Alloy 827182 Welds Paper

DEVELOPMENT OF CRACK GROWTH RATE DISPOSITION CURVES FOR PRIMARY WATER STRESS CORROSION CRACKING (PWSCC)

OF ALLOY 82, 182, AND 132 WELDMENTS

G. A. White1, N. S. Nordmann1, J. Hickling2, C. D. Harrington31Dominion Engineering, Inc., 11730 Plaza America Drive, Suite 310, Reston, VA 20190

2EPRI, 3412 Hillview Avenue, Palo Alto, CA 94304 3TXU Energy, P.O. Box 1002, Glen Rose, TX 76043

Keywords: Nickel-base weld metals, Alloy 82, Alloy 132, Alloy 182, Primary Water Stress Corrosion Cracking, Crack growth rate

Abstract

Nickel-based austenitic alloys, including wrought Alloy 600 and Alloy 82/182/132 weld metals, are used extensively in pressurized water reactor (PWR) applications. In 2003, the authors reported the results of work sponsored by the EPRI Materials Reliability Program (MRP) to develop a crack growth rate (CGR) disposition curve for primary water stress corrosion cracking (PWSCC) of thick-section Alloy 600 material. This deterministic CGR equation has been adopted by Section XI of the ASME Boiler & Pressure Vessel Code for continued-service evaluation of PWSCC flaws detected (or postulated to exist) in PWR reactor vessel upper head nozzles, including control rod drive mechanism (CRDM) nozzles. Following observations of cracking in primary circuit welds with high residual stresses and in some J-groove welds attaching CRDM nozzles to the reactor vessel upper head, the need for a similar equation for Alloy 82/182/132 weldments was identified. A preliminary MRP CGR curve for Alloy 182 material was published in 2000, but this was based on a fairly limited experimental database and simplifying assumptions. Weld metals are by definition as-cast structures and, as such, are much more inhomogeneous than wrought materials. The scatter introduced by the inhomogeneous nature of weld metals necessitated the development of a more sophisticated approach. Analogous to the procedure that resulted in the CGR equation for Alloy 600 wrought material, an international panel of PWSCC experts supported the MRP in its development of a deterministic CGR equation for Alloy 82/182/132 weldments. After reviewing the key metallurgical aspects of Alloys 82, 182, and 132, the data and methods used to develop the CGR equation for such weldments are described. The laboratory testing techniques that have been used to generate CGR data for these weld metals in simulated PWR primary water environments were analyzed. Appropriate screening procedures were developed and applied to produce the final MRP database before using an agreed data reduction methodology to derive separate CGR curves as a function of temperature and stress intensity factor KI for these weld metals, including consideration of the effects of dendrite orientation. For stress intensity factors greater than 20 MPam, the new CGR curve for Alloy 182/132 weld metal is nearly parallel to, and about four times higher than, the previously reported curve for Alloy 600 wrought material. Comparisons are made with other laboratory data not used in derivation of the new MRP lines, with the limited field data available from repeat non-destructive examination inspections of a cracked primary circuit butt weld at the Ringhals PWR in Sweden, and with the CGR disposition curves that have been proposed by other workers.

Finally, an example is provided of the way in which the curves can be applied to the assessment of further growth through PWSCC of piping butt weld flaws that might be detected in service.

I. Introduction

Nickel-based austenitic alloys, including wrought Alloy 600 and weld metals Alloy 82, 182, and 132, are used extensively in pressurized water reactor (PWR) applications. These materials offer a useful combination of good mechanical properties and fracture toughness, compatibility with other vessel or piping materials, and corrosion resistance. However, recent incidents of primary water stress corrosion cracking (PWSCC) of Alloy 600 components other than steam generator tubes in the primary circuits of PWRs [1] have highlighted the need for a qualified equation for crack growth rates (CGRs) to evaluate flaws found by in-service inspection. This requirement was fulfilled for the wrought Alloy 600 base material, after much deliberation involving an international panel of PWSCC experts, by the issuance in 2002 of the Materials Reliability Program (MRP) MRP-55 report [2], whose main contents were later published as Reference [3]. The disposition curve established in that work has since been incorporated into the ASME Section XI Code for flaw evaluation [4]. A similar requirement has also been identified for Alloy 82/182/132 weldments following observations of cracking in primary circuit welds with high residual stresses and in some J-groove welds attaching control rod drive mechanism (CRDM) and bottom mounted instrumentation (BMI) nozzles to the reactor upper head [1,5]. A preliminary MRP CGR curve for Alloy 182 material was published as a proprietary report in 2000 [6] and later made public [7], but this was based on a fairly limited experimental database and an assumption that the results could be described by application of a simple multiplication factor to the then current Scott model for base metal, which had been derived from field data on thin-walled steam generator tubing [8]. Weld metals are by definition as-cast structures and, as such, are much more inhomogeneous than wrought materials. The scatter introduced by the inhomogeneous nature on a microscopic scale of weld metals makes the simple multiplication factor approach not suitable for extensive use, and necessitated the development of a more sophisticated methodology. The approach taken here to deriving a more appropriate model for nickel-based weld metals was analogous to that used to establish the MRP-55 curve for thick-walled wrought Alloy 600, namely detailed consideration and screening by the MRP PWSCC Expert Panel of all available laboratory data from relevant CGR

Proceedings of the 12th International Conference onEnvironmental Degradation of Materials in Nuclear Power System Water Reactors

Edited by T.R. Allen, P.J. King, and L. Nelson TMS (The Minerals, Metals & Materials Society), 2005

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tests in simulated PWR water and statistical derivation of best-fitcurves, taking into account, as far as possible, the particular natureof PWSCC in weld material and its possible influence on theexperimental results that had been obtained. The present articlestarts by describing key metallurgical aspects of Alloys 82, 182, and 132. It continues with a description of the laboratory testingtechniques that have been used and details the screeningprocedures that were applied to produce the final MRP database.After setting out the data reduction methodology used to deriveseparate CGR curves as a function of the stress intensity factor KIfor nickel-based weld metals, comparisons are made with otherlaboratory data not used in derivation of the new CGR lines, thelimited field data available from repeat non-destructiveexamination (NDE) inspections of a cracked primary circuit buttweld at the Ringhals PWR in Sweden, and with the CGRdisposition curves that have been proposed by other workers. Finally, an example is provided of the way in which the curvescan be applied to the assessment of further growth of cracks byPWSCC which might be detected in service. Report MRP-115 [9]presents the full methodology and results of the MRP study ofCGRs for PWSCC of Alloy 82/182/132 weld metals.

II. Background on Metallurgical Aspects of Nickel-BasedWeld Materials and Their Effects on Crack Growth Rates

The incentive for covering metallurgical aspects is that highvariability is observed in the measured CGRs of Alloys 82, 182,and 132, and it is therefore important to understand howmetallurgical factors contribute to this variability. However, while discussion of these factors provides useful backgroundinformation, firm correlations between metallurgical factors andCGR are not available except for differences in CGR between Alloys 182/132 and Alloy 82. For this reason, the CGRs fordifferent weld alloys need to be addressed on a statistical basis,and not by correlation with specific metallurgical features.

II.A Macrostructural and Microstructural Features of Nickel-Based Weld Metals

The CGR in Alloy 82/182/132 is strongly affected by themicrostructure of the weld and by the orientation of the crack growth with respect to the microstructure. For this reason, it is important to develop an understanding of the main features ofAlloy 82/182/132 microstructure.

Weld metal forms by solidification from a molten state,which leads to the formation of dendrites growing in the directionof the heat flow, i.e., perpendicular to the solid material on whichthe weld is deposited. Most welds are made with multiple passes.The grain structure of dendrites in subsequent passes is normallyrelated to that of previous passes as a result of epitaxy, i.e., by thetendency of a crystal forming on a substrate to have the same structural orientation as the substrate. This results in the dendritespersisting through several or many weld passes. Themacrostructure of a typical weld is shown in Figure 1 [9]. Theweld shown in Figure 1 was made with Alloy 82H (Alloy 82H is the same as Alloy 82 but with C content of 0.030.10wt% instead of 0.10wt% max.). Features to note regarding Figure 1 include: The weld was made with over 30 weld passes. There is a strong pattern of columnar grains formed by

dendrites, and the pattern persists through many weld passes.

The dendrites tend to be perpendicular to the base material atthe weld-base material interface, and tend to become vertical(root to crown direction) as the weld thickness increases.The dendrites are mainly vertical in the central region of theweld.

Crown

Root

1 mm

Fig. 1. Transverse Section of Alloy 82H Weld ShowingColumnar Grain Structure [9]

The typical weld structure is shown schematically in Figure 2 [10], which reflects growth of the dendrites perpendicular to thegroove wall and intersecting near the middle of the weld, wherethey form a high energy solidification grain boundary. This situation is somewhat different than that shown in Figure 1 where,because of different groove geometry, the dendrite growth wasmore nearly vertical.

Fig. 2. Sketch Showing Various Kinds of Grain Boundaries in Weld Metals (Courtesy of Ohio State University [10])

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As the weld metal forms, parallel bundles of dendrites with nearly identical crystallographic orientation form and grow intothe melt. The boundaries between these similarly oriented dendrites are called solidification subgrain boundaries (SSGBs)and tend to have low angular mismatches, as well as low energy,and are believed to form paths for PWSCC relatively infrequently.Where different bundles of dendrites intersect or overlap, largerangular mismatches often occur between the grains of the bundles. In this case, the resulting grain boundaries, termed solidificationgrain boundaries (SGBs), can be high energy and are believed to be more common paths for PWSCC. However, it has been found that some high-angle grain boundaries or sections of boundaries have relatively low energy since they have coincident site lattices(meaning that the crystal orientations are such that the atomicstructures of the two grains have a significant level of matching);this causes those particular high-angle boundaries to be relatively more resistant to PWSCC propagation [11]. Such low-energy,high-angle grain boundaries may be present in cases for which thecrack path is kinked, connecting two planes that are offset by asmuch as 12 mm (even with high-angle boundaries in closeproximity that would have facilitated a straighter crack path).

A typical wavy pattern of high energy grain boundaries isobserved in Alloy 82 and 182 weld metals. The structure of the grain boundaries is illustrated in Figure 3 [10], which also shows migrated grain boundaries (MGB) that develop as the result ofsubsequent weld passes. In multipass welds, the SGBs canmigrate on cooling after solidification and during re-heating andresult in a straighter, migrated grain boundary. Visible in this figure are SSGBs formed between dendrites, SGBs, and MGBs.PWSCC cracks in weld metals typically follow the higher energySGB and/or MGBs.

The orientation of the crack in the weld, i.e., relative to thewelds columnar microstructure, has a strong influence on theCGR. Thus, it is necessary to include the relative crackorientation (i.e., parallel or perpendicular to the predominantdirection of the weld dendrites) in the development of a CGRmodel. The convention used for identifying crack orientation is shown in Figure 4.

Cracks grow fastest along high energy grain boundaries inthe direction of grain growth (TS and LS orientations), and nextfastest along high energy grain boundaries perpendicular to thedirection of grain growth but parallel to the welding direction.Cracks that grow perpendicular to the high energy grain

boundaries, i.e., perpendicular to the columnar dendrites, growsignificantly slower.

SSGB

SGB

MGB

Fig. 3. Micrograph Showing Solidification (SGB), SolidificationSubgrain (SSGB), and Migrated (MGB) Grain Boundaries in

Weld Metals (Courtesy of Ohio State University [10])

II.B Effects of Chemical Composition on Crack Growth Rate

The nominal chemical compositions of Alloys 82, 182, and132 are shown in Table I.

The only well explored effect of the compositionaldifferences among the weld alloys on PWSCC is the influence ofchromium. Buisine, et al. evaluated the PWSCC resistance ofnickel-based weld metals with various chromium contents rangingfrom about 15% to 30% chromium [12]. The results indicatedthat weld metals with 30% chromium were resistant to cracking,with a threshold for PWSCC resistance being between 22 and30% chromium.

Crack growth is along (parallel to) the direction of the dendrites for the TSand LS orientations.

Crack growth is across (perpendicularto) the direction of the dendrites for the TL, LT, ST, and SL orientations.

Nomenclature for crack orientationThe first letter denotes the direction normal to the planeof the crack face. The second letter denotes the direction of crack growth.

Fig. 4. Terminology Used for Orientations of Cracks in Test Specimen with Respect to Welds

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Table I. Nominal Chemical Composition for Alloy 82, 182, and 132 Weld Metal

wt. % Alloy

Ni Cr Fe Mn Nb Ti

82 71 20 2 3 2.5 0.5

182 67 15 8 7 1.8 0.5

132 70 15 9 1 2.5 ---

Other laboratory investigations related to compositional effects have been reviewed and the following conclusions can be drawn [9]: (1) increasing the average chromium concentration of the material correlates with increasing resistance to PWSCC, (2) the reduction in local chromium concentration that occurs at grain boundaries as the result of exposure to sensitizinga heat treatments does not increase the materials susceptibility to PWSCC, and the reduction of residual stresses provided by certain heat treatments is helpful, and (3) there are no known consistent effects of impurities (Si, P, and S) on PWSCC susceptibility.

II.C Effect of Weld Design and Fabrication on Crack Growth Rate

Weld design and fabrication can affect CGR in weld metals in several ways, such as by their effects on residual stresses, local material composition, strength level of the weld material, microstructure, and presence of micro-flaws [9]. Residual Stresses. It is considered that residual stresses

associated with welding can have a strong influence on PWSCC CGR because residual stresses often are a strong contributor to the stress intensity factor at crack tips within the welds and laboratory tests have shown the crack-tip stress intensity factor to be a key parameter.

Local Material Composition. Depending on welding conditions, the chemical content of regions of a weld adjacent to a base metal with lower alloy content may be significantly affected by dilution from the base metal.

Strength Level. The effect of strength level on CGR of weld metals is expected to be similar to that of wrought material, and crack growth rates are expected to increase as the materials yield and tensile strengths rise.

Microstructure. The major microstructural feature of Alloy 82/182/132 consists of dendritic grains, as discussed above. PWSCC in those welds involves an intergranular (IG) cracking mechanism, whereby cracks propagate along high-angle grain boundaries [13,14,15]. Cracks in welds usually have an undulating or wavy character that reflects the wavy morphology of the grain boundaries. In tests at constant load or displacement, unbroken ligaments often form in the wake of advancing crack fronts because the most SCC-resistant boundaries tend not to fail. In some regions, uncracked ligaments can be massive and extend back to the fatigue precrack commonly produced in the specimen for CGR testing, thereby resulting in incomplete engagement of the

a Sensitization refers to the precipitation of chromium carbides leading to low chromium concentration at grain boundaries, making the material susceptible to rapid corrosion in acid-oxidizing environments.

stress corrosion crack to the precrack over the full width of the specimen. Such uneven crack fronts and incomplete engagement of stress corrosion cracks are sources of potential uncertainty in making laboratory measurements of CGR for welds.

Weld Defects. Micro-fissures and other weld defects such as pores and slag inclusions are often present in Alloy 82/182/132. The PWSCC CGR in the weld metal could plausibly be affected by latent defects. However, recent investigations [11,13,14] indicate no discernable effect of hot or ductility-dip cracking on PWSCC. On the other hand, relatively large and sharp defects, such as some lack of fusion areas, could potentially promote PWSCC by acting as stress concentrators and increasing the local stress intensity factor.

III. Specimen Manufacture and Crack Front Patterns

Special purpose welds are typically fabricated to make compact tension (CT) specimens for conducting crack growth rate tests for nickel-based alloy welds used in PWR applications. The majority of the specimens included in the Alloy 82/182/132 crack growth rate database were manufactured using a single-sided, V-shaped, butt weld preparation similar to that commonly used in plants. The welding parameters corresponded to normal industry practice, with attention given both to the weld macro- and microstructure (as described above) and to several additional factors discussed below.

The crack growth rate data applied directly in the development of the MRP-115 CGR equation [9] were obtained from CT specimens having a width of 0.50.6 inches except for the 1-inch CT specimens used by Studsvik [9]. Practice varied regarding the extent of Alloy 600 base metal present in the CT specimens away from the Alloy 82, 182, or 132 crack plane.

III.A Weld Chemical Composition

Test specimen fabrication utilized typical vendor fabrication practices and ASME/AWS specified weld metals and are thus considered representative of weld materials in operating PWRs.

III.B Convention for Identifying Crack Orientation

Because CGR is strongly affected by the direction of crack orientation relative to the microstructure (as discussed above), it is important during tests to identify and control the direction of crack growth relative to the microstructure. The convention used for identifying crack orientation relative to weld fabrication, and thus relative to weld microstructure, is shown in Figure 4, which is an extension of the standard convention for rectangular sections of wrought material [16].

III.C Restraint

Nickel-based alloys and the other structural materials used with these welding materials start to liquefy when heated in the range from approximately 1350 to 1450C. When welds start to cool from these elevated temperatures, significant levels of thermal displacements are involved which can lead to considerable stresses and strains. These stresses and strains depend on the degree of mechanical constraint in the weld and are

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usually sufficient to plastically deform the underlying weld metalas well as the base metal.

A J-groove weld will normally be much more constrainedthan a butt weld. The resulting solidification-strain-inducedstresses are considered likely to make the material moresusceptible to stress corrosion cracking, and it is also possible thatthe weld shrinkage strains increase the strength of the weld metalsomewhat and thereby also have some effect on CGR. Estimatesof the residual stresses present in and adjacent to the weld can be determined by various techniques such as x-ray diffraction orfinite-element analyses, and estimates of material strengtheningcan be obtained from mechanical tests, such as hardness andtensile tests.

The stresses and strains induced during the welding process take the form of macroscopic, local residual stresses and strains aswell as microscopic strains at the grain boundaries. Typically, theremoval of the CT specimen from the sample weld is expected to relieve most of the macroscopic residual stresses in the sample,although the extent of such relief might be expected to depend onthe volume of base metal material remaining in the CT specimen.

In laboratory CGR testing, it is standard practice for the levelof residual stress remaining in the test specimen after its removalto be ignored when estimating the stress intensity factor applied tothe growing crack. However, for the constrained geometry typicalof plants applications, welding residual stresses often dominatethe other stresses and cannot be ignored in crack growth calculations.

III.D In-Process NDE

The ASME Boiler and Pressure Vessel Code setsrequirements for in-process NDE of PWR pressure boundarywelds. Because of the favorable laboratory conditions under which weld test samples are fabricated, it is considered that theirquality may, in general, be better than that of typical plant welds.However, even under ideal welding conditions some defects areexpected, and the presence of defects in some laboratory testwelds has been reported.

III.E Weld Defects

Nickel-based welds made with Alloys 82/182/132 can be affected by various forms of solidification cracking, liquationcracking, and ductility-dip cracking during manufacture. Modifications in welding consumables and procedures have verymuch improved resistance to solidification and liquation cracking.However, processes such as ductility-dip cracking can be expected to produce subsurface weld defects or surface defectssmall enough to be accepted for service during pre-service NDE. It is seemingly plausible that weld defects such as hot cracking or ductility-dip cracking could affect the CGR in those welds.However, recent investigations appear to provide convincingevidence that such weld defects do not play a significant role inPWSCC initiation and propagation [11,13,14].

Relatively large and sharp weld defects such as some weld lack of fusion regions may have the potential to promote PWSCCby creating a local stress concentrator and a high local crack-tipstress intensity factor. Lack of fusion areas at the weld wettedsurface would be expected to be detected during pre-service NDE.Subsurface defects would necessarily have to become wetted bythe primary coolant through some cracking process before theycould grow via PWSCC. Potential types of cracking to cause a

subsurface lack of fusion region to become wetted include ductiletearing, environmental or mechanical fatigue, and PWSCC (crackgrowth in from the wetted surface). Although there is notuniversal agreement among experts, it is possible that at leastsome of the cracking observed in BMI nozzles at South Texas Project Unit 1 in 2003 may have involved the wetting ofsubsurface weld lack of fusion areas [11,17,18].

III.F Post-Welding Heat Treatment

Post-weld heat treatment (PWHT) of the buttering used in primary weld preparations is standard practice, and PWHT is in some cases also performed on the filler metal following finalwelding. In addition, PWHT of some nickel-based weld metalsoften occurs indirectly as a consequence of stress relief heat treatments performed on adjacent low-alloy steel components per ASME Code requirements. In these latter cases, the stress relieftemperature is well below that which would be optimum for nickel-based alloys.

Le Hong et al. [19] report that CGRs are reduced by a factorof 2.0 for stress relieved specimens compared to otherwise similaras-received specimens.

No stress relief was applied to the test welds for all of the testspecimen crack growth rate data used directly in the developmentof the deterministic CGR model for Alloy 82/182/132 welds in thepresent study.

III.G Crack Front Patterns

As mentioned earlier, CGR specimens for Alloy 82/182/132weld metals often exhibit irregular crack fronts, with regions ofnon-engagement (no SCC crack initiation) and with large differences in the extent of SCC crack growth. Photomicrographsof example fracture surfaces that illustrate some of the variety ofcrack front types that can occur are shown in Figures 5 and 6.These figures reflect test specimens that are included in the MRPdatabase. Figure 5 illustrates a case where the crack front ismoderately well behaved, i.e., it is not very irregular. Figure 6 illustrates a case with a highly irregular crack front. Theprocedure adopted for evaluating CGRs in the case of uneven crack fronts is described later in Section IV.C.

Fig. 5. Example of a Fracture Surface of Alloy 182 Weld Metal with Moderately Uniform Crack Front

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Fig. 6. Example of a Fracture Surface of Alloy 182 Weld Metal with Highly Irregular Crack Front

IV. Crack Growth Rate Testing Techniques and Consideration of Incomplete Engagement to SCC

Key testing techniques regarding specimen loading and thetest environment are discussed below, followed by a discussion of the effect of incomplete engagement to stress corrosion crackgrowth across the specimen width.

IV.A Specimen Loading

The welding specimens applied directly to develop the MRP-115 CGR model [9] reflect a number of loading variables,including stress intensity factor, cyclic loading parameters, andweld orientation. Test results in the database (after the screeningprocess) reflect the following: Stress intensity factors ranging from 19.7 to 60.0 MPam for

Alloy 182/132 specimens and from 28.0 to 56.8 MPam for Alloy 82 specimens.

A combination of purely constant loading and periodic,partial, cyclic unloading. Cyclic loading parameters includeload ratios, R, between 0.65 and 0.75 and hold times between3600 and 100,000 s.

Three of the six possible weld orientations (TS, LS, and TL).TS and LS represent orientations where crack growth isparallel to the direction of the weld dendrites, and TLrepresents an orientation where crack growth is across(perpendicular to) the direction of dendrite growth. During CGR testing, addressing the following issues can

yield more consistent and meaningful test results: Application of a stress intensity factor within accepted limits.

All of the laboratories that contributed CGR data to the MRP database applied linear elastic fracture mechanics (LEFM)validity criteria (e.g., ASTM E399 [16] and E647 [20]).

Use of side grooving to maintain the crack plane. Limiting large variations in the stress intensity factor during

constant load testing. Use of periodic, partial cyclic unloading to maintain a

straight crack front.

Appropriate procedures for precracking the specimen. A transgranular fatigue precrack is generated first in order to provide a sharp, linear initiation site for the PWSCC growth mechanism. The length of the precrack, and loading detailsassociated with generation of the precrack, must be considered. In-situ transitioning to an intergranular crackgrowth path is sometimes used to encourage development of a uniform SCC crack front.

Use of continuous crack monitoring, which is a valuable toolthat can: 1) aid in determining when SCC initiated from thefatigue precrack, and 2) assist in estimating CGRs during different phases of multi-condition testing. The mostcommonly used technique for such monitoring is reversedDC potential drop.

Careful control and documentation of machining, surface condition, and pre-oxidation in high-temperature water (e.g.,to obtain proper corrosion potentials).

Control of test temperature, with stability ideally withinr0.5C.

Control and monitoring of water purity and dissolved gaschemistry.

Measurement of the corrosion potential of the CT specimenitself and a separate platinum electrode.

Table II. Environments Used for Crack Growth Rate Tests (MRPDatabase after Screening)

Li BTest Org. Weld ID

Tem

p.(

C)

(ppm)

Diss.H2

(cc/kg)(Note 3)

Ref.

Westinghouse D545/D582,33644,PP751

323to

342

2.0 1200 25 [21]

26B2 343,345

2.0,2.2

1202to

1273

28.8to

30.4

[22]

6892 343 2.2 1212 29.5 [9]

Studsvik

WC05F8 319 2.28 1297 29.6 [9]

Bechtel Bettis A-1,C-1, C-2, C-3, C-4

338 Note 1 50 [13]

LM 182-1 328 Note 2 35 [23]

LM 182-2 338 Note 2 40 [23]

LM 82-1 360 Note 2 40 [23]

LM 82-2 338 Note 2 20, 40 [23]

LockheedMartin KAPL

LM 82-3 316to

360Note 2

30to40

[23]

MHI MG-7,132 Heat

325 3.5 1800 30 [15]

Notes:1) High-temp. hydrogenated water with a room-temp. pH of 10.1 to 10.32) High-temp. hydrogenated water with a high-temp. pH of 6.63) The unit for dissolved hydrogen of cc(standard temperature and pressure (STP))/kg H2O is abbreviated as cc/kg

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IV.B Test Environment

The environments for the tests that were directly used to develop the MRP-115 CGR model are shown in Table II. All of the environments were hydrogenated high-temperature water, often with lithium and boric acid additions to simulate the PWR primary coolant environment. Concentrations of impurities such as chlorides and sulfates were limited to low levels. The temperature range covered was 316 to 360C. Because of the sensitivity of PWSCC CGRs to the hydrogen concentration, special care was taken to screen out data from tests where careful control of hydrogen was not demonstrated.

IV.C Derivation of an Appropriate Crack Growth Rate Considering Uneven Crack Fronts

Alloy 82/182/132 weld specimens tend to exhibit uneven stress corrosion cracks, which introduce uncertainty into crack growth rate measurements. For specimens with small to intermediate amounts of SCC, it is common for intergranular cracking in these materials to incubate along the transgranular (TG) fatigue precrack front non-uniformly versus test time, so that only portions of the crack front may exhibit IG cracking in limited duration tests. Even when intergranular cracks are almost fully engaged, fingers of SCC often jut out beyond the overall crack front, thereby demonstrating the heterogeneous nature of SCC growth in these weld metals. The incomplete engagement and uneven crack front issues are important because they can introduce a bias into CGR measurements. Because these issues are complex, there is no single approach for assuring that any bias is removed from the CGR database. Specific analysis methods designed to account for incomplete engagement and uneven crack fronts were proposed and debated in detail at the EPRI-MRP CGR Expert Panel meetings. Fundamental differences in expert views involve the underlying cause of the absence of crack incubation in some regions of weld metal. At present, it is not known definitively if the non-incubated regions are associated with pinned transgranular fatigue precracks, with microstructural regions of enhanced SCC resistance, or with a combination of both effects. These differences are potentially important because they govern how unengaged and shallow crack extension regions are treated when calculating average CGR. It is noteworthy, however, that the issue of engagement is quantitatively less important for average SCC crack lengths of more than about 2 mm based on a comparison of average and maximum crack extensions for CGR tests for which both these parameters were available [9]. While no consensus was reached concerning methods that address the incomplete engagement and uneven crack front issues for future testing, a consensus was reached on how to address this issue in analyzing the existing CGR database. Specifically, there was agreement that including the zero crack extension values within the unengaged portion of the precrack produces non-conservative CGR estimates. To avoid this problem, the EPRI-MRP CGR Expert Panel agreed that it is appropriate that the average crack extension used to compute CGR be based solely on the engaged segments of the stress corrosion crack (i.e., a simple average of all non-zero crack extension values across the specimen width). In addition, data points with less than 50 percent engagement and less than 0.5 millimeters of average crack extension were excluded from the screened database.

V. Development of the Screened MRP Database and Derivation of CGR Disposition Curves

After a screening process was applied to the set of worldwide laboratory CGR data for Alloy 82/182/132 collected by the MRP, a multiple linear regression statistical model was applied in order to derive recommended deterministic CGR disposition curves for these weld metal materials.

V.A Screening Criteria

The starting point for screening the available stress corrosion crack growth database for nickel-based welds Alloy 82/182/132 was the same as that adopted for the earlier MRP-55 [2,3] study of Alloy 600. The EPRI expert panel for PWSCC revised those screening criteria in consideration of the issues that are particularly relevant to weld metals, and Table III lists the key factors that were considered during the screening process for the Alloy 82/182/132 weld CGR data. It should be noted that the main reasons leading to exclusion of Alloy 600 data from further consideration in the MRP-55 study were: No measurable growth. Less than 50% of crack front with IGSCC initiated (hereafter

called engagement) or lack of crack front mapping to enable this feature to be assessed and average growth rates to be calculated in addition to the maximum rates supplied.

Out of specification PWR primary water chemistry (particularly hydrogen).

Cyclic or ripple loading with less than 1 hour hold time at constant load during each cycle.

A similar pattern emerged for the nickel-based weld metal Alloys 82/182/132 with the addition of a few instances of data rejected because of loading beyond LEFM criteria. However, the second criterion listed above (relating to crack engagement) assumed much higher importance for the weld metals. Due to difficulties with lack of uniform crack initiation from starter fatigue cracks and the development of irregular crack fronts, an additional requirement to that of greater than 50% engagement used here was a minimum crack growth increment averaged across the specimen width (aave) of at least 0.5 millimeters. A sensitivity study established that the precise choice of the aavecutoff used as the screening criterion did not have an arbitrary influence on the acceptable screened data set and the eventual outcome of the data analysis [9]. Please see Appendix A to this article for further details regarding the screening process. A detailed treatment of hydrogen effects, which are known to be potentially significant for crack growth in nickel-based weld metals [23], was not possible with the limited number of results contained in the screened database [9]. It was noted, however, that one set of KAPL data which apparently illustrates a significant KI dependency actually resulted more from testing at two distinct H2 levels [23].

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Table III. Key Factors for Consideration in CGR Testing and Data Reporting

1 Material within specifications including composition/condition/heat treatment

2 Mechanical strength properties

3 ASTM specimen size criteria and degree of plastic constraint

4 Pre-cracking technique (including straightness criteria, plastic zone size, crack morphology)

5 Special requirements for testing welds (e.g. pre-crack location, residual stresses/strains)

6 Environment (chemistry, temperature, electrochemical potential (ECP), flow rate at specimen, neutron/gamma flux)

7 Loop configuration (e.g., once-through, refreshed, static autoclave)

8 Water chemistry confirmation by analysis (e.g., Cl, SO4, O2, Cr, total organic carbon (TOC), conductivity)

9 Active constant or cyclic loading versus constant displacement loading (e.g., using wedge)

10 On-line measurement of crack length versus time during test (including precision)

11 Actual crack length confirmed by destructive examination (assessment method/mapping)

12 Appropriateness of crack characteristics (fraction SCC along crack front, uniformity, adequate SCC increment, transgranular portions within IGSCC fracture surface, etc.)

13 Possible effects of changes in loading or chemistry conditions during a test (including heat up and cool down)

14 Calculation and reporting of K or K values

15 Reporting of raw a vs. t data and derivation of da/dt values

16 Reproducibility of data under nominally identical test conditions

V.B Development of MRP Database for Alloy 82/182/132

After the screening process described above was applied, a multiple linear regression statistical model was applied to the MRP database of CGR data for Alloy 82/182/132 to develop theMRP-115 deterministic CGR model. The MRP expert panel concluded that Alloy 182 and Alloy 132 can be regarded assufficiently similar to be described by one CGR curve.

Figure 7 is a CGR versus stress intensity factor plot showingthe complete set of available data for which average CGRs werereported, adjusted to a common reference temperature of 325C assuming a thermal activation energy of 130 kJ/mole (31.0 kcal/mole). This is the same activation energy that was applied tothe CGR data for Alloy 600 in MRP-55 [2,3]. The expert paneljudged that there were insufficient data to develop reliableactivation energy values for Alloy 182/132 and for Alloy 82, sothe accepted activation energy value for Alloy 600, which has a similar composition, was used. Multiple independent studies ofAlloy 600 have resulted in thermal activation energy valueswithin about 1015% of the value of 130 kJ/mole (31.0kcal/mole) [2,3].

Figure 8 is the corresponding plot for the available data forwhich CGRs based on the maximum crack increment across thespecimen width were reported. Figures 9 and 10 show theaverage CGR data in the MRP database following the screeningprocess: Figure 9 shows the data for Alloys 182 and 132, and Figure 10 shows the data for Alloy 82. Note that for referencepurposes, two previously developed CGR curves are shown in Figures 7 through 10.

V.C Data Reduction

The statistical methodology for developing the deterministicCGR equation for Alloy 82/182/132 weld metal is describedbelow. The procedure is similar to that presented in MRP-55[2,3] for Alloy 600 wrought material, but includes a linearizedmultiple regression model in order to determine a best-fit stressintensity factor exponent, , while still treating the data on a weld-by-weld basis:1. Collect data including reported initial or average K, CGR

based on the crack increment averaged across the entirespecimen width, average crack increment, test temperature,and percentage engagement of the crack front to IGSCC(%eng). Average CGR data were used, rather than themaximum measured CGR across the specimen width,because it is believed that the average CGR is a bettermeasure of the fundamental material behavior, whereas themaximum CGR is more dependent on the spatial variabilityin resistance to PWSCC. In addition, the maximum CGR appears to be more dependent on test duration than the average CGR [9]. Finally, it is standard practice in fatiguetesting of CT specimens to average the crack extensionacross the specimen width [20].

2. Perform data screening using the key factors listed inTable III.

3. Modify the reported CGR to account for the effect of incomplete initiation of PWSCC across the crack front bydividing by the engagement fraction (effectively excludingzero crack extension points from the average across thespecimen width):

'%100

CGRCGReng

(1)

518

1.E-12

1.E-11

1.E-10

1.E-09

1.E-08

0 10 20 30 40 50 60 70 8Stress Intensity Factor, K (MPam)

Cra

ck G

row

th R

ate,

da

/ dt

(m/s

)

0

All data adjusted to 325C (617F)using an activation energy of130 kJ/mole (31.0 kcal/mole)

1mm/yr

All CGRs are reported averageCGRs and are not adjusted toaccount for percentageengagement across the crack front,alloy type, or crack orientation

MRP-55 Curvefor Alloy 600


Fig. 7. Complete Set of Worldwide Alloy 82/182/132 Average CGR Data before Screening Process (144 points)

1.E-12

1.E-11

1.E-10

1.E-09

1.E-08


Cra

ck G

row

th R

ate,

da

/ dt

(m/s

)

0


1mm/yr

All CGRs are reportedmaximum CGRs and arenot adjusted to account foralloy type or crack growthorientation



Fig. 8. Complete Set of Worldwide Alloy 82/182/132 Maximum CGR Data before Screening Process (158 points)

519

1.E-12

1.E-11

1.E-10

1.E-09

1.E-08


Cra

ck G

row

th R

ate,

da/ d

t (m

/s)

0


1mm/yr



All CGRs are adjusted to accountfor percentage engagement acrossthe crack front but not alloy typeor crack orientation

Fig. 9. Average CGR Data for Alloys 182 and 132 after MRP Screening (43 Points)

1.E-12

1.E-11

1.E-10

1.E-09

1.E-08


Cra

ck G

row

th R

ate,

da/ d

t (m

/s)

0

1mm/yr




All CGRs are adjusted to accountfor percentage engagement acrossthe crack front but not alloy typeor crack orientation

Fig. 10. Average CGR Data for Alloy 82 after MRP Screening (34 Points)

520

4. Adjust the data to a common reference temperature of 325C using an activation energy of 130 kJ/mole (31.0 kcal/mole).

5. Assume no stress intensity factor threshold for PWSCC ofthe weld metals (i.e., Kth = 0). The EPRI expert panel for PWSCC concluded that, for the weld metal materials, therewere insufficient data to justify a stress intensity factorthreshold other than zero. See Appendix B to this article for discussion of this assumption.

6. Assume the following form to model the set of screenedaverage CGR data:

'weld alloy orient

temp

CGR f f f Kf

ED (2)

where = power-law constantftemp = factor adjusting CGR to common reference

temperature of 325Cfweld = common factor applied to all specimens

fabricated from the same weld to account forweld wire/stick heat processing and for weld fabrication (see discussion below)

falloy = factor accounting for effect of compositiondifference between Alloy 182/132 and Alloy 82(taken as 1.0 for Alloy 182/132)

forient = factor accounting for difference in CGR resulting from crack growth perpendicular tothe direction of the weld dendrites versus parallel to the direction of the dendrites (takenas 1.0 for the parallel case)

K = crack-tip stress intensity factor = power-law exponent

7. Linearize the assumed form of the CGR equation by taking the natural logarithm of the adjusted CGR.

8. Perform a least-squares multiple linear regression fit treatingthe weld factor (fweld) as a normally distributed random variable.

9. Choose the alloy factor (falloy) for Alloy 82 based on thevalue that makes the log-mean for the set of weld factors forthe Alloy 182/132 welds equal to the log-mean for the set ofweld factors for the Alloy 82 welds.

10. Determine the orientation factor (forient) for crack growth in the direction perpendicular to the weld dendrites (TL, LT,ST, SL) versus growth parallel to the dendrites (TS, LS) based on the best-fit from the regression model.

11. Similar to the procedure in MRP-55, base the deterministicCGR equation on the 75th percentile of the log-normal distribution for the 19 weld factors.

The fweld factor serves the same function as the heat factor(fheat) applied in MRP-55 [2,3] to all specimens fabricated fromthe same heat of Alloy 600 wrought material to account for theeffect of material processing differences on the CGR. The weldfactor is necessary in the statistical treatment of the CGR data to account for the systematic biases associated with particular testwelds. Because of material and fabrication differences, differentwelds (of the same alloy type) will display a range of CGRs evenwhen loading and environmental factors are identical. In practice,the weld factor for a particular application is not known, so asdescribed below, the 75th percentile of the distribution ofcalculated weld factors is adopted for the recommendeddeterministic CGR disposition equation.

Table IV. Calculated Normalization Factors for Alloy Type(82/182/132) and Weld Heat/Processing

WeldRank Alloy

AlloyFactor

falloy(Note 1)

WeldFactor

fweld(Note 1) falloyfweld

1 182 1.00 2.17 2.172 182 1.00 2.12 2.123 132 1.00 1.70 1.704 182 1.00 1.25 1.255 182 1.00 1.15 1.156 182 1.00 0.91 0.917 132 1.00 0.89 0.898 82H 0.38 2.04 0.789 82H 0.38 2.03 0.78

10 182 1.00 0.76 0.7611 182 1.00 0.74 0.7412 82H 0.38 1.54 0.5913 82 0.38 1.32 0.5114 82H 0.38 1.32 0.5115 182 1.00 0.51 0.5116 182 1.00 0.38 0.3817 82 0.38 0.61 0.2418 82H 0.38 0.47 0.1819 82 0.38 0.31 0.12

Note:1Assuming form CGR = ftempfalloyfweldforientK1.6

The linearized multiple regression model fit to the set of 77points in the screened MRP database resulted in the following: the set of 19 weld factors (fweld) tabulated in Table IV and

plotted in Figure 11 an alloy factor (falloy) of 1/2.6 = 0.385 for Alloy 82 an orientation factor (forient) of 0.5 for crack growth

perpendicular to the direction of the dendrites a stress intensity factor exponent () of 1.6 a constant factor of 9.82u10-13

Figure 11 shows the log-normal distribution fit to the set of19 weld factors. Because it is fit to the weld factors, this distribution describes the variability in CGR due to difference inweld wire/stick material heat processing and weld fabrication.The 75th percentile value of this distribution is a weld factor of1.49. For the purpose of producing a single deterministic CGR model, the 75th percentile weld factor is absorbed into theconstant factor, resulting in a value of of 1.510-12.

V.D MRP Disposition Curves

The deterministic CGR curves for Alloy 182/132 and Alloy82 are shown in Figure 12. The MRP database indicates that theCGR for Alloy 82 is on average 2.6 times lower than that forAlloy 182/132, so the MRP-115 curve for Alloy 82 is 2.6 times lower than the curve for Alloy 182/132. For crack propagationthat is clearly perpendicular to the dendrite solidificationdirection, a factor of 2.0 lowering the CGR may be applied to the curves for Alloy 182/132 and Alloy 82.

Figures 13 and 14 show the results of the statistical analysisin comparison with the MRP screened database for Alloys182/132 and 82, respectively. Note that, unlike for Figures 9 and10, the data in Figures 13 and 14 have been normalized for theeffect of crack orientation. The raw CGRs for points for which the crack growth was perpendicular to the dendrite direction have been increased by a factor of 2.0.

521

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

0.1 1. 10.

Weld Factor, f weld

Cum

ulat

ive

Dis

trib

utio

nF

9 182 Welds8 82 Welds2 132 WeldsLog-Normal Fit

Weld factors for 19 welds of Alloy 82/182/132material with fit log-normal distribution(most likely estimator), K th = 0, and best fit E

25th Percentile

75th Percentile

Median

The Alloy 82 data have been normalized(increased) by applying a factor of 2.61:1/f alloy = 2.61

Fig. 11. Log-Normal Fit to 19 Weld Factors for Screened MRP Database of CGR Data for Alloy 82/182/132

1.E-12

1.E-11

1.E-10

1.E-09

1.E-08


Cra

ck G

row

th R

ate,

da/ d

t (m

/s)

The reference temperature for theMRP curves is 325C (617F); therecommended thermal activationenergy for temperature adjustmentis 130 kJ/mole (31.0 kcal/mole),the same value recommended inMRP-55 for base metal.

1 mm/yr

MRP-115 Curve for Alloy 182/132CGR = 1.510-12K 1.6

MRP-115 Curve for Alloy 82CGR = (1.510-12/2.6)K 1.6

For crack propagation that isclearly perpendicular to thedendrite solidification direction, afactor of 2.0 lowering the CGRmay be applied to the curves forAlloy 182 (or 132) and Alloy 82.

MRP-55 Curve forAlloy 600 Base Metal

Laboratory testing indicates thatthe CGR for Alloy 82 is on average2.6 times lower than that for Alloy182/132, so the MRP-115 curvefor Alloy 82 is 2.6 times lowerthan the curve for Alloy 182/132.

Fig. 12. MRP-115 Deterministic Curves for Alloy 182/132 and Alloy 82 Weld Materials

522

1.E-12

1.E-11

1.E-10

1.E-09

1.E-08


Cra

ck G

row

th R

ate,

da/ d

t (m

/s)

0

1mm/yr



MRP-21 Curvefor Alloy 182MRP-115 Curve for

Alloy 182/132da/dt = 1.510-12K 1.6

All CGRs are adjusted to accountfor percentage engagement acrossthe crack front and crackorientation but not alloy type

Fig. 13. Average CGR Data for Alloys 182 and 132 after MRP Screening (43 Points) Normalized to a Crack Orientation Parallel to theWeld Dendrites with MRP-115 Curve for Alloy 182/132

1.E-12

1.E-11

1.E-10

1.E-09

1.E-08


Cra

ck G

row

th R

ate,

da

/ dt

(m/s

)

0

1mm/yr




MRP-115 Curve forAlloy 82da/dt = (1.510-12/2.6)K 1.6

All CGRs are adjusted to accountfor percentage engagement acrossthe crack front and crackorientation but not alloy type

Fig. 14. Average CGR Data for Alloy 82 after MRP Screening (34 Points) Normalized to a Crack Orientation Parallel to the WeldDendrites with MRP-115 Curve for Alloy 82

523

The mathematical form of the MRP-115 CGR curve forAlloy 182/132 at 325C (617F) is:

CGR (in m/s) = 1.510-12 K1.6 (3)(for K in MPam)

CGR (in inches/hr) = 2.4710-7 K1.6 (4)(for K in ksiin)

The general form of the MRP-115 equation is as follows:

1 1exp g alloy orientref

Qa

R T Tf f K ED

(5)

where: a = crack growth rate at temperature T in m/s (or

in/h)Qg = thermal activation energy for crack growth

= 130 kJ/mole (31.0 kcal/mole)R = universal gas constant = 8.314u10-3 kJ/mole-K

(1.103u10-3 kcal/mole-R)T = absolute operating temperature at location of

crack, K (or R)Tref = absolute reference temperature used to

normalize data = 598.15 K (1076.67R)

= power-law constant = 1.510-12 at 325C for a in units of m/s and K

in units of MPam (2.4710-7 at 617F for ain units of in/h and K in units of ksiin)

falloy = 1.0 for Alloy 182 or 132 and 1/2.6 = 0.385 forAlloy 82

forient = 1.0 except 0.5 for crack propagation that is clearly perpendicular to the dendritesolidification direction

K = crack-tip stress intensity factor, MPam (or ksiin)

= exponent = 1.6

The MRP curve may be interpreted as the mean of the upperhalf of the distribution describing the variability in CGR due tomaterial heat, orin this caseindividual weld. Therefore, theMRP curve addresses the concern that welds that are moresusceptible than average to crack initiation tend to have higherCGRs than average. Cracking detected in operating plants would tend to be located in components using such susceptible welds.The use of a conservative mean CGR is consistent with thegeneral approach of Section XI of the ASME Boiler & Pressure Vessel Code, in which mean CGRs are assumed and allowable crack sizes are based on design loads multiplied by load factors.

It is noted that all of the data points which were derived from1 inch CT specimens fall on or below the MRP-115 line for Alloy182even after correction for dendrite orientation. This is incontrast to data points derived from 0.50.6 inch CT specimens,some of which are above the MRP-115 line. However, it is judged that insufficiently diverse data are available to discernwhether specimen size in fact has a significant impact onmeasured CGRs because only the two sets of Studsvik data reflect1 inch CT specimens (while all others reflect CT specimens of 0.50.6 inches in width).

VI. Comparison of MRP Disposition Curves with Other Data

Several comparisons were made of the MRP-115 deterministic equation to the data that were not included in thefinal screened MRP database of laboratory CGR data, as well asto the limited available field data. These comparisons were madein order to verify the robustness of the MRP-115 multiple linearregression model, given the manner in which the data screeningprocess was implemented. These comparisons were alsoperformed to verify the absence of any hidden effects in theoverall set of CGR data collected. The following specificcomparisons and investigations were performed: comparison with lab data for ex-service weld material, comparison with available plant data for Alloy 82/182/132

weld metal based on repeat NDE crack sizing, comparison with available laboratory data investigating the

potential effect of pH, comparison with average CGR data that were excluded from

use in calculating the MRP-115 equation, comparison with maximum CGR data including for tests for

which average CGRs were not available, and investigation of effect of periodic unloading and hold time.

The first two items, and a comparison with previouslypublished CGR curves for Alloy 182, are presented below.

VI.A Comparison with Laboratory Data Generated for RemovedPlant Weld Material

As described in more detail in MRP-113 [24], boric acid crystal deposits led to the discovery of a small hole in the Alloy82/182 butt weld between the low-alloy steel reactor vessel outlet nozzle and stainless steel primary coolant pipe during the October2000 refueling outage at VC Summer. Destructive examinationsrevealed the presence of several axial cracks, including a through-wall axial crack extending essentially the full weld width, as wellas a short, shallow circumferential crack in the Alloy 182 claddingthat arrested when it reached the low-alloy steel nozzle.

Samples of both Alloy 182 butter and Alloy 82 filler materialtaken from this hot leg safe end weld were used in a series ofcrack growth rate tests completed by Westinghouse [25,26]. Thetest conditions included a test temperature of 325C; a simulated primary water environment with 3.5 ppm Li, 1800 ppm B, and3035 cc/kg dissolved hydrogen; fatigue pre-cracking in air at astress intensity factor below 15 MPam; and active loading with anominal test stress intensity factor of either 20 or 35 MPam.

The specimens were periodically unloaded to a load equal to70% of the full applied load (R = 0.7) in order to break any oxidesthat might affect the accuracy of the crack growth measurements.Three cyclic loading test phases and one constant loading testphase were conducted on each of the two Alloy 182 and twoAlloy 82 specimens [9]. Side grooves were included in thespecimens in an attempt to keep cracking in the intended plane.For the Alloy 82 specimens, testing was in the TS direction (crackplane parallel to the dendrites) while for the Alloy 182 specimens,testing was in the TL direction (crack growth perpendicular to thedendrites).

For each of the four samples, post-test fractography was usedto determine the overall crack increment, and this increment wasdivided into four parts on the basis of on-line DC potential dropmeasurements, thereby facilitating four separate data points foreach sample corresponding to the four test phases.

524

1.E-12

1.E-11

1.E-10

1.E-09

1.E-08


Cra

ck G

row

th R

ate,

da

/ dt

(m/s

)MRP-115 Curve for Alloy 182

MRP-115 Curve for Alloy 82


Lab Data for Summer 182-1 and182-2 CT Samples (As Meas.)

Lab Data for Summer 182-1 and182-2 CT Samples (Adj. forOrient.)Lab Data for Summer 82-1 and 82-2 CT Samples

The Summer data were producedat a test temperature of 325C,which is the reference temperaturefor the MRP-115 curves for Alloys182/132 and 82 weld metal

1 mm/yr


The Summer data for Alloy 182are shown with and without anadjustment factor of 2.0 to account forcrack orientation


Fig. 15. Comparison of MRP-115 Curves for Alloys 182/132 and 82 with Westinghouse CGR Data for Weld Material Removed from VC Summer Reactor Hot Leg Safe End Butt Weld [25,26]

The crack growth rate data produced from these tests areplotted in Figure 15 along with the MRP-115 curves for Alloy 182and Alloy 82. As noted in the figure, two sets of data points areincluded for the Alloy 182 specimens: the as-measured CGRs, and CGRs increased by a factor of 2.0 to account for the crackorientation (perpendicular to the dendrites). After the crackorientation correction of the CGRs for the two Alloy 182 weldsamples, these data are still in reasonable agreement with theMRP-115 lines for both Alloy 182/132 and Alloy 82.

The VC Summer data were screened from the database usedto develop the MRP-115 deterministic model because theselaboratory data were generated during multi-condition tests inwhich the loading type was changed.

VI.B Comparison with Available Field Data

During the Ringhals Unit 3 refueling outage in 2000, two axially oriented defects were detected in one of the reactor vesseloutlet nozzle-to-safe-end Alloy 182 butt welds using a qualifiededdy current technique [27,28,29].b During the 2000 outage, thedepth of each defect was measured to be 93 mm and the length1610 mm with ultrasonic testing (UT). After additionaloperation for approximately 8000 effective full power hours, thefirst defect (Crack 1) had grown to a depth of 133 mm while thesecond defect (Crack 2) measured 163 mm, as shown in Table V.

b Note that the root pass of each of the double-V type welds at Ringhals is reported to have been produced using Alloy 82 weld metal. However, the reported cracks did not extend to the root region. Hence, both cracks were located exclusively in Alloy 182 material.

The left portion of Table V lists the initial and final crackdepths and corresponding crack extensions associated with best-estimate, statistical upper- and lower-bound, and worst case crackgrowth. The best-estimate case assumes that the initial and finalcrack depths are subject to no error (or, more precisely, that eachis subject to the same error). For example, for Crack 1, the best-estimate initial depth, final depth, and extension are 9, 13, and 4mm, respectively. The worst-case crack growth assumes that theinitial and final depths are at the extreme values implied by themeasurement uncertainty (e.g., for Crack 1, the initial depthwould have been 93 = 6 mm, the final depth 13+3 = 16 mm, andthe extension 166 = 10 mm). The upper and lower statistical bounds assume that the initial and final depth measurements are independent (i.e., that the measurement errors in each case are notsubject to a common bias). Based on standard engineeringtolerance stack-up assumptions, this implies that the uncertainty inthe measurement difference is equal to 32 = 4.24 mm. If halfof this uncertainty tolerance is assigned to both the initial andfinal depth measurements, the values in Table V for Stat. LowerBound and Stat. Upper Bound are obtained (e.g., initial depthof 92.12 = 6.88 mm, final depth of 13+2.12 = 15.12 mm, and extension of 15.126.88 = 8.24 mm for the statistical upper boundfor Crack 1).

The stress intensity factors that apply at the locations of theRinghals Unit 3 defects were calculated by Efsing and Lagerstrm[27] based on stresses (including welding residual stresses)calculated using the finite-element method and fracture mechanics calculations assuming standard superposition assumptions and arereported in the rightmost portion of Table V.

525

Table V. Data Reported for Ringhals Unit 3 Hot Leg Safe End Nozzle Weld Cracks

AverageCGR (m/s)

Stress Intensity Factor(MPam)

Crack Statistical Case

InitialDepth

a1(mm)

FinalDepth

a2(mm)

Extensiona

(mm)

Oper.at

319C

Adjustedto

325CInitial

K1Final

K2MeanKave

Stat. Lower Bound 11.12 10.88 No Growth No Growth 32.2 32.2 32.2Best Estimate 9.0 13.0 4.0 1.4E-10 1.8E-10 29.5 33.5 31.5

Stat. Upper Bound 6.88 15.12 8.24 2.9E-10 3.7E-10 24.0 35.5 29.71

Worst Case 6.0 16.0 10.0 3.5E-10 4.5E-10 21.0 36.5 28.8Stat. Lower Bound 11.12 13.88 2.76 9.6E-11 1.3E-10 32.3 34.6 33.5

Best Estimate 9.0 16.0 7.0 2.4E-10 3.2E-10 29.5 36.5 33.0Stat. Upper Bound 6.88 18.12 11.24 3.9E-10 5.1E-10 24.0 38.3 31.2

2

Worst Case 6.0 19.0 13.0 4.5E-10 5.9E-10 21.0 39.5 30.3

1.E-12

1.E-11

1.E-10

1.E-09

1.E-08


Cra

ck G

row

thR

ate,

da/ d

t (m

/s)

MRP-115 Curve for Alloy 182/132



Ringhals 3 / Crack 1 / DepthIncrease from 2000 to 2001

Ringhals 3 / Crack 2 / DepthIncrease from 2000 to 2001

1 mm/yr



All curves adjusted to 325Cusing an activation energy of130 kJ/mole (31.0 kcal/mole)

The points for the Ringhals 3 hot leg safe end weld cracks are based on thedepth measurements made in 2000 and 2001 and the stress intensity factorscalculated by Ringhals (points shown at average of initial and final Kcorresponding to best estimate initial and final depths). The Ringhals datawere adjusted from the operating temperature of 319C (606F) to thereference temperature of 325C (617F) using the activation energy of130 kJ/mole (31.0 kcal/mole).

Fig. 16. Field Crack Growth Data for Ringhals Unit 3 Hot Leg Safe End Alloy 182 Weld

1.E-12

1.E-11

1.E-10

1.E-09

1.E-08


Cra

ck G

row

th R

ate,

da/ d

t (m

/s)


Ringhals Two-Part Curve(Adjusted to 325C)

EDF Alloy 182 Curve



All curves adjusted to 325C (617F)using an activation energy of130 kJ/mole (31.0 kcal/mole)

1 mm/yr



MRP-115 Curve forAlloy 182CGR = 1.510-12K 1.6

Ringhals Two-Part Curve @325C

EDF Alloy 182 Curve

Fig. 17. Comparison of MRP-115 Curve for Alloy 182 WeldMetal with Other Disposition Curves

Figure 16 shows the Ringhals Unit 3 field data plotted with the deterministic MRP-115 curve for Alloy 182 weld metal. Thesolid and open squares in Figure 16 represent the best-estimatecrack growth rates for the depth increase of the two Ringhals Unit 3 cracks. These two points have been adjusted to thereference temperature of 325C (617F) using the standardthermal activation energy of 130 kJ/mole (31.0 kcal/mole), and they reflect the stress intensity factors calculated by Efsing andLagerstrm [27]. (The points are shown at the average of theinitial and final stress intensity factors corresponding to the best-estimate initial and final measured depths.) The tolerance bars on the points illustrate the uncertainty in the average crack growthbetween the two UT inspections based on the statistical tolerancefor the crack extension discussed above (4.2 mm). Note that thestatistical lower bound crack growth rate for Crack 1 correspondsto no growth because the 4.2 millimeters tolerance is greaterthan the best-estimate extension for this crack.

VI.C Comparison with Other Published Curves for Alloy 182

Figure 17 compares the MRP-115 curve for Alloy 182/132with the following four deterministic curves: The two-part curve developed by Ringhals [28]. The plateau curve published by EDF based on laboratory test

data [19,30].

The MRP-21 curve that has previously been applied in theU.S. based on a limited set of laboratory CGR data [6,7].

The MRP-55 curve for thick-wall Alloy 600 material [2,3].Like the MRP-115 curve, the MRP-55 equation for Alloy

600 was derived using a multiple regression statistical fit (basedon a heat-by-heat treatment of the data). Unlike the MRP-115 curve, however, the MRP-55 curve assumes a threshold stressintensity factor value of 9 MPam and also uses Scotts value for the exponent (1.16) based on field data for Alloy 600 steam generator tubes [8] rather than letting the exponent be determinedby the statistics associated with the regression fit to the set of available laboratory data.

Examination of the five curves in Figure 17 and theunderlying data leads to the following observations: The MRP-115 curve is based on a worldwide database of

CGR measurements for both Alloy 182/132 and Alloy 82from numerous laboratories. CGRs based on crack extensionaveraged over the specimen width (excluding any segmentswith zero crack extension) were used in the statistical modelthat yielded the MRP-115 curve.

The MRP-115 curve is about 25% lower than the MRP-21 curve for stress intensity factors greater than about20 MPam. At stress intensity factors less than 15 MPam,the MRP-115 curve is higher.

526

The MRP-115 curve is nearly parallel to, and about fourtimes higher than, the MRP-55 curve for stress intensityfactors greater than 20 MPam.

The MRP-115 curve crosses the Ringhals curve at about22 MPam and again at 49 MPam. For stress intensityfactors outside this range, the MRP-115 curve is higher.

Similarly, the MRP-115 curve crosses the EDF curve atabout 9 MPam and again at 27 MPam. For stress intensityfactors outside this range, the MRP-115 curve is higher.

VII. Example Application

Now that the crack growth model has been developed, it is helpful to illustrate its application to a typical geometry whereflaws have been found in the field. Subsequent to this example, conclusions are presented regarding a series of examplesinvestigating the effect of the new MRP-115 model with no stressintensity factor threshold, compared to the previous model ofMRP-21 [6,7], which had a threshold.

Before proceeding to the example calculations, the followinggeneral steps constitute a deterministic crack growth evaluation(additional guidance on the overall approach is provided inSection XI of the ASME Code [31]): Calculate the stress field in the region of interest including

the effect of welding residual stresses and normal operatingstresses. Either a conventional strength-of-materialsapproach or, alternatively, finite-element analysis (FEA) can be used to determine the stresses. Use of FEA is normallyrequired if there are weld repairs to the inside and/or outsideweld surfaces (see, e.g., MRP-106 [32]).

Determine the stress intensity factor K that corresponds tothe postulated weld geometry as a function of crack size.References [33], [34], and [35] provide standard Kexpressions from LEFM that are often applied to calculatestress intensity factors from the corresponding stress field.These standard K expressions are based on LEFMsuperposition assumptions, so they do not take any credit forrelaxation of the residual stress field as the crack grows.

Choose an initial flaw size based on the size crack that is detected in the field, the detectability limit for a particulartype of inspection, or another criterion such as the size crackthat results in a CGR of engineering significance. Choose afinal crack size based on criteria such as the size crack thatproduces coolant leakage, the allowable crack size forcontinued service, or the critical crack size for pressureboundary rupture. Typically, an assumption also is made regarding the flaw aspect ratio (length vs. depth) during the growth process.

Calculate the time for crack growth by integrating a deterministic CGR equation such as the MRP-115 equation(Equation 5) for the variable K as a function of crack size. Typically, the normal operating temperature is assumed, and the number of points in the numerical integration is selectedto be large enough so that the result is insensitive to the step size.

VII.A Example Application: PWR Piping Butt Welds

The location chosen for this example is the reactor vesseloutlet nozzle-to-safe-end weld, corresponding to a plant designedby Westinghouse, where the nozzle is low-alloy steel and the

piping is stainless steel. The stainless steel safe end is welded tothe pipe in the field in most applications, and is connected to thenozzle with an Alloy 182 weld, as shown in Figure 18.

In this example, a flaw is postulated in the Alloy 182 weldmaterial, oriented circumferentially, with a range of aspect ratios.The calculations discussed here have considered all theappropriate loadings, including dead weight, thermal expansion,welding residual stress, and pressure. Since PWSCC is a long-term phenomenon controlled by steady-state stresses, seismic loads and thermal transient loads are not included. Standardresidual stress distributions were assumed from expressions developed in the 1980s based on residual stress measurements for test mockups of BWR piping butt welds [36].c Details regardinghow the strength-of-materials approach may be applied todetermine stresses for piping butt welds are provided in MRP-109 [38] and MRP-112 [39].

Results were generated for the example location using the MRP-115 CGR model (Equation 5) to determine the time for anassumed initial part-depth, circumferential flaw to grow in thethrough-wall direction to a depth of 75% of the wall thickness, for a range of crack aspect ratios (Figure 19). The nominal outsidediameter of the pipe is 30 inches, and the wall thickness is2.6 inches. The temperature used for the example was 617F (325C), so no temperature adjustment was required when applying Equation 5.

Fig. 18. Geometry of Weld Region Used for the Crack GrowthIllustration (Reactor Vessel Outlet Nozzle-to-Safe-End Weld)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0 60 120 180 240 300

time (months)

a/t (f

law

dep

th/w

all

thic

knes

s)

T=617 F

AR=3AR=6

AR=10

AR=2

MRP-115 CGR Model

AR = Aspect Ratio= flaw length/flaw depth

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0 60 120 180 240 300

time (months)

a/t

(fla

w d

epth

/wal

l th

ickn

ess)

T=617 F

AR=3AR=6

AR=10

AR=2

MRP-115 CGR Model

AR = Aspect Ratio= flaw length/flaw depth

Fig. 19. Results of Sample Calculations for a Range of Flaw Shapes: Time for Through-Wall Growth for a Part-Depth

Circumferential Flaw at a Reactor Vessel Outlet Nozzle Safe EndWeld (Including Residual Stress)

c Although many BWR piping butt welds are substantially thinner than PWR piping butt welds of corresponding diameter, finite-elementcalculations of welding residual stresses in PWR piping butt welds (in the absence of weld repairs) [37] indicate that the standard residual stressdistributions are generally conservatively high.

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Note that the crack growth is most rapid for the largest aspect ratio, 10, which corresponds to a flaw with length 10 times its depth. The assumed initial flaw depth was about 2% of the wall thickness, or 0.05 inches. The assumption of a shallower flaw would have resulted in a longer growth time, but the goal here was to illustrate how a realistic flaw would grow. As can be seen in the figure, the assumed initial flaw size can have a significant effect on the result.

VII.B Effects of a Stress Intensity Factor Threshold Assumption

The MRP-115 model contains no stress intensity factor threshold for crack growth, because no basis could be found for the existence of one in the MRP database, which includes no data for crack-tip stress intensity factors less than 19.7 MPam. As such, it differs from the previous work reported in MRP-21 [6,7], where a threshold used for Alloy 600 was assumed to apply to the weld metal as well. Calculations were carried out to examine the effect of the stress intensity factor threshold assumption at various PWR piping butt weld locations [9]. The effect of the elimination of the stress intensity factor threshold on crack growth through the thickness of the weld was in each case detrimental, in that the time for a flaw to propagate through the wall was shorter. Once a flaw was through the wall and leaking, the time required for it to reach a critical circumferential length was actually calculated to be somewhat longer with the MRP-115 model, although the difference was sometimes small. This behavior is the result of the MRP-115 CGR curve being higher than the MRP-21 curve for low stress intensity factors but about 25% lower than the MRP-21 curve for stress intensity factors greater than about 20 MPam (see Figure 17), in combination with residual stresses that often become compressive near the center wall region of a thick weldment. These negative stresses tend to produce low stress intensity factors for part-depth flaws that are propagating through the wall. On balance, it may be concluded that the use of a model with no stress intensity factor threshold value is conservative for the PWR piping butt weld application. In addition to this advantage, this approach is more strongly based technically, as discussed in Appendix B to this article.

VIII. Summary and Conclusions

The following are the key conclusions regarding the MRP study of stress corrosion CGRs of Alloy 82/182/132 nickel-based weld metals under PWR primary water conditions: An international expert panel was formed and collected

detailed laboratory test data for the relevant set of worldwide laboratory CGR tests using pre-cracked fracture mechanics specimens.

The expert panel developed screening criteria to qualify data for use in the development of a deterministic CGR model for Alloy 82, 182, and 132 weld metals. The screening criteria were based upon the criteria previously applied to Alloy 600 wrought material in MRP-55 [2,3], but were necessarily extended to cover the special test considerations associated with the weld metal materials.

Based on a literature review and the laboratory experience of the expert panel members, a methodology was developed for considering the potentially non-conservative effect of incomplete engagement to intergranular SCC across the

specimen width and over test duration. Engagement fractions were estimated for all the specimens in the screened database, and, in the case of incomplete engagement, the reported CGRs were adjusted by dividing by the respective engagement fractions. This approach is appropriate regardless of whether the incomplete engagement is caused by isolated islands of more crack-resistant material or is a testing artifact due to the difficulty of the crack transitioning from the transgranular fatigue pre-crack to the intergranular stress corrosion crackor a combination of the two.

The expert panel concluded that there are currently insufficient data available to include a stress intensity factor threshold in the deterministic CGR model for the nickel-based weld metals. Analyses of weld metal cracking that involve the existence of pre-existing defects (either real or postulated) could be strongly influenced by assuming an arbitrary stress intensity factor threshold value.

A linearized, multiple regression statistical model was fitted to the screened database including an Arrhenius temperature correction, an alloy factor (Alloy 182/132 or Alloy 82), a crack orientation factor (parallel or perpendicular to the weld dendrites), a crack-tip stress intensity factor exponent, and a weld factor that accounts for the randomness associated with the heat of weld wire/stick material and welding process. Insufficient data were available to include dissolved hydrogen concentration (or electrochemical potential), cold working, post-weld heat treatment stress relief, or loading type (constant or periodic unloading) in the model.

For the purpose of producing a single deterministic CGR model, the 75th percentile weld factor was absorbed into the statistical model. The MRP recommends that Equation 5 be applied for the disposition of PWSCC flaws detected in Alloy 182/132 and Alloy 82 in PWR primary circuits (analogous to [4,40]) and used in safety case calculations that assume hypothetical PWSCC flaws [38,39]. Furthermore, data such as those in Table IV and Figure 11 may be used to determine statistical CGR distributions for use in probabilistic fracture mechanics models of the growth of PWSCC flaws in the weld metal materials.

Detailed comparisons with the available worldwide laboratory CGR data that were not included in the final screened database used to produce the MRP-115 deterministic model were performed. These comparisons verified the robustness of the MRP-115 multiple linear regression model, given the manner in which the data screening process was implemented, and verified the absence of any hidden effects in the overall set of CGR data collected.

Evaluation of the only known set of repeat PWSCC crack sizing data for nickel-based weld metals in an operating PWR plant (2 cracks in Alloy 182 reactor vessel outlet nozzle-to-safe-end weld at the Swedish plant Ringhals Unit 3) produced best-estimate CGRs bounded by the MRP-115 curve for Alloy 182/132, as shown in Figure 16.

In other countries, different approaches have been applied to develop CGR disposition curves for the nickel-based weld metals, resulting, as would be expected, in CGR curves somewhat different than the MRP-115 model (see Figure 17).

The MRP-115 equation (Equation 5) was applied to calculate the time for flaws in piping butt weldments to grow to larger sizes. As expected, the assumption of no stress intensity

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factor threshold has a significant effect for relatively small part-depth flaws.

Acknowledgements

This paper is a summary of work sponsored by EPRI on behalf of the Alloy 600 Issue Task Group d of the Materials Reliability Program [9]. Valuable input and review comments were received from many sources, but the authors wish to express their special gratitude to P. Andresen, S. Attanasio, W. Bamford, W. Cullen, J. Daret, P. Efsing, S. Fyfitch, R. Jacko, C. Jansson, A. Jenssen, A. McIlree, W. Mills, R. Pathania, P. Scott, W. Shack, T. Yonezawa, and K. Yoon.

Appendix A: Reasons for Data Exclusion

The original set of worldwide laboratory CGR data collected by the MRP comprised 261 individual data points. The technical issues that were addressed by the screening process are listed in Table III. In practice, 184 data points were excluded from the statistical evaluation for the following objective reasons: reported CGR based only on the maximum crack increment

along the crack front because the MRP data reduction was based on the average crack extension (95 points),

no measurable crack growth (24 points), less than 0.5 mm of crack extension averaged along the crack

front (23 points), hold time less than 1 hour for periodic unloading tests

(18 points), complex loading changes during the test (16 points), hydrogen concentration outside standard plant range

(12 points with 150 cc/kg), loading exceeding the nominal linear elastic fracture

mechanics (LEFM) limit (9 points), engagement to intergranular (i.e., stress corrosion) cracking

along less than 50% of the crack front (4 points), flutter loading (2 points), and temperature change during the test (1 point). For 20 of the excluded data points, two of the above reasons applied. Note, however, that the data excluded because only maximum CGRs were available were not evaluated for compliance with the screening criteria regarding minimum average crack extension and minimum engagement to intergranular cracking and might have been affected by these further considerations.

Appendix B: Assumption of Zero Stress Intensity Factor Threshold for the Alloy 82/182/132 Weld Metals

For the CGR equation developed for the weld metal materials, it was concluded that insufficient data exist to justify a stress intensity factor threshold other than zero. This appendix briefly describes the factors taken into consideration by the expert panel in this recommendation. Note that in MRP-55 [2,3] a threshold KISCC of 9 MPamwas assumed as a curve-fitting parameter for Alloy 600 based on the arrest of cracks in Alloy 600 steam generator tubes once stress intensity factor values decreased to about 9 MPam [8], and MRP-55 cautions that use of its equation for Alloy 600 at low

d Chairman: C. Harrington; EPRI project manager: C. King.

crack-tip stress intensity factors (< approximately 15 MPam (14 ksiin)) involves assumptions not currently substantiated by actual CGR data for CRDM nozzle materials. First, it must be recognized that the threshold stress intensity factor, KISCC, is a concept that is difficult to implement as a practical engineering tool for PWSCC of nickel-based alloys. Because SCC is a time-dependent process, KISCC is not an absolute material property, but depends on test procedure and duration. KISCC is sometimes designated at an arbitrary, slow, selected crack growth rate, but slow crack growth over long periods can, however, be significant. Furthermore, in nearly all practical cases, SCC initiates in circumstances where linear elastic fracture mechanics (LEFM) cannot be applied, and consequently the initiation of short cracks involves physical processes different from those responsible for the arrest of longer cracks. No data in the MRP database following the screening process are available for Alloys 182/132 at stress intensity factors less than about 20 MPam and for Alloy 82 at stress intensity factors less than about 27 MPam. In addition, for the weld metals, no comparable field data to those for Alloy 600 steam generator tubes are available that might allow such a threshold to be reasonably estimated. Moreover, analyses of weld metal cracking that involve the existence of pre-existing defects (either real or postulated) could be strongly influenced by assuming an arbitrary KISCC. Finally, trial fits of the laboratory data with and without an imposed threshold did not support the assumption of a particular KISCC value. Based on these considerations, the MRP-115 CGR equation for the nickel-based weld metals does not take credit for a stress intensity factor threshold greater than zero. This conservative approach has been adopted until data become available specific to the weld metals that justify assuming a KISCC threshold.

References

1. W. Bamford and J. Hall, A Review of Alloy 600 Cracking in Operating Nuclear Plants: Historical Experience and Future Trends, Proceedings of 11th International Conference on Environmental Degradation of Materials in Nuclear Power SystemsWater Reactors, ANS, 2003, pp. 10711081.

2. Materials Reliability Program (MRP) Crack Growth Rates for Evaluating Primary Water Stress Corrosion Cracking (PWSCC) of Thick-Wall Alloy 600 Materials (MRP-55NP) Revision 1, EPRI, Palo Alto, CA: 2002. 1006695-NP. NRCADAMS Accession No. ML023010510.

3. G. A. White, J. Hickling, and L. K. Mathews, Crack Growth Rates for Evaluating PWSCC of Thick-Wall Alloy 600 Material, Proceedings of 11th International Conference on Environmental Degradation of Materials in Nuclear Power SystemsWater Reactors, ANS, 2003, pp. 166179.

4. Paragraph IWB-3660, Evaluation Procedure and Acceptance Criteria for PWR Reactor Vessel Upper Head Penetration Nozzles, ASME Boiler & Pressure Vessel Code, Section XI, ASME International, July 2004 Edition.

5. R. S. Pathania, A. R. McIlree, and J. Hickling, Overview of Primary Water Cracking of Alloys 182/82 in PWRs, Fontevraud V International Symposium on Contribution of Materials Investigation to the Resolution of Problems Encountered in Pressurized Water Reactors, SFEN, 2002, pp. 1327.

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6. Crack Growth of Alloy 182 Weld Metal in PWR Environments (PWRMRP-21), EPRI, Palo Alto, CA: 2000. 1000037.

7. W. H. Bamford, J. Foster, K. R. Hsu, L. Tunon-Sanjur, and A. McIlree, Alloy 182 Weld Crack Growth, and its Impact on Service-Induced Cracking in Operating PWR Plant Piping, Proceedings of 10th International Conference on Environmental Degradation of Materials in Nuclear Power SystemsWater Reactors, NACE International, 2002.

8. P. M. Scott, An Analysis of Primary Water Stress Corrosion Cracking in PWR Steam Generators, Presented at NEA/CSNI Specialist Meeting on Operating Experience with Steam Generators, Brussels, Belgium, September 1620, 1991.

9. Materials Reliability Program Crack Growth Rates for Evaluating Primary Water Stress Corrosion Cracking (PWSCC) of Alloy 82, 182, and 132 Welds (MRP-115),EPRI, Palo Alto, CA: 2004. 1006696. Nonproprietary version (MRP-115NP, 1006696-NP) planned to be made public through submittal to the US NRC.

10. J. C. Lippold and A. J. Ramirez, Elevated Temperature Grain Boundary Embrittlement and Ductility-Dip Cracking of Nickel-Base Weld Metals, USNRC-ANL Conference on Vessel Penetration Inspection, Cracking, and Repairs,September 29October 2, 2003, Gaithersburg, Maryland.

11. P. M. Scott, et al., Examination of Stress Corrosion Cracks in Alloy 182 Weld Metal after Exposure to PWR Primary Water, Proceedings of 12th International Conference on Environmental Degradation of Materials in Nuclear Power SystemsWater Reactors, The Minerals, Metals & Materials Society, 2005.

12. D. Buisine, et al., PWSCC Resistance of Nickel B

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