGRs under cyclic loading was determined for the Alloy 600 

and Ni-alloy welds tested relative to the CGRs that would be expected under the same loading conditions 

for the respective alloy in air. The cyclic CGRs were analyzed by a superposition model to establish the 

individual contributions of mechanical fatigue, corrosion fatigue, and SCC. Environmental enhancement 

was observed for solution-annealed Alloy 600 and for some specimens in the thermally treated condition. 

Although the results show considerable scatter, the CGRs for the Davis-Besse CRDM nozzle Alloy 600 

are generally higher than those for the heats of Alloy 600 tested previously in the program. Both 

laboratory-prepared and field Ni-weld alloys showed minimal environmental enhancement for loading 

ratios R " 0.5 and rise time greater that 1000 s. 
 

49 

References 

1. Bamford, W. H., and J. F. Hall, “A Review of Alloy 600 Cracking in Operating Nuclear Plants: 

Historical Experience and Future Trends ,” The Vessel Penetration Inspection, Crack Growth and 

Repair Conference, Oct. 2, 2003, Gaithersburg, MD, 2003. 

2. Scott, P., “An Analysis of Primary Water Stress Corrosion Cracking in PWR Steam Generators,” 

Proc. of the Specialists Meeting on Operating Experience with Steam Generators, Brussels, 

Belgium, pp. 5-6, 1991. 

3. Cattant, F., “Lessons Learnt from the Examination of Tubes Pulled from Electricite de France 

Steam Generators,” Nucl. Eng. Des. 168, pp. 241-253, 1997. 

4. Diercks, D. R., W. J. Shack, and J. Muscara, “Overview of Steam Generator Tube Degradation and 

Integrity Issues,” Nucl. Eng. Des. 194, pp. 19-30, 1999. 

5. U.S. NRC Information Notice No. 90-49, “Stress Corrosion Cracking in PWR Steam Generator 

Tubes,” Aug. 1990. 

6. U.S. NRC Information Notice No. 91-43, “Recent Incidents Involving Rapid Increases in Primaryto-Secondary Leak Rate,” July 1991. 

7. U.S. NRC Information Notice No. 94-05, “Potential Failure of Steam Generator Tubes with 

Kinetically Welded Sleeves,” Jan. 1994. 

8. U.S. NRC Information Notice No. 89-33, “Potential Failure of Westinghouse Steam Generator 

Tube Mechanical Plugs,” March 1989. 

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50 

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51 

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52 

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Stress Corrosion Resistance,” Scripta Metallugica et Materialia, 25 (8), pp. 1775-1780, 1991. 
 

53 

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54 

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Crack Growth and Repair Conference, Gaithersburg, MD, 2003. 

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Intl. Conf. on Environmental Degradation of Materials in Nuclear Power Systems-Water Reactors, 

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Crack Growth and Repair Conference, Gaithersburg, MD, 2003. 

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Rate of Nickel Based Alloy 132 Weld Metal in PWR Water,” Proc. 11th Intl. Conf. on 

Environmental Degradation of Materials in Nuclear Power Systems-Water Reactors, NACE 

International, Houston, TX, 2003. 
 

55 

Appendix A: Crack Growth Rate Data for Davis-Besse and 

V. C. Summer Ni-base alloys 

A.1 Specimen N3CL-1 of Alloy 600 in PWR water at 316°C 

Table A1. Crack growth results for 1/4–T CT specimen N3CL-1 of Alloy 600 in PWR watera at 316°C 

Test 

Test 

Time, 

Conduc -

tivity 

R 

Load 

Rise 

Time,

Down

Time,

Hold 

Time, Kmax, !K, 

Growth 

Rate, 

Allowed 

Kmax, 

Kapp – 

Kmax,

b

Crack 

Length

Period (h) (μS/cm) Ratio (s) (s) (s) (MPa m1/2) (MPa m1/2) (m/s) (MPa m1/2) (%) (mm) 

Pre a 51 21.0 0.26 1 1 0 19.0 14.1 5.07E-09 20.3 -6 5.993 

Pre b 71 – 0.28 5 5 0 18.4 13.3 3.09E-09 20.1 -8 6.117 

1 115 – 0.53 30 4 0 20.0 9.4 4.30E-10 19.9 1 6.199 

2 194 18.5 0.54 300 4 0 20.3 9.4 2.55E-10 19.9 2 6.240 

3a 225 18.5 0.52 300 4 0 21.5 10.3 8.88E-11 19.8 9 6.269 

3b 316 14.7 0.51 300 4 0 23.2 11.4 2.63E-10 19.7 18 6.353 

4 436 12.8 0.71 300 4 0 24.0 7.0 1.10E-10 19.6 22 6.400 

5 676 15.4 0.70 1000 12 0 23.1 6.9 negligible 19.6 18 6.398 

6 722 15.4 0.51 300 4 0 23.7 11.6 4.42E-10 19.5 22 6.471 

7a 818 13.0 0.50 1000 12 0 23.2 11.6 5.45E-11 19.4 20 6.493 

7b 868 13.0 0.51 1000 12 0 24.2 11.8 3.73E-10 19.3 25 6.562 aSimulated PWR water with 2 ppm Li, 1000 ppm B, and !2 ppm dissolved hydrogen. Effluent DO was <10 ppb. bBased on flow stress. 
 

56 

A.2 Specimen N3CC-2 of Alloy 600 in PWR water at 316°C 

Table A2. Crack growth data for 1/4–T CT specimen N3CC-2 of Alloy 600 in PWR watera at 316°C 

Test 

Test 

Time, 

Conduc -

tivity 

R 

Load 

Rise 

Time,

Down 

Time, 

Hold

Time, Kmax, !K, 

Growth 

Rate, 

Allowed 

Kmax, 

Kapp – 

Kmax,

b

Crack 

Length

Period (h) (μS/cm) Ratio (s) (s) (s) (MPa m1/2) (MPa m1/2) (m/s) (MPa m1/2) (%) (mm) 

Pre a 189 22 0.31 5 5 0 21.9 15.1 7.52E-10 19.8 10 6.243 

1 383 20 0.51 300 4 0 21.1 10.3 2.13E-11 19.8 6 6.270 

2 457 17 0.51 30 4 0 21.7 10.6 6.05E-11 19.7 10 6.335 

3 550 16 0.51 300 4 0 21.6 10.6 9.46E-12 19.7 10 6.338 

4 580 15 0.51 30 4 0 23.6 11.5 8.59E-10 19.5 20 6.415 

5 765 15 0.50 300 12 0 23.5 11.7 8.87E-11 19.5 21 6.471 

6 885 15 0.70 300 12 0 23.7 7.1 5.28E-11 19.4 22 6.492 

7 912 16 0.51 300 12 0 23.7 11.6 1.01E-10 19.4 22 6.502 

8 1197 20 0.52 300 12 3600 24.3 11.7 6.92E-11 19.3 26 6.541 

9 1365 21 0.48 300 12 0 23.2 12.1 7.53E-11 19.3 25 6.617 

10 1530 21 1.00 – – – 23.3 – 9.67E-11 19.2 21 6.658 aSimulated PWR water with 2 ppm Li, 1000 ppm B, and !2 ppm dissolved hydrogen. Effluent DO was <10 ppb. bBased on flow stress. 
 

57 

A.3 Specimen N3CC-3 of Alloy 600 in PWR water at 316°C 

Table A3. Crack growth data for specimen N3CC-3 of Alloy 600 in PWR watera at 316°C 

Test 

Test 

Time, 

Conduc -

tivity 

R 

Load 

Rise 

Time,

Down

Time,

Hold 

Time, Kmax, !K, 

Growth 

Rate, 

Allowed 

Kmax, 

Kapp – 

Kmax,

b

Crack 

Length

Period (h) (μS/cm) Ratio (s) (s) (s) (MPa m1/2) (MPa m1/2) (m/s) (MPa m1/2) (%) (mm) 

Pre a 50 20 0.36 0.25 0.25 0 19.7 12.6 1.35E-08 27.6 -29 12.777

Pre b 55 20 0.35 0.25 0.25 0 19.1 12.4 8.41E-09 27.5 -30 12.872

Pre c 72 19 0.33 5.0 5.0 0 18.8 12.6 2.03E-09 27.4 -31 12.943

Pre d 78 19 0.32 0.5 0.5 0 19.6 13.3 1.31E-08 27.2 -28 13.065

1 97 18 0.50 12 4 0 19.5 9.7 9.08E-10 27.1 -28 13.157

2 122 18 0.49 60 4 0 19.6 10.0 3.55E-10 27.1 -28 13.186

3 147 18 0.48 300 4 0 19.8 10.3 2.80E-10 27.1 -27 13.206

4 174 17 0.63 1,000 12 0 20.0 7.4 2.18E-10 27.0 -26 13.249

5a 230 16 1.00 - - - 19.3 0.0 7.85E-11 27.1 -29 13.218

5b 339 15 1.00 - - - 19.7 0.0 1.25E-10 26.9 -27 13.319

6a 386 14 1.00 - - - 24.9 0.0 2.28E-10 26.8 -7 13.404

6b 529 12 1.00 - - - 25.1 0.0 1.55E-10 26.7 -6 13.510

7 648 12 1.00 - - - 30.1 0.0 negligible 26.5 13 13.637 aSimulated PWR water with 2 ppm Li, 1000 ppm B, and !2 ppm dissolved hydrogen. Effluent DO was <10 ppb. bBased on flow stress. 
 

58 

A.4 Specimen J11CC-1 of Alloy 182 J–groove weld in PWR water at 316°C 

Table A4. Crack growth data for specimen J11CC-1 of Alloy 182 J–groove weld in PWR watera at 

316°C. 

Test 

Test 

Time, 

Conduc -

tivity 

R 

Load 

Rise 

Time,

Down

Time,

Hold 

Time, Kmax, !K, 

Growth 

Rate, 

Allowed 

Kmax, 

Kapp – 

Kmax,

c

Crack 

Length

Period (h) (μS/cm) Ratio (s) (s) (s) (MPa m1/2) (MPa m1/2) (m/s) (MPa m1/2) (%) (mm) 

Pre a 145 17.2 0.32 0.5 0.5 0 21.3 14.8 4.05E-08 20.9 2 6.153 

Pre b 186 17.2 0.32 50 50 0 20.6 14.3 1.27E-09 20.7 0 6.264 

1 331 15.6 0.51 300 4 0 20.6 10.3 1.13E-10 20.5 1 6.333 

2 453 13.0 0.70 1000 12 0 21.4 6.5 5.71E-12 20.5 4 6.342 

3a 644 22.2 0.51 1000 12 0 21.2 10.8 3.88E-11 20.5 4 6.363 

3b 745 20.0 0.50 1000 12 0 21.1 10.7 negligible 20.4 3 6.382 

4 937 14.7 0.50 1000 12 0 22.5 11.5 8.25E-12 20.4 10 6.392 

5 987 14.7 0.50 300 12 0 22.5 11.5 5.53E-11 20.4 10 6.401 

6 1106 18.2 0.49 300 12 3600 22.6 11.6 negligible 20.4 11 6.402 

7 1178 18.9 0.52 30 4 0 22.9 11.2 7.93E-10 20.2 13 6.525 

8 1248 17.2 0.49 300 12 0 22.9 11.9 1.06E-10 20.1 14 6.550 

9 1415 16.1 0.49 300 12 3600 22.9 6.0 negligible 20.2 14 6.543 

10 1443 18.2 0.49 30 12 0 23.2 12.1 1.13E-09 20.0 16 6.616 

11 1513 18.5 0.49 300 12 0 27.3 14.2 3.00E-10 19.8 38 6.742 

12a 1548 18.5 0.49 300 12 3600 27.6 14.3 3.98E-11d 19.8 40 6.747 

12b 1607 18.2 0.52 300 12 3600 29.9 0.0 3.01E-11e 19.7 52 6.777 

13 1948 16.7 1.0 – – – 28.5 0.0 2.20E-11f 19.7 45 6.797 aSimulated PWR water with 2 ppm Li, 1000 ppm B, and !2 ppm dissolved hydrogen. Effluent DO was <10 ppb. bBased on flow stress. 

dBased on total crack extension during the period. From superposition model, CGR during the constant load was estimated to 

1.68E-11 m/s. 

eFrom superposition model, CGR during the constant load was estimated to 1.63E-11 m/s. 

f

Based on total crack extension during the period. 
 

59 

A.5 Specimen J11CC-3 of Alloy 182 J–groove weld in PWR water at 316°C 

Table A5. Crack growth data for specimen J11CC-3 of Alloy 182 J–groove weld in PWR watera at 

316°C 

Test 

Test 

Time, 

Conduc -

tivity 

R 

Load 

Rise 

Time,

Down

Time,

Hold 

Time, Kmax, !K, 

Growth 

Rate, 

Allowed 

Kmax, 

Kapp – 

Kmax,

b

Crack 

Length

Period (h) (μS/cm) Ratio (s) (s) (s) (MPa m1/2) (MPa m1/2) (m/s) (MPa m1/2) (%) (mm) 

Pre a 78 21 0.31 0.25 0.25 0 17.8 12.3 6.55E-08 29.4 -39 12.373

Pre b 104 21 0.32 10 10 0 17.9 12.2 7.17E-09 29.1 -38 12.605

1 221 18 0.50 60 4 0 18.7 9.4 7.44E-10 28.5 -34 13.049

2 412 20 0.51 300 12 0 18.8 9.2 1.33E-10 28.4 -34 13.134

3 507 20 1.00 – – – 18.9 0.0 3.92E-11c 28.5 -34 13.099

4 580 19 0.50 300 12 0 20.8 10.4 2.30E-10 28.3 -26 13.233

5 742 15 0.50 1000 12 0 21.0 10.5 7.12E-11 28.3 -26 13.254

6a 864 13 1.00 – – – 21.0 0.0 3.93E-11c 28.3 -26 13.204

6b 934 13 0.51 12 12 3600 20.9 10.2 3.66E-11 28.3 -26 13.204

7a 937 13 0.51 1000 12 0 21.0 10.3 – 28.3 -26 13.240

7b 941 19 0.70 1000 12 0 26.5 8.0 – 28.2 -6 13.295

8 1079 33 0.70 12 12 3600 26.6 8.0 4.00E-11 28.2 -6 13.271

9 1084 33 0.70 1000 12 0 26.8 8.0 – 28.2 -5 13.312

10 1248 31 0.70 12 12 3600 30.9 9.3 4.65E-11 28.2 10 13.327

11 1349 29 0.70 1000 12 0 31.0 9.3 8.23E-11 28.0 10 13.410

12 1516 28 0.70 12 12 3600 30.7 9.2 negligible 28.1 9 13.368

13 1537 28 0.70 300 12 0 30.9 9.3 – 28.0 10 13.405

aSimulated PWR water with 2 ppm Li, 1000 ppm B, and !2 ppm dissolved hydrogen. Effluent DO was <10 ppb. bBased on flow stress. 

cValues estimated from the total crack extension during the test period and not from the slope of the crack length vs. time plot. 
 

60 

A.6 Specimen WCR–01 of Alloy 82 weld in PWR water at 320°C 

Table A6. Crack growth results for specimen WCR–01 of Alloy 82 weld in PWR watera

 at 320°C. 

Test 

Test Time, O2 

Conc., 

Load 

Ratio 

Rise 

Time, 

Down 

Time, 

Hold 

Time, Kmax,

b !K, CGRenv, 

Estimated 

CGRair, 

Crack 

Length, 

Period (h) (ppb) R (s) (s) (s) (MPa m1/2) (MPa m1/2) (m/s) (m/s) (mm) 

Pre a 100 <10 0.32 1 1 0 22.9 15.5 2.83E-09 4.16E-08 12.289 

Pre b 124 <10 0.31 0.5 0.5 0 22.8 15.8 4.86E-09 8.60E-08 12.351 

Pre c 140 <10 0.29 5 5 0 21.9 15.6 1.72E-09 7.80E-09 12.379 

Pre d 146 <10 0.30 0.25 0.25 0 22.7 15.9 9.34E-09 1.72E-07 12.484 

1 258 <10 0.30 30 2 0 22.9 11.4 1.39E-10 6.34E-10 12.649 

2 380 <10 0.51 30 2 0 22.4 11.0 2.06E-10 5.58E-10 12.736 

3 523 <10 0.50 300 12 0 22.4 11.2 3.18E-11 5.95E-11 12.779 

4 593 <10 1.00 – – – 22.9 0.0 4.11E-11 – 12.789 aSimulated PWR water with 2 ppm Li, 1000 ppm B, and !2 ppm dissolved hydrogen. Conductivity was 22 ± 3 μS/cm, 

and pH was 6.4. 

bAt the end of the test the maximum allowed Kmax based on the final load was 28.8 MPa m1/2. 
 

61 

A.7 Specimen BCR–01 of Alloy 82 SMA weld in PWR water at 320°C 

Table A7. Crack growth data for specimen BCR–01 of Alloy 182 SMA weld in PWR watera at 320°C 

Test 

Test Time, O2 

Conc., 

Load 

Ratio 

Rise 

Time, 

Down 

Time, 

Hold 

Time, Kmax,

b !K, CGRenv, 

Estimated 

CGRair, 

Crack 

Length, 

Period (h) (ppb) R (s) (s) (s) (MPa m1/2) (MPa m1/2) (m/s) (m/s) (mm) 

Pre a 31.0 <10 0.30 0.25 0.25 0 20.0 14.0 4.00E-09 1.03E-07 12.046 

Pre b 54.0 <10 0.30 0.25 0.25 0 20.7 14.5 1.82E-08 1.20E-07 12.269 

Pre c 71.0 <10 0.30 0.25 0.25 0 21.2 14.6 2.50E-08 1.27E-07 12.407 

Pre d 75.0 <10 0.30 0.25 0.25 0 20.7 14.3 2.42E-08 5.75E-08 12.573 

1 386.0 <10 1.00 – – – 20.8 0.0 2.35E-11 - 12.616 

2a 389.0 <10 0.50 0.25 0.25 0 24.2 12.1 1.49E-08 9.76E-08 12.715 

2b 413.0 <10 0.50 0.25 0.25 0 24.4 12.2 1.25E-08 1.00E-07 12.905 

3 507.0 <10 0.50 50 2 0 25.2 12.6 3.16E-10 5.76E-10 13.041 

4 839.0 <10 1.00 – – – 25.4 0.0 2.34E-11 - 13.057 

5 842.0 <10 0.30 1 1 0 28.0 19.6 5.04E-08 1.02E-07 13.291 

6a 845.0 <10 0.50 1 1 0 27.9 13.9 8.18E-09 4.36E-08 13.369 

6b 869.0 <10 0.50 1 1 0 28.7 14.3 9.17E-09 4.88E-08 13.514 

7 939.0 <10 0.50 300 12 0 28.8 14.4 2.79E-10 1.66E-10 13.569 

8 1,340.0 <10 1.00 – – – 30.2 0.0 4.49E-11 - 13.594 

9 1,343.0 <10 0.30 1 1 0 33.0 23.1 5.73E-08 2.01E-07 13.791 

10 1344.0 <10 0.50 1 1 0 33.6 16.8 1.44E-08 9.36E-08 13.869 

11 1346.0 <10 0.50 300 12 0 33.2 16.6 5.78E-10 2.95E-10 13.901 

12 1702.0 <10 1.00 – – – 33.9 0.0 4.14E-11 - 13.930 

aSimulated PWR water with 2 ppm Li, 1000 ppm B, and !2 ppm dissolved hydrogen. Conductivity was 22 ± 3 μS/cm, and pH 

was 6.4. 

bAt the end of the test the maximum allowed Kmax based on the final load was 27.5 MPa m1/2. 
 

62 

A.8 Specimen WLR–01 of Alloy 82 SMA weld in PWR water at 320°C 

Table A8. Crack growth data for specimen WLR–01 of Alloy 82 SMA weld in PWR watera at 320°C. 

Test 

Test Time, O2 

Conc. 

Load 

Ratio 

Rise 

Time, 

Down 

Time, 

Hold 

Time, Kmax,

b !K, CGRenv, 

Estimated 

CGRair, 

Crack 

Length, 

Period (h) (ppb) R (s) (s) (s) (MPa m1/2) (MPa m1/2) (m/s) (m/s) (mm) 

Pre a 103 <10 0.30 0.5 0.5 0 19.8 13.8 8.22E-09 5.05E-08 12.640 

Pre b 167 <10 0.30 50 2 0 19.6 9.8 2.61E-10 2.12E-10 12.719 

Pre c 196 <10 0.30 1 1 0 16.7 8.4 1.35E-10 5.54E-09 12.729 

Pre d 215 <10 0.30 0.5 0.5 0 19.7 9.8 2.82E-09 2.15E-08 12.837 

Pre e 226 <10 0.30 0.25 0.25 0 19.3 13.5 1.20E-08 9.13E-08 13.017 

Pre f 265 <10 0.50 1 1 0 23.0 11.5 1.52E-08 2.04E-08 14.251 

1 315 <10 0.50 50 2 0 22.6 11.3 5.93E-10 3.73E-10 14.349 

2 389 <10 0.70 1000 12 0 22.8 7.1 2.31E-10 5.35E-12 14.418 

3 504 <10 0.70 1000 12 0 23.5 7.7 1.60E-10 6.96E-12 14.531 

4 841 <10 1.00 - - - 23.2 0.0 2.07E-11 - 14.531 

5 1082 <10 1.00 - - - 31.2 0.0 5.43E-12 - 14.590 

6 1203.0 <10 0.70 1000 12 0 31.0 9.3 2.04E-11 1.71E-11 14.590 

7 1392.0 <10 0.70 300 12 0 31.3 9.4 1.94E-11 6.01E-11 14.602 

8 1398.0 <10 0.30 0.5 0.5 0 34.7 24.3 1.03E-07 5.03E-07 15.614 

9 1417.0 <10 0.50 100 12 0 38.1 19.1 7.01E-09 1.62E-09 16.099 

10 1439.0 <10 0.50 300 12 0 41.9 20.9 8.00E-08 7.95E-10 16.410 

11 1466.0 <10 0.50 50 2 0 43.7 21.8 2.03E-09 5.55E-09 16.656 

12 1531.0 <10 0.50 300 12 0 44.6 22.3 8.95E-10 1.00E-09 16.854 

13 1634.0 <10 0.50 1000 12 0 46.6 23.3 4.12E-10 3.60E-10 16.964 

14 2469.0 <10 1.00 - - - 49.8 0.0 7.71E-11 - 17.182 

aSimulated PWR water with 2 ppm Li, 1000 ppm B, and !2 ppm dissolved hydrogen. Conductivity was 22 ± 3 μS/cm, and pH 

was 6.4. 

bAt the end of the test the maximum allowed Kmax based on the final load was 21.7 MPa m1/2. 
BIBLIOGRAPHIC DATA SHEET

(See instructions on the reverse)

NRC FORM 335

(9-2004)

NRCMD 3.7

U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER

 (Assigned by NRC, Add Vol., Supp., Rev.,

 and Addendum Numbers, if any.)

NUREG/CR 6964

ANL- 07/12

3. DATE REPORT PUBLISHED

MONTH

May

YEAR

2008

4. FIN OR GRANT NUMBER

Y6388

2. TITLE AND SUBTITLE

Crack Growth Rates and Metallographic Examinations of Alloy 600 and

Alloy 82/182 from Field Components and Laboratory Materials Tested in

PWR Environments

5. AUTHOR(S)

B. Alexandreanu, O. K. Chopra, and W. J. Shack

6. TYPE OF REPORT

Technical

7. PERIOD COVERED (Inclusive Dates)

February 2007

8. PERFORMING ORGANIZATION - NAME AND ADDRESS (If NRC, provide Division, Office or Region, U.S. Nuclear Regulatory Commission, and mailing address; if contractor,

Argonne National Laboratory

9. SPONSORING ORGANIZATION - NAME AND ADDRESS (If NRC, type "Same as above"; if contractor, provide NRC Division, Office or Region, U.S. Nuclear Regulatory Commission, 

Division of Fuels, Engineering and Radiological Research

provide name and mailing address.)

and mailing address.)

10. SUPPLEMENTARY NOTES

Samantha Crane and Hipolito J. Gonzalez, NRC Project Managers

11. ABSTRACT (200 words or less)

In light water reactors, components made of nickel-base alloys are susceptible to environmentally assisted cracking. This report

summarizes the crack growth rate results and related metallography for field and laboratory-procured Alloy 600 and its weld

alloys tested in pressurized water reactor (PWR) environments. The report also presents crack growth rate (CGR) results for a

shielded-metal-arc weld of Alloy182 in a simulated PWR environment as a function of temperature between 290°C and 350°C. 

These data were used to determine the activation energy for crack growth in Alloy 182 welds. The tests were performed by

measuring the changes in the stress corrosion CGR as the temperatures were varied during the test. The difference in

electrochemical potential between the specimen and the Ni/NiO line was maintained constant at each temperature by adjusting

the hydrogen overpressure on the water supply tank. The CGR data as a function of temperature yielded activation energies of

252 kJ/mol for a double-J weld and 189 kJ/mol for a deep-groove weld. These values are in good agreement with the data

reported in the literature. The data reported here and those in the literature suggest that the average activation energy for Alloy

182 welds is on the order of 220-230kJ/mol, higher than the 130 kJ/mol commonly used for Alloy 600. The consequences of

using a larger value of activation energy for SCC CGR data analysis are discussed.

12. KEY WORDS/DESCRIPTORS (List words or phrases that will assist researchers in locating the report.)

Crack Growth Rate

Stress Corrosion Cracking

Alloy 600

Ni-base welds

PWR Environment

14. SECURITY CLASSIFICATION

13. AVAILABILITY STATEMENT

unlimited

(This Page)

unclassified

(This Report)

unclassified

15. NUMBER OF PAGES

16. PRICE

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