ML081570607

New England Coalition, Inc., Contentions 4 Prefiled Exhibits NEC-UW_14 - NEC-UW_22, Volume 2

ML081570607
Person / Time
Site:	Vermont Yankee File:NorthStar Vermont Yankee icon.png
Issue date:	04/28/2008
From:	New England Coalition
To:	Karlin A, Wendy Reed, Richard Wardwell Atomic Safety and Licensing Board Panel
SECY RAS
References
50-271-LR, ASLBP 06-849-03-LR, RAS M-39
Download: ML081570607 (378)

Text

Unredacted UNITED STATES NUCLEAR REGULATORY COMMISSION ATOMIC SAFETY AND LICENSING BOARD Before Administrative Judges:

Alex S. Karlin, Chairman Dr. Richard E. Wardwell Dr. William H. Reed In the Matter of )

)

ENTERGY NUCLEAR VERMONT YANKEE, LLC ) Docket No. 50-271-LR and ENTERGY NUCLEAR OPERATIONS, INC. ) ASLBP No. 06-849-03-LR

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(Vermont Yankee Nuclear Power Station) )

NEW ENGLAND COALITION, INC.

CONTENTIONS 4 PREFILED EXHIBITS NEC-UW 14- NEC-UW 22 April 28, 2008 Volume 2

NEC-UW_14

• n~han"6Re-Re; - LRU-date to CHECWOR-KS3 ... ...........

__*....... ___ ---- _1 From: Beth Sienel To: Jonathan Rowley Date: 02/20/2008 9:03:30 AM

Subject:

Re: Update to CHECWORKS Jonathan, I talked to the FAC program owner (Jim Fitzpatrick) and he said the update is in progress. More details:

The Fleet has upgraded to the new version of Checworks.and VY put EPU conditions into the program.

They are now in the process of verifying.

Hope this helps, Beth

>'>> Jonathan Rowley 2/19/2008 4. 16 PM >>>

Beth I (and OGC) need to find out if VY has updated the CHECWORKS computer program they used to predict and track pipe thinning to account for powerjuprate conditions. VY stated -during the EPU process that the FAC Program (using CHECWORKS) would be~updated to account for uprated power conditions. There has been one outage since the EPU was granted duringwhich the updating was to have initiated, that is my understanding.

Could you contact the FlowrAccelerated Corrosion Program owner and verify if they have started updating the program?

CC: Raymond Powell; Ricardo Fernandes

I c~ep\G}OOO6MPPage, 1 Mail Envelope Properties (47BC3325.EC4: 5 : 55534) U

Subject:

Re: Update to CHECWORKS Creation Date 02/20/2008 9:03:17 AM From: Beth Sienel Created By: BEK@nrc.gov I Recipients nrc.gov J TWGWPO03.HQGWDOOI JGR (Jonathan Rowley) nrc.gov kpl_po.KP_DO RAF 1 CC (Ricardo Fernandes) I RJP CC (Raymond Powell)

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I NEC-UW_15 PENNSTATE I Department of Mechanical and Nuclear Engineering (814) 865-2519 College of Engineering Fax: (814) 8634848 The Pennsylvania State University 137 Reber Building University Park. PA 16802-1412 Dr. Brian W. Sheron Associate Director for Project Licensing and Technical Analysis U.S. Nuclear Regulatory Commission MS 05E7 11555 Rockville Pike Rockville, MD 20852-2738

Dear Dr. Sharon:

Enclosed are the results of a project given to my Penn State Graduate Students on finding pipe failure data over a range of pipe sizes and conditions. We specifically looked for stainless steel data as well as carbon steel pipe data. Since the data is from several sources other than nuclear the pipe wall thickness may not always be comparable to reactor pipe wall thicknesses. In some of the reports the students did separate the failure and leakage data by mechanism such that we could then screen the data.

I I had the students normalize the data in such a fashion that we could then compare to the break frequency spectrum curves generated by the NRC experts group. I did talk to Rob Tenoning on the best way of normalizing our data such that we would be consistent with the break frequency plots. The key findings from the students work is that the data, when plotted in the same manner as the break frequency spectrum plots from the NRC experts work, shows a much flatter behavior at the larger pipe sizes indicating a more similar probability level for failure as compared to a more significant decrease in the failure probability as given by the NRC break frequency spectrum.

II am complying all the independent sets of data in a spread sheet and will attempt a further screening. Once complete, I will send you a copy of the data. I wanted you to have these report now with all the data so you could make an independent assessment.

Please let me know if you need anything else.

I Very truly yours, L.E. Hochreiter Professor of Nuclear and Mechanical Engineering I

College of Engineering An Equat Opportunity University

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I NucE 597D - Project 1 I

DATA COLLECTION OF PIPE FAILURES OCCURING IN STAINLESS STEEL AND CARBON STEEL PIPING I I

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Pennsylvania State University Dr. L.E. Hochreiter I April 2005 I

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Executive Summar-y Currently the Nuclear Regulatory Commission (NRC) is contemplating changing the acceptance criteria for Emergency Core Cooling Systems (ECCS) for light-water nuclear power reactors contained in NRC Regulation 10 CFR 50.46. This regulation sets specific numerical acceptance criteria for peak cladding temperature, clad oxidation, total hydrogen generation, and core cooling under loss-of-coolant accident (LOCA) situations. Furthermore, the regulation requires that a spectrum of break sizes and locations be analyzed to determine the most severe case and to ensure the plant design can meet the acceptance criteria under such conditions.

Currently the regulation states that breaks of pipes in the reactor coolant pressure boundary up to, and including, a break equivalent in size to the double-ended rupture of the largest pipe in the reactor coolant system must be considered. While this restricts the design, it maintains a large safety margin ensuring the plant-is covered under all LOCA situations. However, an impetus for change has resulted from materials research, analysis, and experience that indicate that the catastrophic rupture of a limiting size pipe at a nuclear power plant is a very low probability event.

If approved, the proposed change would divide the break spectrum into two categories based upon the likelihood of a break. Breaks of higher likelihood, breaks smaller than 10 inches, would need to meet the current requirements set forth in 10 CFR 50.46. Breaks of a lower likelihood, those larger than 10 inches, would only need to meet the requirements of maintaining a coolable geometry and having the capability for long term cooling.

The purpose of this project was to collect data on instances of pipe failures including cracks, leaks, and ruptures. For each instance of failure the plant type, pipe diameter, type of pipe, failure mechanism, and type of failure was recorded. The data was then collapsed based on plant type (PWR or BWR), type of pipe (carbon or stainless steel), pipe size, and failure mechanism.

Then, normalized failure frequencies were calculated as a function of both pipe size and failure mechanism per reactor year. Plots of the frequency distributions were generated on a semi-log scale, and the frequency distributions as a function of pipe size were compared to the NRC predicted failure frequencies.

For this project our group collected two, independent sets of data. The first set was provided by the OECD Pipe Failure Data Exchange Project (OPDE), with a total of 2891 data points. The second set consists of 67 data points collected by our group from various sources. The two sets of data were not combined due to the lack of information accompanying the data presented in the OPDE database, such as plant name or exact failure size. This made it impossible to identify overlapping coverage and combine the information. Rather, within this report we have analyzed each data set individually in order to make an overall comparison of the trends observed for each data set and the NRC predictions.

The results from both the OPDE and the independent sets of data detailed in this report do not support the NRC's assertion that larger sized pipes do not break frequently enough to be used as design criteria. The overall trends of both sets of data show that the frequency of failures does not decrease as sharply with increasing pipe size as the NRC predicts.

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Table of Contents 1.0 Detailed Introduction to the Problem ................................................................................ 6 2.0 Data Collected ............................................................................................................. 8 2.1 OECD Pipe FailureData Exchange Project...................................................... 8 2.2 Independently Collected Data .............................................................................. 9 3.0 Collapsing and Analyzing the Collected Data ................................................................... 12 4.0 Results and comparisons .................................................................................................... 15 4.1 FailureFrequency as afunction of Pipe Size ...................................................... 15 4.2 FailureFrequencyas afunction of FailureMechanism..................................... 25 5.0 Conclusions ............................................................................................................................ 31 6.0 References ............................................................................................................................... 33 Appendix A - OPDE-Light Database Appendix B - Independent Database Appendix C - Collapsed OPDE Data Appendix D - Copies of References 3

List of Fiaures Figure 4.1-1. Normalized pipe failure frequencies as a function of pipe group size for both carbon and stainless steel pipe failures in both BWR and PWR plants Figure 4.1-2 Normalized rupture frequencies as a function of pipe group size for both carbon and stainless steel pipe failures in both BWR and PWR plants Figure 4.1-3. Normalized Failure Frequency Distribution for PWRs Figure 4.1-4. Normalized Failure Frequency Distribution for BWRs Figure 4.1-5. Normalized pipe failure frequencies as a function of pipe size f6r PWRs Figure 4.1-6. Normalized pipe failure frequencies as a function of pipe size for BWRs Figure 4.1-7. Normalized pipe failure frequencies as a function of pipe size for PWRs using the Modified Analysis Method.

Figure 4.1-8. Normalized pipe failure frequencies as a function of pipe size for PWRs using the Modified Analysis Method.

Figure 4.2-1. Normalized pipe failure frequency as a function of Pipe Group Size for PWRs Figure 4.2-2. Normalized pipe failure frequency as a function of Pipe Group Size for BWRs Figure 4.3-1. PWR Failure Frequency for Carbon and Stainless Steel Pipes as a Function of Failure Mechanism Figure 4.3-2. BWR Failure Frequency for Carbon and Stainless Steel Pipes as a Function of Failure Mechanism Figure 4.3-3. PWR and BWR Failure Frequency for Carbon and Stainless Steel Pipes as a Function of Failure Mechanism Figure 4.3-4. Pipe Failure by Corrosion as a Function of Pipe Size (PWR & BWR)

Figure 4.3-5. Pipe Failure by Fatigue as a Function of Pipe Size (PWR & BWR)

Figure 4.3-6. Pipe Failure by Mechanical Failures as a Function of Pipe Size (PWR & BWR)

Figure 4.3-7. Pipe Failure by Stress Corrosion Cracking as a Function of Pipe Size (PWR &

BWR) 4

List of Tables Table 1-1. NRC Total Preliminary BWR and PWR Frequencies Table 2-1. Excerpt from "OPDE-Light" Database Table 2-2. Description of Plant Systems and Type of Piping Table 2-3. Definition of OPDE Pipe Size Groups Table 2-4. OPDE Pipe Failure Definitions Table 3-1. Definition of Pipe Size Groups Table 3-2. Definition of NRC LOCA Groups Table 4.1-1. OPDE Calculated, and NRC Predicted, Normalized Failure Frequencies (1/cal-yrs).

Table 4.1-2. Normalized Rupture Frequencies Table 4.1-3. Summary of PWR Pipe Failures from the OPDE Database as of 2-24-05 Table 4.1-4. Summary of BWR Pipe Failures from OPDE Database as of 2-24-05 Table 4.1-6. Summary of PWR Pipe Failures from OPDE Database as of 2-24-05, using the Modified Analysis Method.

Table 4.1-7. Summary of BWR Pipe Failures from OPDE Database as of 2-24-05, using the Modified Analysis Method.

Table 4.2-1. OPDE Calculated, NRC Predicted, and Independent Database Calculated, Normalized Failure Frequencies (1/cal-yrs)

Table 4.3-1. Failure Frequencies of Pipes for each Failure Mechanism 5

1.0 Detailed Introduction of Problem In order to ensure the safety of nuclear plants the cooling performance of the Emergency Core Cooling System (ECCS) must be calculated in accordance with an acceptable evaluation model, and must be calculated for a number of postulated loss-of-coolant accidents (LOCA) resulting from pipe breaks of different sizes, locations, and other properties. This is done to provide sufficient assurance that a plant can handle even the most severe postulated LOCA. LOCA's are hypothetical accidents that would result from the loss of reactor coolant, at a rate in excess of the capability of the reactor coolant makeup-system. Currently, the evaluation criteria for these types of accidents state that pipe breaks in the reactor coolant pressureboundary up to and including a break equivalent in size to the double-ended rupture of the largest pipe in the reactor coolant system must be considered. In the case of such an event the NRC has set forth the following criteria that must be met for a design to be considered acceptable [37]:

a. Peak cladding temperature must not exceed 22000 F.

b. Maximum cladding oxidation must not exceed 0.17 times the total cladding thickness before oxidation.

c. Maximum hydrogen generation. The calculated total amount of hydrogen generated from the chemical reaction of the cladding with water or steam shall not exceed 0.01 times the hypothetical amount that would be generated if all of the metal in the cladding cylinders surrounding the fuel, excluding the cladding surrounding the plenum volume, were to react.

d. A coolable geometry of the core must be maintained.

e. After any calculated successful initial operation of the ECCS, the calculated core temperature shall be maintained at an acceptably low value and decay, heat shall be removed for the extended period of time required by the long-lived radioactivity remaining in the core.

While requiring that all plants be analyzed in the case of a double-ended guillotine break of the largest pipe restricts the design, it does maintain a large safety margin ensuring the plant is covered in all pipe break situations. However, an impetus for change has resulted from materials research, analysis, and experience which indicate that the catastrophic rupture of a large pipe at a nuclear power plant is a very low probability event. The hypothesis that is currently being set forth is that small pipes break more frequently than large pipes. The criteria would change so that the NRC would refocus their analysis efforts because they want to make sure that the appropriate amount of time and money are being invested in the areas of most concern.

Furthermore, risk analyses indicate that large break LOCA's are not significant contributors to plant risk. According to a presentation given by Dr. Brian Sheron of the NRC at Penn State in the Fall 2004, "using the double ended break of the largest pipe in the reactor coolant system as the design basis for the plant results in ECCS equipment requirements which are inconsistent with risk insights and places an unwarranted emphasis and resource expenditure on low risk 6

~I contributors. This also places constraints on operations which are unnecessary from a public health and safety perspective." Therefore, the proposed rule change would use the pipe size with the largest break frequency as the design basis for pipe rupture and accident analysis of the plant.

A pipe size with a 10 inch diameter is currently being suggested. [37]

I The proposed change would divide the break spectrum into two categories based upon the likelihood of a break. Breaks of higher likelihood, or those smaller than 10 inches, would need to meet the current requirements set forth in 10 CFR 50.46. These include criteria (a) through (e) above. On the other hand, breaks of a lower likelihood, or those larger than 10 inches up to and including a double-ended guillotine break of the largest pipe in the reactor coolant system, would I

only need to meet the requirements of maintaining a coolable geometry and having the capability for long term cooling. Thus, criteria (a), (b), and (c) would be eliminated for these cases. [37]

The purpose of this project was to collect data on instances of pipe breaks, leaks, and cracking.

These failures included pipe failures from broken pipes either by splits, ruptures, or guillotines, and cracks in pipes, either circumferential or length wise. For each instance found the plant type, pipe diameter, type of pipe, failure mechanism, and type of failure was recorded. Only stainless steel and carbon steel pipes were considered. Then, normalized failure frequency distributions were developed and compared to NRC predictions.

The predicted NRC failure frequencies were taken from Table 3 on page 14 of 10 CFR 50.46, LOCA Frequency Development [38]. This table is replicated below. m Table 1-1. NRC Total Preliminary BWR and PWR Frequencies.

Current Day Estimates (per cal. yr)

U Plant Effective Type Break Size (inches) 1/2 5_%

3.OE-05 Median 2.2E-04 Mean 4.7E-04

• 95%

1.7E-03 I

1 7/8 3 1/4 7

2.2E-06 2.7E-07 6.6E-08 4.3E-05 5.7E-06 1.4E-06 1.3E-04 2.4E-05 6.OE-06 5.0E-04 9.4E-05 2.3E-05 I

18 1.5E-08 1.IE-07 2.2E-06 6.3 E-'06 41 1/2 3.5E-1I 7.3E-04 8.5E-10 3.7E-03 2.3E-06 6.3E-03 8.6E 2.OE-02 I 1 7/8 .6.9E-06 9.9E-05 2.3E-04 8.5E-04 PWR 3 1/4 7

18 1.6E-07 1.IE-08 5.7E- 10 4.9E-06 6E-0-07 7.5E-09 1.6E-05 2.3E-06 3.9E-08 6.2E-05 8.8E-06 I.5E-07 I

41 4.2E-1.1 I1.4E-09 2.3E-08 7.OE-08 I

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I 2.0 Data Collected I For this project our group collected two, independent sets of data. The first set was provided by the OECD Pipe Failure Data Exchange Project (OPDE), with a total of 2891 data points. The second set consists of 67 data points collected by our group from various sources listed as references in this report. The two sets of data were not combined due to the lack of information accompanying the data presented in the OPDE database, such as plant name and exact failure size, which made identifying overlapping coverage impossible. Rather, within this report each data set was individually analyzed in order to make an overall comparison of the trends observed for each data set and the NRC predictions.

OECD Pipe FailureData Exchange Project [3]

OECD Pipe Failure Data Exchange Project (OPDE) was established in 2002 as an I international forum for the exchange of pipe failure information. It is a 3-year project with participants from twelve countries, including Belgium, Canada, Czech Republic, Finland, France, Germany, Japan, Republic of Korea, Spain, Sweden, Switzerland and 3 the United States. "The objective of OPDE is to establish a well structured, comprehensive database on pipe failure events and to make the database available to project member organizations that provide data." [3] The OPDE database evolved from what existed in the "SLAP database" at the end of 1998 [2].

OPDE covers piping in primary-side and secondary-side process systems, standby safety systems, auxiliary systems, containment systems, support systems and fire protection systems. Furthermore, ASME Code Class 1 through 3 and non-Code piping has been considered. At the end of 2003, the OPDE database included approximately 4,400 records on pipe failure. The database also includes an additional 450 records on water hammer events where the structural integrity of piping was challenged but did not fail.

Access to the actual OPDE database is restricted to organizations providing input data.

However, a."OPDE-Light" version of the database will be made available later this year to non-member organizations contracted by a project member to perform work or which pipe failure data is needed. This version will not include proprietary data, such as the exact pipe diameter, where failure occurred, and preclude any plant identities or dates.

Our group was fortunate enough to get a copy of this "light" version of the database for BWR and PWR pipe failures reported as of February 24, 2005. A total of 2891 failures (1536 for PWR plants and 1355 for BWR plants) were provided in this database, and considered for this project.

I The database listed the plant type, reactor system, apparent cause of failure, pipe size group, number of total failures for each cause and pipe size group, and then a break down of the type of failure within the category. An excerpt from the OPDE-Light database has been provided for clarification in Table 2-1 on the following page. The database, in its entirety, has been included in Appendix A of this report.

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However, there are a few problems with this database related to the purpose of this project. First, since the database did not provide the type of pipe (carbon or stainless) for each failure, a reasonable prediction of what type of pipe was involved in the failureI based on the plant system, which was given, was made. The type of pipe assumed for each system is also given in the following page in Table 2-2.3 Additionally, as previously mentioned, no explicit pipe diameters were given for each failure due to the proprietary nature of this information. Rather, the failures were collected into group sizes before it was sent out. A total of six group sizes were utilizedI by OPDE. The range of pipe diameters that comprise each group is given in Table 2-3.

The main problem with these groupings, and the database in general, is that pipes larger than 10 inches in diameter are all grouped together and there is no way of determiningI how much larger than 10 inches they actually were. Finally, for the purpose of this analysis any crack, leak, or issue (i.e. wall thinning) with the pipe was considered to be a failure. However, the OPDE database lists the information by type of failure. The3 definitions of each faiure type have been included in Table 2-4.

Independently Collected Data [5-36]

For the purpose of this project -our group collected separate information on instances of piping failures and their causes. The information was collected primarily from NuclearI Regulatory Commission (NRC) bulletins, informnation notices, event reports, and generic letters. Our group was able to compile a total of 67 instances of piping failures. This database is provided in Appendix B. While our database is much smaller than the one compiled by the OECD Pipe Failure Exchange Project, it provides an independent check of the trends observed by that database.3 A list of references is provided at the end of this report, and some of the actual references, printed from the NRC website, have been included in Appendix D.3 9I

Table 2-3. Definition of OPDE Pipe Size Grou ps.

Pipe Size Corresponding Corresponding Pipe Diameters Pipe Diameters Group (mm) (inches) 1 DN < 15 DN < 0.6 2 15 < DN < 25 0.6 < DN < 1.0 3 25<DN<50 1.0 < DN < 2.0 4 50 <DN< 100 2.0 < DN <4.0 5 100 < DN < 250 4.0 < DN < 10.0 6 DN > 250 DN > 10.0 Table 2.4. OPDE Pipe Failure Definitions.

Type Description Crack - Part Part through-wall crack (_ 10% of wall thickness)

Crack - Full Through-wall but no active leakage; leakage may be detected given a plant mode change involving cooldown and depressurization.

Wall Thinning Internal pipe wall thinning due to flow accelerated corrosion - FAC Small Leak Leak rate within Technical Specification limits Pinhole Leak and the from Differs "small leak" only in terms of the geometry of the throughwall defect underlying degradation Or damage mechanism Large Leak Leak rate in excess of Technical Specification limits but within the makeup capability of safety injection systems Severance Full circumferential crack- caused by external impact/force, including high-cycle mechanical fatigue - limited to small-diameter piping, typically Large flow rate and major, sudden loss of structural integrity. Invariably caused Rupture by influences of a degradation mechanism (e.g., FAC) in combination with a severe overload condition (e.g., water hammer)

4.3 Pipe Failuresas a function ofFailureMechanism I

This section of the report summarizes the frequency of failure mechanisms for carbon and I

stainless steel pipes. The information presented in figures 4.3-1 through 4.3-3 represents the normalized failure frequencies for each failure mechanism. This data is also presented in tabular form in table 4.3-1. The data was collapsed by pipe sizes and broken apart by steel type and I plant type. The data was normalized for each type of steel based on the number of reactor years and the total amount of failures (carbon +stainless) for each plant.

I Table 4.3-1. Failure Fre uencies of Pipes for each Failure Mechanism.

Plant Failure Mechanism Carbon Steel Stainless Steel Total Failure I

Type Failure Frequency Failure Frequency Frequency PWR Corrosion 2.04E-05 5.38E-06 2.57E-05 I

PWR FAC 2.29E-05 2.32E-05 4.61 E-05 PWR PWR MIC Erosion 8,26E-06 1.84E-05 1,92E-07 2.30E-06 8.45E-06 2.07E-05 I

PWR Fatigue 1.77E-05 9.62E-05 1.14E-04 PWR PWR Human Factors Mechanical Failures 6.91 E-06 4.23E-06 2.42E-05 7.11E-06 3.11E-05 1.13E-05 I

PWR SCC 9.60E-07 3.25E-05 3.34E-05 PWR PWR Water Hammer Misc 0.OOE+00 I.15E-06 3.84E-07 2.69E-06 3.84E-07 3.84E-06 U

BWR BWR Corrosion FAC 6.31E-06 1.26E-05 6.97E-06 1.37E-05 1.33E-05 2,63E-05 I

BWR MIC 1.31E-06 2.18E-07 1.52E-06 BWR BWR Erosion Fatigue 8.71E-06 1.55E-05 1.96E-06 4.90E-05 1.07E-05 6.44E-05 I

BWR Human Factors 5.22E-06 1.85E-05 2.37E-05 BWR BWR Mechanical Failures SCC 3.92E-06 4.14E-06 5.44E-06 1.36E-04 9.36E-06 1.40E-04 I

BWR Water Hammer 4.35E-07 2.18E-07 6.53E-07 BWR Misc 8.71E-07 4.14E-06 5.01E-06 I I

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1.2E-04 1.0E-04 F 0 a Carbon Steel Stainless Steel I

EoCarbon &Stainless Steel E 80-0 8.OE-05

0. 46.OE-05

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LL. 4.0E-05 2.05-05 IA Corrosion FAC 771J7 MIC Erosion Fatigue Human Factors Mechanical Failures SCC Water Hammer Misc Failure Mechanism Figure 4.3-1. PWR Failure Frequency for Carbon and Stainless Steel Pipes as a Function of Failure Mechanism I

I Corrosion FAC MIC Erosion Fatigue Human Mechanical .SCC Water Misc Factors Factors Hammer Failure Mechanism Figure 4.3-2. BWR Failure Frequency for Carbon and Stainless Steel Pipes as a Function of Failure Mechanism 26

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1.000E-04 9.000E-05 r Carbon Steel a Stainless Steel 0 Carbon and Stainless Steel I I

6.000E-05 I

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7.000E 4) 5.OOOE-05I -7 5.000E.05, S3.000-E05 I

2.000E-05 1.OOOE-05 I

iA Corrosion FAC "In MIC Erosion Fatgue Human Mechanical M.

SCC Water Misc I Factors Faiures Hammer Failure Mechanism Figure 4.3-3. PWR and BWR Failure Frequency for Carbon and Stainless Steel Pipes as a Function of Failure Mechanism U

From these plots it was determined that PWR plants are dominated by fatigue failures and BWR I

plants are dominated by stress corrosion cracking failures. However, in general the most frequent failure mechanisms for both plants are corrosion, fatigue, mechanical factors, and stress corrosion cracking.' These four failure mechanisms were analyzed as a function of pipe size in I

figures 4.3-4 through 4.4-7.

For these plots corrosion includes general corrosion, flow accelerated corrosion, and I

microbiological corrosion. Stress corrosion cracking was not included with corrosion because the pipe failure method for stress corrosion cracking is different than the other corrosion types.

Though mechanical failure frequency was not the highest, mechanical failures were chosen I because they appear to be independent of pipe type and plant type. Human factors were ignored because they are a factor of quality assurance as opposed to the other failure mechanisms which are primarily a factor of operation. In regards to human factors it is not known if they have U decreased with reactor operating experience because the dates of failures was not included with the OPDE data.

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3 PWR i sS RAS 1 4 2 1 1~ 1 A 2 PWR ss RAS Fiow-cawabohn 4 2 2 PWR Excessve V~trabon PWR FAC - Flow Accterated CoTrOnf PWR FAC. Flow Accelerted Corroswn PWR ss RAS Fret~ng I I t PWR PWR PWR PWR HF.CONSTANST - 3 6271 I - 1 4 PWR RAS RAS RAS PWR SS HF:Hu1an er1or 3 1 5s RAS HFFWe": Erro 4 2 1 1 PWR sS RAS 2 PWR PWR GSC-kC C PWR PWR &S RAS 2 2 PWVR ISS PWR sS PWR as RAS PwSCC 3 1 2 PWVR as RAS RAS ... eo0~i~~*

S-we, ove'oadrg 3 I.

RAS TGSCC- rrugarxiar SCC 5 I 5-RAS TGSCC. Transgar a~n CC 2 t1 PWR 3 4

Thermal Fobe 3 5 .I 3 PWR *ss RAS Thermal Foebg 4 2 PWR f S 3 4

-l 1

Vbrabon-la1 9 ..e 10 1 1 Ss RAS MWrato-taIue -I i1~ 7 3

V'iralofasgue 6 4 PvNR T S RAS RCPa B/A-SCC PWR ss RCPB BCAr-SCC RCPB C.*T-oJcon 2 1 2 1 1 1 PWVVR SS RCPB RCPB RCPB 1 RCP8

- - m m - ------ - -- -- --

m m - - - --

m - - mm - m m M M HFCONSTANST 1 6 2 3 12 7 2

PVVR SS RCPB PWR Ss RCPB ss PAR Ss PAR I -

SS PAR 5S RCPS PWR SS RCPB HF:Wekk~q Error PWR RCPB 2 HF.W,4,g Error 3 2 PVVR HF W&dN ewDr Ooson emrbnernent PWR xC e~as C PAR SS RCPB PWR RCP88 PWSCC 2 44 26 2 1 4 I I PWR RCPB RCP6 PAR PAVR RCPB RCPB PWR PWSCC 6 7 2 2 Severe overil&a" sever e overloar,51 PMR I TGscc -.rwsTauar SCC PAR SS RCPB TGSCC- Trersremular SCC PWR RCPB TGSCC TrarisgsoartrSCC 4 I 4

PWR A 1 PWR SS I RCPB Thermal league 6 1 RCPB Ther*., Fatgue- cyc"9 3 31 PWR 82 2 I 3 1.

PWR RCPB Vibrabotms4tg 3 4 PWR PWR PWR PWR SS RC.S4NSTR t so- Corrosion 1 I I PAR SG 1eefnaAsecneL Fhgue 2 3

I-f-3 PAR CS SG PWSCC 1 PWR CS I GSCC. ,r.,.r~aruer5CC - 2 1 PAR CS Vs*o*ron-Fabgue 2 2 PVVR CS 4 1 f -

PAR S5 sir BtA-SCC 3 PWR B/A-SCC 6 PAR Cavttben-rso1on 3 PAVR 6SS SIR Ceavtabon-eroion 6 2 PWR SS SIR ~_ Caosrosie I PWR 3 I -

PWR PwR SIR 3 PWR SIR FAC.Fiow AceielId CoI¢ossc 2 1 Free*zin 5 -

PAR r6 SIR_#

PAVR I1 SS SIR 2 4 Ss SIR HF.CONSTINST HF'Httian error 2 HF.REPAIRA"NT lIP. We1, Erorw 1

=3

-- I- - ..iI ss HF-W6fia 7 P',/VR SIR

. .. ss.. SIR

I r ~ I I SS SIR HF.WeAV Error 3 PWR HF.Wek"l, Error 4 PWR HF:W6*ng Error 5 PVYR HF.We" Error 6 SS SIR OveCsres,6ed I 3 FWSOC 2 _

SIR PVWSCc PWVR S SIR PWSCC PWR PwScc 2 3 1o I 1 asve ove&loa&V 2 -3 Severe overlactng 2 6

Severe0 oeca 6 2 I 1 PWR SS SIR TGSCC- Traas9 uarr SCC PWR S SIR TGSCC- Tru* w U&nui 2 I TQSCC. Transgsnur ,SC PWR I SS TGSCC- Transg7sAr SCC Thermafaligue PWR SS SIR 71rie='rail ralue 4 3 2 1 T6h6m*Vtatquo 5 8 4 PWR SIR Lkneponed 6 1 I PWVR UnrepOrted vb-a ton-ata=*u PVWR SIR f [ - - 2 6 SS " 2 o 42 3 9

, Vorabon-fab" 4 3 V~aton4a~ue PWR SS SIR 5 7 1 I - 4 PwR CS STEAM Corrosion 3 1 STEAM PWR :1-CS I STEAM FAC-Fbc Aceerated Corrosio 2 10 FAC. AmwAcceeratd CTosm 3 FAC. FlowAceraWed Commono 4 FC. *,iwAc..*,*,edl Corosion I ,~ 3 9- I -

PTR CS STEAM AO. FlowAccr.. atd Corrosio 6 14 1 PWR CS STEAM HF Human error 6 I1 PWVR CS STEAM HF.Woetn Error HF:Wekhfg Enror-HF.WeldingwTr Overstressed I1- 1 I -~ --I PWR CS STEAM_ Savere oveucaok Svaere oversoa, 2 ~~~~~1~

PMR CS STE-.AM 2

-l 9I I-4 PSVR -- I CS ATFAM V.h~afl-abO~e 1541

- -- ~ -- - - - - - - - - - -

m m m m - m m - m m m -I m m - m PPEPEPYE TOTALNO .

PLANTTYPEI PIPETYPE ISYSI"EMGROUP APPARENT CAUSE URO I.. OF RECORDSI Cxa*-FtAl Crack-Pan I eformallon Lare Lea Leak IPAi-Leak IRupture ISevere" SancILeak.. Wahifwngn f3WRi CS AUAU C~orrlof I 1 CS AUXC Coroiron 2 2 1 2 - -- I~1~~ I EWR FCSW RM CS AUXC oaoron.,

AUXC Eroancaitaton 3 t 1 AUXC Eloson-ca,,o.on 6 1 AUXC Erosion-corrosi0n 3 4 bWR CS AUXC 4 7 EWR CS AUXC Ero.uo,-corroJan 5 9 3 ~1~

15 EWR CS AUXC j -

BWR CS AUXC MIC- )Aacrvoog9cay k)vjced CDrosrcai 2 8WR CS AUXC Severe ova 3 3 OWR -CS I- 2 8WR CS 8WR CS 6I 6WR CIS AUXC Vibraab-otn-a ~- 1 I ~ 2 8 BWR CS AUXC Vitabon-'Fa*ue 3 1 I 4

5 6-6WR Ca-1,vea Syse-n SS cn4ai5 sys C,*TosOn 2 1 6WR -SS C,*riawnw Systen 1

_WR SS 6 2 1 BWR SS BWR SS 0 1 BWR 8S CS 4 1 BWR CS TGSCC,-Transgrarnur SCC 6 1 BWR CS 2 1 BWR CS Fretora* 2 BWR CS EHC HF:CONSTANST 1 BWR CS EHC 1 BWR CS EHC- 4 1 BWR CS EHC del 2 1 3 3 BWR CS EHC -7 I

BWR (S EP$ I BW4R ES EPS 7 21 2 BWR 6S EPS 41 1 1-Ccfonaon B7WR FPS FPSF BWR CS 546 1 BWR CS 1 CS 3' I BWR 1 BWR CS FPS HF.Huran ofr I SWR CS FPS BWR EWR 1 BWR FPS j C-LkrodooCalyh xkcedCorrosion BWVR CS . FPS severe oarioaar 4.

--BWR CS PS 3 17-

ss Fw c Corcnosn 2 2 2

,.

SS 6S 2 6S 2 5s FWC GCrosion 6 FWC Coaroslon-lfahgue '2 1 FWC 3 1 FWC I 8WR 6s FWC JO5~ 2 2 2

'3 1 8WR Ss FWC -oslon FWC_ 4 1 1

BWR 8WR FWC BWR SS FWC FAG.-Fow Accelerated Corosion I SWIR 3 8WR 1 Ess v FWC F AG Fow Accelerated Corroso 4 3 210 8 BWR SS FWC FAC. Flow AcCel~ated Conossoi 6 22 1 1 1 BWR 6S -FwC FAC - Flow Accelecaled COrro,.on

-S FWC Fatigue BWR SS FWC HF.CONSTANST 8WR

,SS FwVC HF.CONSTANST -- I I

BWR HF.CONSTANST Si 1 6 1 BWR FWC HF:H-rw error $ I FWC 2 FWC I F'WC 1 GSCC - overoardai S( 4 I I -

BWR 65 FVWC SevereOvroadr I 1 BWR SS FWC SevereOveoNd9 0WR Ss FWC Severe Overloading 62 1 8WR4 SSJ SbCC.- Slrausrle kwi.Aed Corrosion ri acd& .1~

FWC SsCC-Stka 7 4 3 OWR SS FWC 3 I

FWC '5 FWC ThnraJf.tUqje 6 t BWVR SS FWC T T

FWC T

VibraOtOf-laIue 1 2 I

________________________

B3WR FWVC 2 wxaoratontaigje wVaon,-fabgue 3 Vibra*borabqJe 4 5 1 FWC Vorstort-fas que 6 1 BWR cs IASA 2 6WII CS 2 BWR CS 2 11 cs HFI:Xrnan' eoro - 2 I 1 BWR CS tA-SA 2 Cs- Lk.SA I BwFr CS 2 1 BWF* CS LA-SA evvw -cr WASA Vibabofl-Fabg.m 2 BWF* -CS 3 CS Corrosion, I 1 BsR

-CS, Commono. 3 BW*4 PCs 2

CS 36 1 4 6 PCS FAG- Flow Accelerated Co-o6.on Bw'i pCS FA.. Fr . Sted CO.TO.1o.

4 I 86

- - m - -- - - m - - m - - - - -

- = m m - -"--"- = m i -

DVVfl____jtO I -. t-------r------,----,

BWR CS PCs FAC-FZ7- ccekr*WCma"on -- 2 I I

1 - -

SWR CS PCs PCs PCS WAarxonaIgue I I Vorabon-faugue 2 7 4 3 PCs I 1 BWR ~CS PCS 2 BWR $S RAS BWR SS RAS Corroso$4n 3 RAS Corrosion 4 -I ~ 4 6

RAS C,<ITO$On RAS Ccroson-al wu 3 BWR &S RAS 1 BVVR ~SS~ RAS EOSO oExteral Ct1onde L'nII~edSOC I 2 17 I - ~ 8 ECSCC- ExtenW Citdrlde Whiced SOC ECSCC- Exiemal Chlorlde kxuc4d SCC BWR 1 .as FAC. Flow A.celeraola CoIosion SWR RAS Fatgbe 2 1I~

HF.(

RAS RAS BWVR SS RAS I BWR SS HAS HFHLman error GWR SS RAS H F:Hxgrnm error I BWR SS RAS HF REPAJRAlAJNT 2 2I BWH SS j RAS HF REPAIR/"AINT T BWR Ss RAS HF REPAiR/"ANT HF.We"orr*g r HF:WeWdng error HF.Wei" error i

BWR RAS HF.Weln Error 6 4 1 BWR .SS IDSCC0. Inlrdeninbc 5CC BWR OWTI BVVR RAS I-SCO - h 4 32 9 11j13 BWR RAS 0csc5. igrvarar S-* 2 4 7 BWR Ss RAS Severe ovartoaoe 2 3 BWR &S RAS Severe overiecia 4 1 1 BWR Ss RAS TGSC. Trweag SCC 3 7 6 .I BWR RAS 4 RAS I 2 2 1-IWR SS RAS Tl-e'mall* je 3 1 1 RAS RAS RAS I -- I 3 I 3 BWR as RAS Thermnal Fflgue - yci7 5 1 -

BWR ES RAS Tharnal Faegue Cy.ng, 6 1 8WR 3 1 BWR ES RAS RAS Lrnrepoaled 6 1 HAS vlbralI~On-etu 1 4 1 3 S.S BWH RAS 2 15 1 EWR ES RAS 3 7 SS BWR RAS 4 2 RAS VtLflII0g-fu1e 5 1 BWR ss 71 1 1 BSSR S RCP6 Corr. won I BWR Ss RCPB ECSCC -External C2iond* uced SCC 1I3 2

8WR as E CCC - Exlenma C"kon. t BWR SS Eroslon BWR as RCPB ex*n*rl dwnmage 3 - I RCPB I 1 ST RCPB BWR ST RCPB BWR SS RCPB HF:Fai*cabon EIoT I ~

BWR wS RCPB HF:FaLbrlcalon ErTro 6 1 1 BWR RCPB BWR RCPB BWR RCPB HF.We&U'q error 2 SS RCPS HF:WekUg Error 3 7 4 RCPB HF.WeklN eror I I RCPB HF.WVeng error a 8WR HCPB Hot caradN I 11 SWR SS RC-P8 iGSCC.- Inwrr, " SCC t 4 2 11WR 11 Iss 2 aWR aas 2 4 BWV aS RCP BGSCC=-k*arg arSCC I SSi~_ tGSCC. k° grr 203 3 174 1 22 3 VWR I SaS BWR SS RCP8 TGSCC- TranS anua SCC I I RCPB' ' 1 TGSCC - lfra -nslarta SCC 2 H' SS T05CC. I r&.'S=arr 5.CC BW RCPB 3 i BWR I SS RCP8 "hern1*aIFeta, 2 2 6WR SS RCPB TIhemWFalque 3 1 1

BWR I S RCPB V-bratralw¶ e 4 F I BWR Ss RC BWR SS RCS-INSTR TGSCC - TkU&rw. SC - 2 1 i 6WR S$ RCSINSTR 105CC -TransgwasuSCC 2 1 BWR I BWR I 1

4 BWR SS SIR C*;oTaJon CoroUon, 3 BVWR 6s SIR T1 11 CExlerrnl C4'*o-%*0 C.Ekrnal Chloride k Erosion1 -2 8WR SS SIR SS __SIR_-------

1 td Corrosion 4 2 t BWR aS.

BWR 6S BWR SS SWR 6S SIR Fabgue 6 OWN SIR HF.CONST/ANSY HF.CONSTANST HFCONSTANST HF.CONSTANST 5 I 8WR aS SIR HF:FatneC.aon Eror" 6 2 BWR HF.FcWn*bon Erro" HI.YHumwanerror BWR S SIR HF:Ksron error HF.Weo" Error 2 4;

-- _2 4 I: -I 1 p Error 6 6 t 2  !

BWR Gs SIR SCI E"J 2.2 3___11 ___ 1

--- - m - - - --- - --- ----

- --- ~ - - - ~ -. - -- - - -

BWR SS SR lole(Smua SCC 6 22 4 OWR SS SIR MIC- M ccrovlogChiy nlro.ed CoIoon 1 SIR C*erpie.izat0on 6 t 1I 2

BWR 2 Sev&OOVWIGl14NI BWR SS Severe overai~aw I BWR Ss SIR Severe overloacn 6 1 I I

TGSCC. T~4",ue'SCC 1 3

BWR as SIR 3 3 SBWR SS STR Therma lague 6

'BWR ... S SIR TmeI-FaIMe -

BWR SS Urvepored 5 SIR 8SWR 6S SIR Wrason.-FOM* 0 BWVR SS SIR t Vrawe-f abue BWR 63 ,SIR Vorabonkaugus 2 27 2 I 1 1 I 21 SWR as SIR 3

BvWR SS SIR 4

BVVR SS* SIJR 5

BWR &S SIR 6 1 WaCv-1a11"e 1 BWR CS STEAM C*TOavo 2 1 _ ¶ BWR CS STEAM BvVR CS STEAM BWR CS S1 EAM BVER CS STEAM FAC. FjowAxrelalw1 C~orowon 2 15 3 1 FAG- FlwACCtIuale CwOIT0on 3 FAC- FlDwACe.rateld CoTosl51o 4 FAC. FlowAC00WO~d COlTOZOn ,

BWR CS STEAM FAC - FlowA~ce..*.od CrrOrswn 6 I 1 BWR CS 2 BWR -CS CS BWR CS STEAM BWR STEAM STEAM STEAM BWR CS STEAM HF:Wek*i 1" m G 6 1I BR CS STEAM HF.W Error BWR CS STEAM werne&Oeloa*Nq 4 I -

BWR C5 6 BWVR 6 CS 2 4 I 2 4 BVVR STEAM ITGSCC - TrerTrW SCC 2 2

3 STEAM Thoelralfdl ,ew 6 1 I BWR CS a oFab ue V*t.br ... .. I 2 Vitabo-fel-. 2 12 2

Vmrat-Fsabe I 6 1 1 Watsr Hanme 6 BWR ,,i CS WaW__Hamrny 6 BWR -~-CS

Appendix B Haddam Neck PWR CS 2.25 4 Erosion GL 89-08 CANDU PWR CS 4 4 Thermal Fatigue Korean CANDU PWR CS 4 4 Thermal Fatigue Korean CANDU PWR CS 4 4 Thermal Fatigue Korean CANDU PWR Cs 4 4 Thermal Fatigue Korean Millstone Unit 3 PWR CS 6 5 Erosion/CorrosIon IN 91-18 Arkansas Nuclear One Unit 2 PWR Cs 14 6 Erosion IN 89-53 DC Cook Unit 2 PWR CS 16 6 Erosion Bulletin 79-13 DC Cook Unit 2 PWR CS 16 6 Erosion Bulletin 79-13 Fort Calhoun Station PWR CS 12 6 FAC IN 97-84 Surry Unit 1 PWR CS 30 6 Not yet determined IN 81-04 SurryUnit 2 PWR CS 18 6 Erosion/Corrosion IN 86-106 Trojan 1 PWR CS 14 6 Erosion IN 87-36 Zion 1 PWR CS 24 6 Human Factor IN 82-25 FR (Framatome Reactors) PWR CS 10 6 Corrosion Korean FR (Framatome Reactors) PWR CS 28 6 Corrosion Korean

• i, ý.:Diablo Canyon U nitT,'.,. "IPWR-..

P R :"-CS-1.-:,!

CN_92-20S,`

.- ',-'rhermal Fatigue ":*'Jý

,.,. Lovilsa Unit 1 -  :.PWR -: ýCS-, , "  : Erosion/Corrosion;:;-:--IN

" 9*-

9118-'

,. .. . .quy .Un I PW R.... .,.t... .,.i.CS-,*. *;. -.*-ThermalFatigue , .. <IN 92-20A*t:.z

~'.~-SurryUnit 1 ~W~C ~ . i '.rsolorso Wolf Creek PWR SS 0.25 1 Vibration IN 89-07 KSNP Korean Standard Nuclear Power Plant PWR SS 0.375 1 Thermal Fatigue Korean Oconee Unit 3 PWR SS 0.75 1 Mechanical Failure IN 92-15 WH-3 PWR SS 0.75 1 Flow Induced Vibration Korean WH-3 PWR SS, 0.75 1 Flow Induced Vibration Korean H.B. Robinson Unit 2 PWR SS 2 3 SCC IN 91-05 Oconee Unit 2 PWR SS 2 3 Vibration IN 97-46 Prairie Island Unit 2 PWR SS 2 3 SCC IN 91-05 WH-3 PWR SS 2 3 Flow Induced Vibration Korean WH-3 PWR SS 2 3 Flow Induced Vibration Korean

• WH--3 PWR SS 2 3 Flow Induced Vibration Korean Crystal River Unit 3 PWR SS 2.5 4 Fatigue IN 82-09 Fort Calhoun Station PWR SS 3.5 4 SCC IN 82.02 Maine Yankee,. PWR SS 3.5 4 SCC IN 82-02 Maine Yankee PWR SS 3.5 4 SCC IN 82-02 Maine Yankee PWR SS 3.5 4 SCC IN 82-02 Maine Yankee PWR SS .3.5 4 SCC IN 82-02 Maine Yankee PWR SS 3'5 4 SCC IN 82-02 Maine Yankee PWR SS 3.5 4 SCC IN 82-02 Ginna PWR SS 8 5 SCC IE Circular76-06 Foreign . PWR SS 8 5 Thermal Stress Bulletin 88-08 Arkansas Nuclear One Unit I PWR SS 10 6 SCC IE Circular76-06 Oconee Unit 2 PWR SS 24 6 Erosion IN 82-22 Sequoyah Unit 1 PWR SS 16 6 Fatigue IN 95-11 Sequoyah Unit 2. PWR SS 10 6 Human Factor IN 97-19 Sum Unit 2 PWR SS 10 6 SCC IE Circular76-06

=3uLq MRad .~ - =m m m m

M m M M- M M M M M M M M n M M M Appendix B (cont.)

__ _ __Plant

____ _....._ _ Type

__ _ Group Material Diameter Pipe Size Failure Mechanism Reference Dresden Unit 2 BWR CS 4 4 Human Factor Bulletin 74-10 Nine Mile Point Unit 2 BWR CS 8 5 Fatigue Event 36016 Vermont Yankee BWR CS 12 6 SCC IN 82-22 Cooper Station BWR SS 0.25 1 Vibration IN 89-07

. Pilgrm BWR SS 1 2 Corrosion IN 85-34 Browns Ferry 3 BWR SS 4 4 SCC IN 84-41 Browns Ferry 3 BWR SS 4 4 SCC IN 84-41 Nine Mile Point Unit I BWR SS 6 5 SCC Bulletin 76-04 Dreseden Unit 2 BWR SS 10 6 Thermal Fatigue IN 75-01 Dreseden Unit 2 BWR SS 10 6 Thermal Fatigue IN 75-01 Dreseden Unit 2 BWR SS 10 6 Thermal Fatigue IN 75-01 Dreseden Unit 2 BWR SS 10 6 Thermal Fatigue IN 75-01 Dreseden Unit 2 BWR SS 10 6 Thermal Fatigue IN 75-01 Hatch Unit 1 BWR SS 22 6 SCC IN83-02 Hatch Unit 1 BWR SS 22 6 SCC IN 83-02 Hatch Unit I .... BWR SS 22 6 SCC IN 83-02 Hatch Unit I BWR SS 22 6 SCC IN 83-02 Hatch Unit I ..... BWR SS 22 6 SCC IN 83-02 Hatch Unit 1 BWR SS 20 6 SCC IN 83-02 Hatch Unit 1 BWR SS 24 6 SCC IN 83-02 Montecello BWR SS 22 6 SCC IN 83-02 Montecello BWR S§ 12 6 SCC IN 83-02 Montecello BWR SS 12 6 SCC IN 83-02 Montecello BWR SS 12 6 SCC IN 83-02 Montecello 8WR SS 12 6 SCC IN83-02 Montecello BWR SS 12 6 SCC IN 83-02

..-. BroWns.FerrydenU .1 i :, ,:", .BWR."4 - . . ,'..*: ;. : ?,t'  :',:. *<:'- IN 9 4- ,),

iHighlighted plants ,were :notused in the-dataanha1ysis due to missing information.. 7T;,,i.;&ý;:--..I.'.'

Appendix C. Collapsed OPDE Database Collapsed OPDE Raw Data as function of Pipe Size Plant Type Pipe Size Group Resulting Number of Failures (inches) CS SS CS+SS 0.0-1.0 154 544 698 1.0-2.0 74 154 228 2.0-4.0 78 75 153 4.0-10.0 126 112 238 I _ _

> 10.0 Total 93 525 126 1011 219 1536 I 0.0-1.0 118 257 375 1.0-2.0 2.0-4.0 32 32 75 227 107 259 I

4.0-10.0 50 234 284 1A" - . .-.-." -.

> 10.0 Total

" '-

39 271

,.

291 1084

',.. .. "..--...

330 1355 I 0.0-1.0 1.0-2.0 2.0-4.0 272 106 110 801 229 302 1073 335 412 U

522 PWR+BWR 4.0-10.0

> 10.0 Total 176 132 796 346 417 2095 549 2891 I

B I

I

I I

Collapsed OPDE Raw Data as function of Failure Mechanism Plant Type Failure Mechanism Resulting Number of Failures Fa Platype __rMchais CS SS CS+SS Corrosion 106 28 134 FAC 119 121 240 MIC 43 1 44 Erosion 96 12 108 Fatigue 92 501 593 PWR Human Factors 36 126 162 Mechanical Failures 22 37 59 SCC 5 169 174 Water Hammer 0 2 2 Misc 6 14 20 Total 525 1011 1536

.,-..:.-; , ..... .*:...-4. . .. ,I. ,..--.',:,:,..-::<:;-::  ::<.i:.*

• Corrosion 29 32 61 FAC 58 63 121 MIC 6 1 7 Erosion 40 9 49 Fatigue 71 225 296 BWR Human Factors 24 85 109 Mechanical Failures 18 25 43 SCC 19 624 643 Water Hammer 2 1 3 Misc 4 19 23 Total 271 1084 1355 Corrosion 135 60 195 FAC 177 184 361 MIC 49 2 51 Erosion 136 21 157 Fatigue 163 726 889 PWR+BWR Human Factors 60- 211 271 Mechanical Failures 40 62 102 SCC 24 793 817 Water Hammer 2 3 5 Misc 10 33 43 Total 796 2095 2891

c This Exhibit Contains' Proprietary Information

This Exhibit Contains Proprietary Information NEC-UW_18 191 ii)920 Calvert ClS U it I Infe ,NK. approvet a -1. percent increase in tre maxinummiicensea power ievel.

19770926 Calvert Cliffs Unit2 The NRC approved a 5.5. percent increase in the maximum licensed power level.

19790625 Millstone Unit 2 Tlie NRC approved a 5 percent increase in the maximum licensed power level.

19-790629 14. B. Robinson urnit 2 Ilhe NRC approved a 4.5 percent increxse in the maximum licensed power level.

198t00815 Forti Calhtoun Unit I The NRC approvedJ a.5.6 percent increase in the maximum licensed pover level.

19811123 StLucie Unit I Thi, NRC approved a 5.5 percent increase in the maxinmum licensed power level.

19850301 St. Lucie Unit 2 The NRC approvred a 5.5 percent increase in the maximum licensed power level-19850327 Duane Anold The NRC approved a 4.1 percent increase in the maximiim licensed power level.

19860206 Salem Unit I The NRC approved a2 percent increase in t(ie minximum licensed power level.

19860825 North Anna Unit 1 [lie NRC approved a 4.2 percent increase in the maximum licensed power level.

19860825 North Anna Unit 2 The NRC approved a 4.2 percent increase in the nmaxiiutm licensed power level.

19880330 Callaway Unit 1 The NRC approved a 4.5 percent increase in the maximum licensed power level.

19880726 Three Mile island Unit I The NRC approved '3 L3 percent increase in the maximum licensed power level.

19920909 Fermi Unit2 The NRC approved a 4 percent increase in the niaxitnin licensed power level.

19930322 Alvin W. Vogtle .Unit I The NRC approved a 4.5 percent increase in the maximum licensed Power level.

19930322 Alvin W. Voetle Unit 2 lhie NRC approved a 4.5 percent increase in the maximum licensed power level.

19931110 WolfCreek Unit J The NRC approved a 4.5 percent increase in the maxinini licensed power levcl.

19940411. Susquehanna Unit 2 The NRC approved a 4.5 percent increase in the maximum licensed power level.

199410(18 Peach Bottom Unit 2 'Ili, NRC approvk ad5 percent inc~rease in die maxinmum licensed power htve.l 19950216 Limerick Unit 2 "he NRC appro *eda 5 percent increase in the maximum licensed power level.

F1.950222 Susquehanna Unit I The NRC approved a 4.5percelt increase in the naximini licensed Power level 19950428 Nine Nife Point Unit 2 lI. NRC approved a4.3 percent increase intdiemaxinmum licensed power le.vL 199505112 Columbia Generating Sta lThe NRC approved a 4.9 percent increase in the maximuns licensed power level.

199t15071.8 Peach Bottom Unit 3 The NRC approved a 5 percent increase in the maximum licenisedpower level.

1995080(3 Sury Unit 1 The NRC approved a 4.3 percent increase in the maximum licensed power level.

1.950803 Sum., Unit 2 The NRC approved a 5 perceta increase inthe maximium licensed power level.

1.995(1831 Edwin L Hatch Unit I The NRC approved a 4. prcent increase in the niaximumlicensed power level.

19950831 Edwin f. Hatch Unit 2 "se NRC approved a 5 percent increase in the maximum licensed power level.

19960124 Limerick 'Unit l The NRC approved a 5 percent inc.rease in the maximum licensed power level.

199610412 Virgil C. Summer The NRC approved a 25percent increase in the maximum licensed power level.

1996)523 Palo Verde Unit 2 The NRC approved a2 percent increase in the maximum licensed powetr level.

19160523 Palo Verde Unit 2 The NRC approved a 2 percent increase in die maximum licensed power level.

The NRC"appoe a eceticreseithmxiumicndlwelee 19960523 Palo Verde. Unit 3 The NRC approved a42 percent increase in the maximum licensed power level.

19960926 Turkey Point Unit 3 The NRC approvtd a 4.5 percent increase in the maximum licensed power level.

19960926 Turkey Point Unit 4 The NRC approved a 4.5 percent increase in tie maximum licensed power level.

U.nit 1 The NRC approvt'ii a 5 percent increase in tie maximum licensed power level.

19961101 Brunswick 199.*611011 Brunswick Unit 2 The NRC approved a 5 percent increase in the maximum licensed power level.

I I

I approvec a 4 percent increase in tile maximum licensed power level.

19980429 Joseph M. Farley 19980429 Joseph M. Farley Unit I Unit 2 The NRC approved a 5 percent increase in the maximum licens4ed power level.

The NRC approved a 5 percent increase in the maximum licensed power level. I 1998090S Browns Ferry Unit 2 The NRC approved a 5 percent intrease in the maximum licensed power level.

19)80908 Browtns Ferry 19980916 Monticello Unit 3 The NRC approved a 5 percent increase in the maximum licensed power level.

"Ahe NRC approved a 6.3 percent increase in lie maximum licensed power level.

I 19981022 Edwin L Hatch Unit I 'lTe NRC approved a 8 percent increase in the maximunm licensed power level.

19981022 Edwin L Hatch 19990930 Comanche Peak Unit 2 Unit i.

Ilhe NRC approved a 9 percent increase in the maximum licensed power level.

Tihe NRC approved a I percent increase in the nilaximum licensed power level.

I 20000509. LaSalle County Unit I 'Ilie NRC approved a 5 pcrcent increase in the maximum licensed poweir level.

200(,0509 LaSalle County 20000N611 Perry Unit 2 Unit I The NRC approved a 5 percent increase in the. maximnu licensed power level.

Tie NRC approveda 5 percent increase in the maximum licensed power level.

I N2(01006 River Bend U..'-nitI The NRC approved a 5 percent increase in the maximum licensed power level.

2000)1026 Diablo Canyon 200l101.19 Watts Bar UJnit I Unit I The NRC approved a 2 percent increase in the niaxinum licensed power level.

The NRC approved a 1.4 percent increase in the maximum licensed power level.

I 200105-04 Braidwood Unit I The NRC approved a 5 percent increase in the maximum licensed power level.

20010504 Braidwood Unit 2 The NRC approved a 5 percent increase in the maximum licenseAd power level.

20010504 Byron 20010504 BvTon Unit..

Unit 2)

Tise NRC approved a 5 percent increase in the maximum licensed power level.

T'he NR approved a5 percent inrease in the maximum licensed power level.

I 2 0 10525 Salem Unit I The NRC approved a 1.4 percent increase in the maximum licensed power level.

2 00 10525 Salem Unit 2 The NRC approved a 1.4 percent increase in the maximum licensed power level.

20010706 San Onofre" 200 10706 San Onofre Unit 2 Unit 2 The NRC issued license amendment 180 increasing the maximum reactor power level to 3.438 megsawatts from 3,390 megawatts.

Tlhe NRC approved a I.4 petrent increase in the maximum. licensed power level.

I I

20010706 San tnofre Unit 3 The NRC approved a 1.4 percenit increase in the maximum licensed power level.

200 1.0706 San Onofre Unit 3 T'le NRC issued license aenesidment 171 increasing the maximum reactor povwer level to 3,438 meagawatts from 3,390 inegaoawattts.

200110706 Susquehanna Unit I The NRC approved a 1.4 percent increase in the maximum licensed power level..

20010706 Susquehanna 20'010719 San Onofre U-nit 2 Unit 3

'lhe NRC approved a 1.4 percent increase in the maximum licensed power Ievel.

The NRC issued license amendment raising maxinmum reactor power level to I

,,438 megawatts.

.20,010730 Hope.Creek 2D(0109241Beaver Valley Unit I Unit I The NRC approved a 1.4 percent increase in the maximum licensed power level.

'llie NRC approved a 1.4 percent increase in the maximum licensed power level. I 20,10924 Beaver Valley Unit 2 "he NRC approved a 1.4 percent increase in the maximum licensed power level.

20)0 1112 Comanche Peak 2(101 PO12 Conianchepeak Unit't Unit 2.

'Ite NRC approved a 1.4 percent increase in the maximum licensed power level.

'he NRC approved a 0.4 percent increase in (ie maximum licensed power level.

I 2,N) 110112 Slhearon Harris Unit 1. The NRC approved a 4.5 percent increase in the maximum licensed power level.

2)011106 Duane Amold The NRC approveil a 15.3 perctnt increase in the maximum licensed power level.

I I

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E,+J I ."UWC-ULLOI(I UC II !NIJUIIsSUCttel auIiUlvilt

ýLvLIJ* I MIL., IIICUMULLI enIaxmIII U UIC reIactu power le vl Ito 1,912 megawatts.

20011221 Dresden Unit 2 The NRC approved a 17 percesnt ihcreaae in the maximum licensed power level.

20011221 Dre"den' Unit 3 The NRC approved a 17 percent increase in the maximum licensed power level.

20011221 Quad Cities Unit I "lleNRC approved a 17.8 percent increase in the maximum licensed power level.

20011221 Quad Cities Unit 2 The NRC approved a 17.8 percent increase in the niaxitnum licensed power level.

20020329 Waterford Unit 3 The NRC approved a 1.5 percent increase in the maximum licensed power level.

2 0 020405 Clinton Unit I The NRC approved a 20 percent increase in tile maximum licensetd power level.

20020412 South Texas Project Unit I Use NRC approved a 1.4 percenl increase in the maxinmum licensed power level.

20020412 South Texas Project Unit 2 The NRC approved a 1.4 percent increase in the nmaximum licensed power level.

2-K)020424 Arkansas Nuclear One Unit 2 The NRC approv'ex a 7.5 peicent increase in the maximum licensed power level.

2W020430 Sequoyah Unit I The NRC approved a11.4 pe-cenl inerease in the maximumnlicensed power level.

.20020430 Sequoyah Unit 2 The NRC approtd a 1.4 percent increase in the maximums licensed powec level.

l2002053l Brunswick Unit. 1 The NRC approved a 15 percent increase in dse maximum liensed power level.

20020531 Brunswick Unit 2 The NRC approved a 15 percent increase in the mnaximum licensed power level.

20021010 Grand Gulf Unit I The NRC approved a 1.7 percent inrcreaae in the maximum licensed power level.

20021105i H. B. Robinson. Unit 2 Thse NRC approved a 1.7 percent increase in the via.ximum licensed power level.

2,102l122 Peach Bottom Unit 2 The NRC approved a .1.62 percent increase in the maximum licensed power level.

2(021122 Peach Bottom Unit 3 The NRC approved a 1.62 percent increase in the maximum licensed power level.

.20021 126 Indian Point Unit 3 The NRC approved a IA percent increase in tlse nmaximum licensed power level.

S 2(0)21129 Point Beach tUnil 1 The NRC approved a 1.4 ptrcent increase in. the maximum licensed pown' level.

20021129 Point Beach Unit 2 The NRC approved a 1.4 percent increase in the maximum licensed power level.

2021204 Crystal River Unit 3 MTe NRC approved a 0.9 percent increase in t(ie maximum licensedpower level.

200212210 Donald C. Cook Unit I The NRC approved a 1.66 percent increa-sc in due maxinmum licensed power level.

20030131 River Bend Unit I "fhe NRC approved a 1.7 percent increase m the inaxiinns licensed power level.

20030204 Crystal River Unit 3 The NRC approved license amendment 205 increasing the maximuns reactor power level to 2,568 megawatts.

2()03050)2 Donald C. Cook Unit 2 TMe NRC approved a 1.66 percent increase in-the maximum licensed power level.

20030509 Pilgrim Unit 1 lhe NRC approved a 1.5 prcnent increase in the maximum licensed power level.

20030522 Indian Point Unit 2 The NRC approved a 1.4 percent increase in the maximum licensed power level.

20030523 Indian Point Unit 2 Tlhe NRC issued license anendment 237 increasing (temaximum reactoir power level to 3,114.4 megawatts.

200307085 Kewaunsee The NRC approved a 1.4 pýrcent increase in Use maximum licensed power level.

20030923 Edwin L IHatch Unit I The NRC approved a 1.5 percent-increase in thie maximuum licensed power level.

2003(h923 Edwin L Hatch Unit 2 The NRC approved a 1.5 perceni increase in the maximum licensed power level.

2W5'309-29 Palo Verde Unit 2 lhe NRC approveA a 2.9 pereni increase in tUe maximum licensed powe lerel.

20040227 Kewaunee The NRC-approved a 6 percntl increase in the maximum licensed power level:

20040623 Palisades TIhe NRC approved a 1.4 percent increase iii the maximunmlicensed power level.

2W041{)28 Indian Point Unit 2 The NRC approved a 3.26 percent increase in thie maximum licensed power ievel.

+?Lir*d;*v, h*l.. - [*;*4:'J ) -'f-I

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I 2(X)50324 Indian Point.

2W050415 Waterford LIlII Unit 3 Unit 3 i iC fi.t. aJ1U£-t.£ d j l'iCCltt ILLLM edC Iii ULl*C iilAtilfll i'IlCeIlaC powe- ICYcI.

The NRC approved a 4.85 percent increase in the rnlaximnm licensed power level.

Ilie NRC approved a 8 percent increase in the maximum licensed power level.

I 2.05[116 Palo Verde 2.0051 1t. Palo Verde 2(06030,2 Vermont Yankee Unit I Unit 3

'he NRC apprve.da-2.9 percent. increase in the maximum licensed power level..

The NRC approved a 2.9 percent increase in the maximum licensed power level.

Thlie NRC approved a 20 percent increase in the maximum licensed power level.

I 20060"522 Seabrook 2M)060711. R. E. Ginna Unit 1. The NRC approved a 1.7 percent increase in the maximunii licensed power level.

The NRC approved a. 16.8 percent increase in the ijiaximiull liceised power level. I 20(060719 Beaver Valley Unit I The NRC approved a 8 percent increase in the maximumin licensed power level.

20060719 Beaver Valley 20070306 Bro.vts Ferec Unit 2 Unit 1

'he NRC approved a S percent increase in dhe maximum licensed 1power level.

The NRC approved a S percent increase in the maximumii licensed power level.

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NEC-UW_19 VERMONT YANKEE NUCLEAR POWER STATION PP7028 ORIGINAL PIPING FLOW ACCELERATED CORROSION INSPECTION PROGRAM USE CLASSIFICATION: INFORMATION Implementation Statement: This procedure supercedes VY Procedure DP 4023 and use of the Vermont Yankee Piping Flow Accelerated Corrosion Program Manual, Revision 2a:, prepared for Vermont Yankee by Yankee Atomic - Nuclear Services Division.

Issue Date: 05/10/01 PP 7028 Original Page I of 15 NECO04914

TABLE OF CONTENTS 1.0 PURPOSE, SCOPE, AND DISCUSSION...........-....... ........................ 3 2.0 DEFIN.ITIONS................... * . 5 3.0 PRIMARY RESPONSIBILITIES....--...*.-.-....* ....... ....... * ..... ...... ....... .....* .. 5 4.0 PR O C ED UR E ....... .......... ............. .. . ................ ....... . ............... ........

4.1. 4 1.Program P o r mM Maneac".

ain tenance .............. ......-

................° .......: ........................ * ............°........................ -9 4.2. Initial Screening and Identification ofFAC Susceptible Piping 9 4.3. CHECWORKS Modeling ........................................ .. 9 4.4. Outage to Outage Activities 10

5.0 REFERENCES

AND COMMITMENTS .............................. 12 6.0 FINAL CONDITIONS 14 7.0 ATTACHMENTS 15 PP 7028 Original Page 20f 15 NEC004915

1.0 PURPOSE, SCOPE, AND DISCUSSION 1.1. Purpose The purpose of the Vermont Yankee Piping Flow Accelerated Corrosion (FAC) Inspection Program is to provide a systematic approach to ensure that PAC does not lead to degradation of plant piping systems and feedwater heaters. This Program Procedure controls the engineering and inspection activities performed to predict, detect, monitor, and evaluate wall thinning due to PAC at the. Vermont Yankee Nuclear Power Station.

1.2. Scope The scope of this program is limited to evaluation and inspection of plant piping systems and feedwater heater shells that could be susceptible to FAC.

FAC is known to occur in piping systems constructed of carbon or low-alloy steels, which carry water or wet steam. All plant piping systems have been screened for susceptibility to damage from FAC. A separate document titled "FAC Susceptible Piping Identification" has been developed to identify, on a line by line basis, the piping which is susceptible to damage from FAC. This documeni is maintained by the Piping FAC Inspection Program coordinator and is updated as required to reflect changes in plant operation and configuration.

There is no finite scope of piping components to be scheduled for inspection on a periodic basis.

Each refueling outage. inspection efforts will be optimized to focus on piping components which have been identified as wearing, or potentially wearing due to FAC. The components selected for inspection each refueling outage are identified using:

• Results ofultrasonic thickness (UT) inspections from previous refueling outages.

• Results of the CHECWORKS predictive software, which incorporates actual inspection data.

" Operating conditions at VY, which may indicate PAC damage is occurring.

" Operating experience and events from other plants.

Carbon steel feedwater, heater shells have experienced thinning and through wall leaks due to PAC. Vermont Yankee has replaced all low pressure feedwater heaters with new heaters constructed of materials resistant to FAC. The four remaining high pressurefeedwater heater shells are carbon steel. Long term monitoring of shell thickness for plant feedwater heaters is included in the scope ofthis program.

1.3. Discussion Following the December 1986 Surry pipe rupture the industry has worked steadily.to develop and implement monitoring programs to prevent the rupture ofhigh energy piping due to single phase erosion-corrosion (FAC). In March: 1987 INFO issued Significant Operating Experience Report (SOER) 87-3 which recommended that a continuing program be established at all U.S.

nuclear power plants, including analyses to predict wear rates and to plan and schedule periodic inspections. USNRC Generic Letter GL 89-08, requires all holders of operating licenses to provide assurances that a systematic program has been implemented to ensure that Flow Accelerated Corrosion does not lead to degradation of plant piping systems.

This Program Procedure (PP) controls engineering and inspection activities performed to assess the susceptible plant piping. This procedure defines the methods and criteria used in the evaluation and inspection of plant piping components which are susceptible to wall thinning due to FAC. The program is based on current industry practice and the latest EPRI recommendations (REF 5.4.8.).

PP 7028 Original Page 3 of 15 LPC#1 NEC004916

1responsibilyio tMLtenanfVce Departmend 1O:eral n~lhe litholfthiedw trhetr, only"11 o She'la le inSpI offeedwat ht II erdetýt ie ytecniino shcle(lis h and r inozle, coordinated to trpEi w)4Wforthe Mainlenaicil lot pirtes u hbuters alodpndn nt on e itio afthte fmaterialenlstbs, ruesuppotpantpjcTes ffimpnemn Overllpeation of theheedater shellawltbersfs otrmyeeried conbyth intrna dintionv oft shellsecinand nozzles rpib-lty of th Mainten sanc ShelloadRo ntezze ignpectinfeedwinthenanr f ill buret tsisodpnden on tecnionfthe tue heater ne rvniveMitemaln:tues tube supportdlaers. migmn repnibcli;IIMl(ty ofte ilMa pintnnce Depatment lld Shellandnozlel LSCTinspctIonsA offeedwae11iiheater 1 ill w be' ptit-i r, din colerlendplate, e io n o m hl is the I cootrdintedf throug the rogramnmasinbleano,mEngineer1i n theI iMai ten uraent JaDestprartiment. U inspectnafsionsofithevheter shells lbae per iormedin jc h !I al unt esiidddygl an c,: urrent tes*tn*geqf emeatenrntuesuner Prmeventie Maintenance *(PM) wor*mordes.

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ITC I, Tepiary purposelectporfong

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ýJIC Tinspectio ns achoug I I accomlaion thindusyompones fioud ithignieficatwl 0C losoto unr tsre p rIt replaced Ilip') are ifmoreMi- repa ir and 4

VAO~ii~epible pStIflTh ll.A colln-tac ýw ltcwic .Cl<"nA ý J elk I an sillfiafdCductiom ppin my ndiatf(-mrf)rh cfrubliiiia ,wnd and ft&tm4di Ue The primary purpose ofperforming UT inspections each outage is to locate piping components degraded3

• by FAG prior to the time that an immediate repair or replacement is reqnired. ThisPPC allows 7028'lisufficient lead Iag

()Ill\a *Id time for a planned replacement which will have a minimum impact on plant operation.

4aof15 Given the costs of inspection and replacement ofpiping components, a long term approach for mitigating.I the effects ofFAC taken under this program will be towards reducing component wear rates. To accomplish this, components found with significant wall loss due to FAG under this programi, will be preferably replaced with materials which are more resistant to FAG damage.I PP 7028 Original Page 4 of 15I LPC#N NEC004917

2.0 DEFINITIONS 2.1. Flow Accelerated Corrosion (FAC) A corrosion process that causes thinning of steel piping exposed to flowing water or wet steam. The rate of loss is dependent on several parameters, which include flow regime, service life, water chemistry, piping material, piping geometry, and hydrodynamics.

2.2. Program

A set.o f activities that benefit from the existence o fa formal, high level "Program Document."

Such documents are meant to provide for a common understanding of program depth, breath and technical bases as well as the responsibilities o fthe program owner and those helping to implement the

,c, program. "Program Documents" are typically created to ensure regulatory requirements are satisfied.

I

• They can also be used to layout the technical bases and personnel responsibilities related to complex, multi-departmental processes.

2.3. Program Owner The individual responsible for maintaining the program, program documents, and assuring proper execution ofthe program requirements. Each program shall have an individual assigned as the program owner. The appropriate Job title is determined by the responsible Department Manager.

A summary of expectations for the program owner are contained in Appendix A ofAP 0098 and shall be referenced in all Program Procedures.

2.4. Single-Phase Flow: The flow in the piping system remains in the liquid phase at all design and operating pressures and temperatures.

2.5.

ITwo-Phase Flow: The flow in the piping system may vary from liquid to wet steam. This depends on the operating pressures and temperatures and varies with the specific location in the piping system.

I 3.0 PRIMARY RESPONSIBILITIES Implementation ofthe tasks performed under this program involve several plant departments. The organization for personnel performing tasks under this program is shown in Figure 1.

3.1. The VY Design Engineering Mechanical! Structural (DE AMIS) Department is responsible for the Piping FAC Inspection Program. The DE MIS Lead Design Engineer (LDE) has responsibility for the overall program management and administration and, for structural evaluation ofthinned piping components.

3.1.1. Establishment and maintenance ofcriteria and procedures for evaluation ofthinned wall piping components.

3.1.2. Performing structural evaluations o f thinned wall piping components.

I 3.2. The Vermont Yankee Piping FAC Inspection Program Coordinator (FACPC) works within the I* Mechanical Structural (DE MIS) Department under the direction of the DE MIS LOE. The responsibilities of the FAC Program Coordinator are:

3.2.1. Maintenance of the Vermont Yankee Piping FAC Inspection Program Procedure and supporting documents to ensure that program meets commitments to GL 89-08 and the "Expectations of I Program Owners" as defined in Appendix A ofAP 0098.

3.2.2. Continual assessment ofFAC inspection program to insure program effectiveness.

3.2.31 Participation in relevant industry working groups, benchmarking with current industry practice, evaluation ofindustry events; and implementation ofrevisions, changes, and process improvements which result from the participation.

3.2.4. Establishment and maintenance of criteria for selection o fpiping systems and components susceptible to FAC and for maintenance ofthe "FAC Susceptible Piping Identification" document which screens all current plant piping systems and identifies piping susceptible to FAG 3.2.5. Establishment and maintenance ofcriteria for selection and scheduling ofcomponents to be inspected during refueling outages including: initial inspections, follow-on inspections, and scope expansion and/or reduction.

PP 7028 Original Page 5 of 15 LPC#l NEC004918

3.2.6. Establishment and maintenance ofcriteria for use and control ofthe CHECWORKS predictive software used to evaluate piping, plan inspections, track inspection results, wear rates, piping component data, and repair and/or replacement history.

3.2.7. Review ofdesign change and maintenance documents as necessary to assess the impact ofthe proposed tasks on the inspection program, and'recommend action when appropriate.

3.2.8. Ensure that all physical and operational changes or additions to plant piping systems are incorporated into theprogram.

3.2.9, Analytical evaluation ofplant piping systems for FAC using the EPRI CHECWORKS codes as appropriate.

3.2. 10. Pre-outage activities including:

• Development of inspection scope for each refueling outage.

• Perform/update analytical evaluations (CHECWORKS models) as required.

Provide pre-inspection implementation support.

3.2. 11. Outage activities including:

" Providing engineering support for inspection implementation.

. Evaluation and disposition ofall inspection results.

" Recommend changes to the planned inspection scope upon discovery oftmacceptable conditions.

Providing assistance as required in the development ofrepair/replacement options.

• Providing written summary ofinspection results to ISIPC prior to plant startup.

• Ensure that cognizant departments and the Control Room are informed of unacceptable conditions discovered during evaluation of inspection results and facilitate completion ofappropriate paperwork (ER's, WOR, IDR, etc.).

3.2.12. Post-outage activities including:

• Development of outage inspection report including trending analyses and long term recommendations.

• Update/maintain the plant CHECWORKS models and maintain a history of all piping inspections.

• Update/maintain 'FAC Susceptible Piping Identification" document to reflect plant changes as required.

3.2.13. Keep DE MIS LDE informed on the progress ofFAC related tasks.

PP 7028 Original Page 60f15 NEC004919

33.. Th-e-V fi-iont o Yankee In--Service Inspection Program Coordinator(ISTPC): works within the System Engineering Department under the direction ofthe SuperintendentofSystern Engineering. The responsibilities ofthe ISIPC include:

3.3.1. Provide for overall coordination with the Vermont Yankee In-Service Inspection Program if inspection results on safety class piping indicate violations ofthe piping design code.

3.3.2. Coordination ofpre-outage activities including:

• Input to the development ofoutage schedules and budgets relative to FAC activities.

P providing oversight ofwork order planning and coordination with ISI Program resources.

• Arrange on-site services as required.

3.3.3. Coordination ofoutage activities including:

• Ensure components scheduled for inspection are properly prepared and accessible.

• Perfonnance ofinspections.

• Postinspection restoration ofcomponents.

• Repair/replacement effort ofunacceptable components.

3.3.4. Interface with the cognizant departments, as needed to insure all safety related repair/replacement ISI examination requirements are satisfied.

3.3.5. Ensure that required piping repairs and/or replacements are perfonned according to plant procedures and repairs to safety class piping and components are performed in accordance with, ASME Section XI requirements.

3.3.6. Ensure that cognizant departments and the Control Room are informed ofunacceptable conditions discovered during evaluation ofinspection results and facilitate completion of appropriate paperwork (ER's, WOR, IDR, etc. ).

• 3.3.7. Ensure that inspection records are temporarily stored per AP 6807 and permanently stored per AP 6.809 and available far the plant lifetime.

3.3.8. Keep the Superintendent afSystem Engineering informed on the progress afFAC related tasks.

3.3.9. Provide technical advice on implementation and inspection aspects ofthe FAC program.

3.3.10. NDE procedure development and maintenance.

PP 7028 Original Page 7 ofl 5 NEC004920

3.4. Level ill I1I51 Supervisor is a certified Level III UT examiner and works un~hdert i&Cihof the ISIPC. The responsibilities ofthe Level i IIIISI Supervisor include:

3.4.1. Review ofapplicable NDE procedures used in pipe tJT wall thickness measurements.

3.4.2. Ensuring that UT inspectors are properly qualified and trained to the applicable inspection procedures.

3.4.3. Review ofinspection results for compliance to the applicable procedures.

3.4.4. Resolution of anomalies found in inspection data.

3.4.5. Recommendations for augmented or special NDE procedures or techniques as required.

3.4.6. Direct supervision ofinspection personnel to ensure that the inspection personnel accurately and efficiently execute the inspection plan, complete inspections, and appropriately document inspection results.

3.4.7. Control ofall inspection data during the refueling outage.

3.4.8. At the completion ofinspections forwarding all inspection records to the ISIPC for pennanent storage per the requirements ofSection 6.2 3.5. Non Destructive Examination (NDE) Personnel 3.5.1. Meet Applicable qualification Standards. Personnel perfonning ultrasonic inspections shall be qualifiedto the requirements ofNE 8043.

3.5.2. Perfonn assigned setup, calibrations, and examinations.

.3.5.3. Documentation ofresults in accordance with approved procedures.

3.6. Plant Support Services' The Project Engineering Department is responsible for providing staging, lighting, insulation removal, surface preparation ofpiping components, and for component restoration after inspections are performed. Activities are controlled through the VY Work Order process in accordance with plant procedures.

PP 7028 Original Page 8 of 15 NEC004921

4.0 PROCEDURE 4.1. Program Maintenance The FACPC shall maintain the Yankee Piping FAC Inspection Program Procedure, PP 7028 and supporting documents to ensure that program meets commitments to GL 89-08 by:

4.1.1. Continual reassessment ofthe piping FAC inspection program to insure program effectiveness. A FAC Program SelfAssessment shall be performed at least once per operating cycle.

4.1.2. Participation in relevant industry working groups, benchmarking with current industry practice, evaluation ofindustry events; and implementation ofrevisions, changes, and process improvements which result from the participation.

4.1.3. Adaptation ofcurrent or developing industry practices: for selection and scheduling of components to be inspected, follow-on inspections, scope expansion and/or reductions, and criteria and procedures for evaluation ofthinned wall piping components.

4.1.4. Review design change and maintenance documents as necessary to assess the impact of the proposed tasks on the inspection program, and recommend action when appropriate.

4.1.5. Incorporate all physical and operational changes or additions to plant piping S'stems into the program as applicable.

4.2. Initial Screening and Identification ofFAC Susceptible Piping 4.2.1. A screening and evaluation ofall plant piping systems for susceptibility to FAC shall be performed. The screening shall use the EPRI Guidelines from reference 5.4.8., industry experience, and previous Vermont Yankee inspection results. The evaluation shall be performed and reviewed by engineers with FAC experience and familiar with plant systems. The resulting document shall be controlled by the FACPC.

4.2.2. The FACPC shall revise the "FAC Susceptible Piping Identification" document as required to reflect changes in plant operation, piping configuration, and/or materials.

4.3. CHECWORKS Modeling 4.3.1. Evaluate the susceptible plant piping systems for FAC using the EPRI CHECWORKS code. The evaluations shall be performed, reviewed, and documented per the requirements ofAppendix D.

PP 7028 Original Page 9 oi 15 NEC004922

4.4. Outage to Outage Activities ..

.Inspection and evaluation efforts performed under the program follow a cyclic pattern. Once inspection data from a given outage is obtained, it is incorporated into the appropriate predictive model and the results are then used in conjunction with otherFAC related information to establish the inspection scope for the next refueling outage.

NOTE Each large bore piping component within the scope ofthis program has been given a unique identification number as described in Appendix A. The location (building and elevation) of each large bore component is obtained from the Component Location, Sketches in Appendix A. Small bore piping inspection locations included in the program are identified in Appendix B.

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The tasks performed each refueling outage to implement the piping inspections under the FAC inspection program are detailed below. These are also broken out chronologically in a flow chart

• included here as Figure 2.

4.4.1. The outage inspection scope is determined by the FACPC using previous inspection data, the results ofthe CHECWORKS models, industry experience, and the guidelines contained in Appendix E.

4.4.2. The outage inspection scope is reviewed by the ISIPC for impact on and conflicts with-the overall outage plan. The ISIPC will plan and organize the on-site resources required to implement the piping inspections.

I 4.4.3. A work package is assembled for each piping component or group ofcomponents. This package includes component location sketches, support requirements such as scaffolding, lighting, etc., surface preparation and gridding requirements, and any special inspection I

requirements as determined by the FACPC.

4.4.4. Prepare piping components for inspection.

4.4.4.1. As directed by the ISIPC, scaffolding, lighting, insulation removal, and surface preparation ofeach piping component to be inspected are performed by on-site services in accordance with the applicable plant procedures.

4.4.4.2. Surface preparation and gridding ofpiping components for inspection shall conform to the guidelines in NSAC 202L (reference 5.4.8.). Specific I

instructions for surface preparation are given in NE 8044. Specific instructions for gridding ofpiping components are given in Attachment A ofNE 8053, or as I further directed by the FACPC.

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pp 7028 Original' Page 10 of 15 NEC004923

NEC-UW 20 PP7028 Piping FACJnspectlon Program FAC INSPECTION PROGRAM RECORDS FOR 2005 REFUELING OUTAGE TABLE OF CONTENTS TAB Pages 1 FAC 2004-2005 Program EWC Program Scoping Memo & Level 3 Fragnet 2-5 (4 pages) 2 2005 Refueling Outage Inspection location Worksheets 6-19 Methods and Reasons for Component Selection (14 pages) 3 VYM 2004/007a Design Engineering - MIS Memo: J.C.Fitzpatrick to 20-37 S.D.Goodwin subject, Piping FAG Inspection Scope for the 2005 Refueling Outage (Revision la), dated 5/5/05. (18 pages) 4 VYPPF 7102.01 VY Scope Management Review Form for deletion of FAG 38-43 Large Bore Inspection Nos. 2005-24 through 2005-35 from RF025, dated 11/1/06 (6 pages) 5 2005 RFO FAG Piping Inspections Scope Challenge Meeting Presentation, 44 -46 5/4/05 (3 pages) 6 ENN Engineering Standard Review and Approval Form from VY for: "Flow 47-48 Accelerated Corrosion Component Scanning and Glidding Standard",

ENN-EP-S-005, Rev. 0. dated 9/22/05 (2 pages) 7 ENN Engineering Standard Review and Approval Form from VY for: "Pipe 49-50 Wall Thinning Structural Evaluation" ENN-CS-S-008, Rev, 0. dated 9/22/05 & VY Email: Communication of Approved Engineering Standard date 9/27/05 (2 pages) 8 EN-DC-147 Engineering Report No. VY-RPT-06-00002, Rev.O, "VY Piping 51 -69 Flow Accelerated Corrosion Inspection Program (PP 7028) - 2005 Refueling Outage Inspection Report (RF025 - Fall 2005) (19 pages) 9 Large Bore Component Inspections: Index and Evaluation Worksheets 70 - 327 (258 pages) 10 Small Bore Component Inspections: Index and Evaluation Worksheets 328 - 347 (20 pages)

Page 1 of 347 NEC037099

I ENN Nuclear Management Manual Non QA Administrative Procedure ENN-DC.183 Rev.1 Facsimile of Attachment 9.10 Program or Component Scoping Memorandum I

.

I 2004-2005 Program Scope Mem.o _ _

I Vermont Yankee - Engineering Department

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Pmrppl NtimhPr-Piping Flow Accelerated Corrosion (I-AC) Inspection Program 2004 &

I 12005 Prooram Related Efforts El Depar-tment: Desi gn Enigineerh~--_echanicaf /Structural Owner. s James fitzpatrick ---

I Backup: Thomas O'Connor Procedure No. PP 7 02 8 "*,Vermont Yankee Piping Flow Accelerated Corrosion

& Tille: Inspection Prooram Detailed Scope of Project(Explanation); Engineering activities to support ongoing I

Inspection Program 10 provide a systemalic approach to insure thai Flow Accelerated Corrosion (FAC) does not lead to degradation of plant piping systems. Currently" Program Procedure PP 7028 controls engineering and inspection activities to predict, detect, monitor, I

and evaiuate pipe wall thinning due to FAC. Activities include modeling of plant piping using the EPRI CHECWORKS code to predict susceptibility to FAC damage, selection of components for inspection, UT inspections of piping components, evaiuation of data, trending, I monitoring of industry events and best practices, participation in industry groups, and recommending future repairs and lor replacements prior to component failure.

• Expected to adopt a new ENN Standard Program Procedure ENN-DC-315 (which is currently under development with an accelerated deveiopment date of 6!30!04),

I E'gpxecte-d-Benefi-ts-AJ-us-tiffcatiom):_V-Y conimi-tted-oha-ve-an-effecti-ve pipingEAC inspRQtionr_

program in response to GI 89-08_

I Conseque-nces of Deferral: Possible hazards to plant personne_l Loss of plant availability, unscheduled repairs, and deviationlrom previous regulatory commitments. I Duration of Progqarn: Life of plant

!004 Key Deliverables Of Milestones: Completion Estimate I

Compiete Focused SA write up & generate appropriate corrective 6/18104 actions' (coordinate ack'ities with program standardization effor*sý..-.

Completion of RFO 24 documentation, write and issue RFO 2004 1nspection Report 7/23104 I.

Software QA on XP platform for CHECWORKS FAC module Version 1.0G 1 8/13/04 I

-TuTMIMTr-!) Oiutage IrispecuonL acope, ricluuilyg coplrlig worksheets.

Update Piping FAC susceptibility screening'To account for piping and 97'1/04 8/13/04 I

drawing updates_ Include effects from NMWC, power uprate, & life extension.

I Update pn abasend devetop new priority 10/01/04 iogic for inspection scheduling, I

Page l of2 I

NEC037100 I

ENN Nuclear Management Manual Non QA Administrative Procedure ENN-DC-183 Rev.1 Facsimile of Attachment 9.10 Program or Component Scoping Memorandum U,.

Completion Estimate Jpdate CHECWORKS models using Version 1.OG with latest 2002 12/31/04 RFO & 2004 RFO Inspection data (Note ideally results are to be used In determining the 2005 inspection scope, however schedule nilestones override pp rOa logic,

,doption ofENN-DC-315 ENN Standard FAC program

• rocedure to include all previous improvements identified 30if Assessmenls.

1

ngoing Program Maintenance, Includes: procedure revisions, 12t3 t/04 3rogram improvements, benchmarking, attendance at industry (EPRI HUG) meetings, evaluation of industry events (industry awareness)

Ifl ffects on VY, license renewal -rolect Irnut, and .feet su.ppo.

25 r--iest nes --------------

Perform Proqram Self Assessment minimum once "er cvci[.. 411/05 C.;version of CHECHWORKS1,OG models to SFA-..VersiOll .. . 2,.....

lx 911/05 RFO 25 sup art 11115/05 Completioll of RFO 25 documentation, develop RFO 25 Outage 12/31/05 IlIsnection Report Ongoing Program Maintellance, Includes: procedure revisions, program improvements, benchmarking, attendance at industry (EPRI CHUG) meetings, evaluation of industry events (industry awareness) for effects on VY, and fleet suwoort.

r- I JIssue 2005 Outage Inspectioll Report 1/15106 I tp qFA Priclicrfivi Mncw*lg with ?NN.F RFf rfqfq All1/ I Ar ml p Ongoing Program Maintenance. Includes: procedure re 12/31/06 program improvements, benchmarking, attendance at industry (EPRI CHUG) meetings, evaluation of industry events (Industry awareness) for effects on VY, and fleet sunnort.

Estimated Bud Wet or Ex enses: AmountlHrs Ca tured in DE Mech.IStructural Base Bud at N/A others impacted B Pro ect ,Esimt ~dHours Desi n En ineerin FlUid S tems En ineerin 40 Electrical!l&C En ineerin a Mechanical! Strucwral Desi n

---

3 F nel: Attached -----------------

Performance Indicators for FAC Program are contained in the Program Health Report (Attached)

Page 2 of 2 1' NEC037101

2004-2005 Piping FAC Inspe, .n Program Level 3 Fragnet YEAR 2004 {2 nd half) (Time Line from 6J01 04 to 12131,04 /

I-Preparer Reviewer TOTAL Est, Est. Delivery Task No, Task Description (HRS} (HRS) (HRS) Sta rt 1 Completion Estimated Estimated. Estimated. Date Complete Focused SA write up & generate appropriate corrective actJions (coordinate actvites Wth program standardizationr 20 10 30 611/04 6/18104 efforts).

Completion of RFo 24 documentation, write and issue RFO 2004 Inspection Report 60 30 90 6114104 7/23104 Software QA on XP platform fur CHECWORKS FAC modulo 04-3 Version I.OG 20. 10 30 1711/04 8/13104 Update Piping FAC susceptibility screening to account fur piping 1 04-4 and draving updates, Include effects from NMWC, power uprate, 40 20 60 7/12104 8113/04

& life extensinn.

& lieetnin Update piping Small bore piping database and develop new 04-5 priority logic for inspection scheduling. 40 20 60 9/6104 10101/04 04-6

-Upd ate UI*-IUVVUI-Kb models usir4-Version-JTO-fwilli latest.-

2002 RFO &2004 RFO Inspection data 160 8G 240 8/23104 12131104 Issue 2005 RFO Outage Inspection Scope. InclUding Seaping 04-7 worksheets. 40 20 60 812104 911/04

+ I*-

048 *DevelapmerlTadoption of ENN-DC-315 ENN Standard FAC program Procedure to include all 80 40 120 6(2/04 10131104 previous improvements idertitfed Self Assessments.

04.9 Ongoing Program Maintenance. Includes: procedure revisions, 160 , 40 200 W11/04 12/31104 program improvements, benchmarking, Attendance at industry (EPRt CHUG) meetings, evaluatLon of industry events (industry

Žwareness'-f--effects-en W. LR uroiect inout,-and-fleet suooorl TOTAL {From end of RFO 24 to December 31, 2004) 620 270 890 H RS 54 Page 1 of2 NEC037102

-- m -- m - m - ----- m m-- m - ---

M M M m M m m m M Mm nM M m M 2004-2005 Piping FAC Inspe, ,n Program Level 3 Fragnet YEAR 2005 (1/1/05 TO 12/31/05)

Preparer Reviewer TOTAL Est. Est.

Task No. Task D&scrlptlon (HRS) (HRS) (HIRS) Start Delivery I Estimated Estimated. Estimatec Completion 1"q fp Perform Program Self Assessment (minimum once p-.r cycle).

05-1 40 20 60 3/1/05 4/01/05 Conversion ofCHECHWORKS O.0G models to SFA Version 2_1x 05-2 360 160 540 411/05 9/01/05 RFO 26 Preparatlin &Outago SIppQrt 05-3 160 60 240 9/1/05 1111510504 05-4 Completion of RFO 25 documentation, develop RFO 25 Outge lnspedlion Report 90 12/31105 60 30 11115/05 05-5 Ongoing Prograrnain6eriace. Includes: pmcedure revislons, program rmprovements, benchmarking, a3ttndanC at irtdlstry 20 60 1/01/05 112/31/05 (EPRI CHUG) meetings, evaluatlon or industiy events (industry 40 1 1

-- - --- VF4 V I. ýttý I IVý%ýý - 1, Total Fh 990 J I I.

Page 2 of 2 NEC037103

Th~z I VY Piping FAC Inspection Program PP 7028 - 2005 Refueling Outage Inspection Location Worksheets I Methods and Reasons for Component Selection

• ly* (* L.-.-

A ~ ~~~ReviewedtV'-*...

...... -  ![ ..

I Note" RIsed for VY and g4dj_ Events and O on 311105 abzien~.e Piping components are selected for irispebtion during the 2004 refueling oUlage based on the following groupings and/or criteria.

Lame Bore ieping LA: Compollents selected from measured or apparent wear found in previous inspection results.

LB: Components ranked high for susceptibility from current CHECWORKS evaluation.

LC: Components identified by industry events/experence via the Nuclear Network or through the EPR] CHUG. U LD: Components selected to calibrate the CHECWORKS models, LE: Components subjected to off normal flow conditions. Primarily isolated lines to the condenser in which leakage is indicated from the lurbine p6rformance monitoring systoem. {through the Systems Engineering Group)_

LF: Engineedng judgment I Other LG: Piping idehtihed froth EMIAC Work Orders (malfuinctioning equip., leaking valves. etc.)

Small B46re FLhino SA: SusbopliN6 piping locations (groups of components) contained In the Small Bore Piping data base which haVe:; rnt received an initial inspeotidn.

S8: componriets seloieed trom measured or apparent wear found In previous inspection results.

SC: romp'6he.tsfd tdi*dLby i0dutVi eav'e sei' ence via the Nuclear Netwotk or thb0gh the EýP'RIlHUG.

S): Comp*rS-ortP subji.tcd to..f normal Ilow cohditions. Primarily isolated lines to the cond~nsor In Whiich lea1ooe Is [nditibd*.rfron 1th turbrie p0rfbrmance monitoring system. (through theSy-tes"Ehgiaerihg Group).

SE: Engineering Judgment! Other.

SG: Piping identified tram EMPAC Work Orders (malfunctioning equip., leaking valves, elc.)

________ Heater Shells I No feedwater heateT §hell Inspectlnris will beperlormed during the 2005 RFO. All 10 of the leedwater heater shells have been replaced With FAC resistant Materials,I I

I Page 1 of 14 NEC037104

VY Piping FAC InspactJon Program PP 7028 - 2005 Refueling Outage Inspection Localion Worksheets / Methods and Reasons for Component Selection LA: Large Bore Components selected(idenlified) from previous Inspection Results From the 19951199611998/1999120011200212004 Refueling Outage Inspections (Large Bore Piping) these components were identified as requiring thiure monitoring. The following components have either yet l0 be inspected as recommended, or the recommended inspection is in a future outage.

Inspect. Loc. ComponenllD Notes IComments I Conclusions No. SK.

q6-18 001 FD13EL05 199ý Repobrt: calculaled time to tmin is 11.5 & t2 cycles based on a 98-19 FD13SP06 single measurement. The 2005 RFO is 6 cycles since the inspection.

liT ins ectelbow and downstream 0l In 209Q8 96-36 002 FD02SP05 1996 fkeport: calculated time to Tmin is 9.5 cycleS based on a single measurement. The 2005 RFO is 6.cycles sinoe the inspection.

U..T _.n.*AUIp.et elb.ow cn dow.A, tream0i1e in 20V'?

98-37 005 FD07SPOI 1AN6 R port: 6Ealbuf6.d Uih6 10 Terh is !96 .cycles based on a Sitgle measureraent The 2005 REO is 6 cycles since, the inspection,

____________ t ekj0w0' flfltea

_Asec i"I 2007 96.39 005 FD07SP02US 1 fl"6d: "A16 t4d t 6 Tiri i 105&ycles based on a -ingIe measurement. The 2005 RFO is 6 cyolds since the inspection.

U0.lii i..aw Id44wflstIfaInl-elI 20.08 9$.0 005 FPZ0EL06 19V08.Repont: Sal8ul~f*4t~hrtotniin s 7:5& 6.76ycies bas- on a W,8-07 FD07EL07 strigle.measurement. The 2005 fRFO is 5 cycles since the iris0ction.

Gvenbno significaft wear lound i,n adjacent .compnents (R-SL =04.3 cyc.es -n FD07SPO.7) defer in~ectioln unlll R.A026. UT f00l6SOt eplboveH QLO7*40- a00d dw ra bi~8I 9913 011 FI'DO8EL04 1109" pot WiTmii. I T.7:9 & 12:5"y.ies 6t.'td dn a FO08SP04 single UIh inspectin. The 211 0.5 :RQ is 4 ciycles sin-ce the:fhjpe'6tton.

99162 011 FDOP040pe ) 10o Fepor: calulad t~iietto dn. .aslte Tmin ii 6.2& cycles based U~te, ln*petit ~thedlown.raUY under lpn'l Inib6l 2 B*PLI=

99-33 CND-Nz32-A sle mesUreti ent. Th e 20tRFO 5 is 4 cylesos incethinpco.

0208 01 FD14Eb01 2002 recohtrfendati0P to$Vnspett'hie at dWinps2r007oae mnonr *:in~

109-29 FDttSPO2S 2me4suGiven single o0,0eum05minth.lrr t.o-nlye The205RFO-h rad elbow dofwerestflrem ssp cyle sineo 4edc pipe i~nt2erb7e wtnspectimon 99-3 99-16C:D-rozat-A F.820180 99-32 019 FD06TEO 1(pipe cap) 1999 Reoport: calculated UT jnsn1telbow and lime deIn00 to Tmi n is ownstr 6-.2 &

fnln 200Cycles basod on a I 68.

99-36 001 04.03 FDncfTE0t CND-Noz32-C 204 ReomeQdto single measurement. worTheinspuedt 2005 AFe> tegaerine20 fsince the d.ef1autr is 4 cyclesbasdnthe inspection.

UTrinsp&1t w~ear loated elbow4and und0,0 downstream r tnh/ce, pip0 lii 2005S Relbw-tInspectuplra FD14 elbow lee&

99320 017 FDO4TE01 (ppeca) 04 Rprt:cammcudatedn toeimse tomin i620&611 yceth one autwa based 02-08 016 FD18ELO1 2002 rchentbnto inspec the elboW in ýýO07 based on a~sirigie 02-09 FiD15SFO2US measurement. Re-inSPeCCt elbow and do'wnstream pipe in 2007(3

_______cycles from 20021-04-03 001 FDOITEOS 2004 recommendation to inspect tee in 2008 based on the default wear rate of 0,005 inch/cycle, Re-inspect upstream elbow and lee in N 002 FD02RDOI 2004 recommendation to re-inspect in 2011 based on the default wear rate of 0,005 inch/cycle. Re-Inspect reduceor with downstream

______elbow and tee In 2007.

Page 2 of 14 NEC03710S

I VY Piping FAC Inspection Program PP 7028 - 2005 Refueling Outage Inspection Location Worksheets I Methods and Reasons for Component Selection I

LA: Large Bore Components selected(idenlifled) from previous Inspection Results -continued I Inspeot.

04.08 Lac.

001 Component 10 FOOZTE01 Notes/Comments I Conclusions Z004 recohmmendation to inspect tee in 07 based on the default I

wear rate of 0.005 inch/cyde. Actual point to poini measurements from 04-09 001 FO03SPOI 1999 to 2004 [ndicate no wear, Given EPU operation, re-inspect wih u streanil*bow and reducer In 2007.

200.4 recommendation to inspetet pipe secticn in 2011 based en a I

single inspection and the default wear rale of 0.005 inch/cycle. Re-04-10 001 FDQ7SP02DS inrsnect in 2011.

2004 recommendation to ihspect pipe section In 2008 based on a I

sin le ins ection. Re-Ins ect with downslr.eam elbow In 2Q08.

04-13 04-23 001 001 FD14EzL03 MSD9TtO1 to 2004 recdmmendai-n dn to intpect Row 13 pup piece1o OS volve In 2008 is based on a sinole UT insoection. Re-Insp~ect In 2008.

2004 recomrmendaticn to inspect pipe section in 2010 due to localized I

MS.9-TE0

._ wear directl under 2 lines. RoIn$p. t In 2010.

I

___.__ _

04-23 001 MSO9OL05 .2004racorhrriendltion to inspect plPe section in 2010 base on a single fnseetion. Re-)*psp't in 2016.

Turbin'e (2ip$3-around Piping:

Previous Intemal Visual UT & Repair History:

I I *,Tp.ac (V i ...

"..tn m. Tb " '.6 "ff SŽ'*0...

SFf1959 RF021 RFO022 RF023 S I

i~t z 8200.1 f,-0.2 ,.'

I

___________V_

1_: V V V V 00WS originial I:

Y V.UT, R

V VJUTI R

VAST Y

V V V I Ut P. :: r...

.r ... I-yWiT4' . V I

I NOTE; Reference Dwg. No. 5920-6841 Sh. 1 of 2 needs to be updated with oorreotlnformation. This will be performed during the EPU design ohange effort.

The HP turbine rotor was replaced in 2004. Internal visual inspection oj all four 36"dlameter lines was Perorrmed, An I

internal visual Inspection of the 30"C line (firsl inspection sincethe 1993 replacement) and the 30" D line W8s performed.

2005 RFO based on increased flows and the possibility of different flow regimes in both the 36 & 30 inch piping, I

perform a visual Inspection. LP tUfbine work in 2005 RFO may provide opportunity ]or access to the 30 " lines. As a

,minimum inspect (2) 36 inch lines and the carbon steel 30" B line.

I Page 3 of 14 I

NEC037106 I

VY Piping FAC Inspection Program PP 7028 - 2004 Refueling Outage Inspection Location Worksheets I Methods and Reasons lor Component Selection LB: Large Bore Components Ranked High for Susceptibility from CHECWORKS Evaluation The current CHECWORKS wear rafe calculations contain inspection data up to the 1999 RFO and wear rate predictions are current to the 2001 RFO. The 2001 and 2002 RFO inspection daja has been entered into the CHECWGRKS database. However, updated wear rate calculations are not cormiMte, and won't be in time to support the schedule date for issuirng Ihe inspection scope for the 2005 outage. Basedon a review ot the 2001 and 2002 RFO inspection data for components onlhe Feedwater, Condensate, and Heater Drain Systems, the CHECWORKS models still appear to over-predict actual wear. Nothing new or unanticipated was observed in either 2002 or 2004.

Feedwater $&siom Listed beWow are components which meet the followihg criteria:

a) negative tihe to Tmin frorf the predictive CHECWORKS runs which include Inspection data up to the 1999 RFO.

b) n'o inspections have been performed on these components or the corresponding components in a parallal train Since the 1999 RFO.

Comrpnent Location Location Notes l15 Sko.toh F Q7..L0.5 605 TB FPP Elev. 241 Conm ortts .- on o.th** train were rin 0.cted S-Oo'flto 006 V. Heater Bay Eisvs 228 Comro'nehts on 6ther tealn were Intcted In 1998.

FDO7EL11 & 248 Results indiGate minimal Wear. -AfterupdatIng the CHEOW.O-Ks mrdel with newer data, assess need

___foradditiqna.lio.pow~ris in 2007 RFs.

FDOIEL12 006 T.B Heater Bay Elev. 248 F:eýedýtr heater irplacemnnt occurred in 2004 RFO, Informal visual inspeotions of intemals and cut pipa profile inticated a stable red oxide and no distinguishable winr pa4eyr,.

FDPSTEt01 012 T.B Heater Bay.Clevs 228 htnerhih Wa6tcomponents FDOSEL06 &,FDQESP06 Weare FD68ELO7 & 248 rlsioect*d In 1988. Re~uats h'dicate minimal wear. After upda'ting OE-IWORKs.model With newer, data, assess nreed for inspecting components on the train vs..thb',

FID08aL4Y 0.12 T.B HWeater Bay E.e'v. 248 Feef t': heater replaeerment ocurred in 0044TRaO.

lnfor'mal vI.Ual inspef.*ons 01 internals a *nd*ot pippe prof1e indiat*,d a s red oxide and rno disting*Uishable 6t]e weAe piat_ srnI FD15ELO8 013 RX Steam Tunnel El. 266 intrnaV1bsuai <if elbow-erforme6d iP 1996 duirlifg 46tok vaivereplacement, no incdicaton of wall loss at thdt.tine.

Corresponding component on line 16"- FDW-14 was inspected in RF024. After updating CHECWORKs model with newer data, assess need for inspecting I this COomlonent in 2007 RFO, Page 4 of 14 NEC037107

I VY Piping FAC Inspection Progr=am PP 7028 - 2005 Refueling Outage Inspection Location Worksheets/ Methods and Reasons for Component Selection I LB: large Bore Components Ranked High for Susceptibility Irom CHECWORKS Evaluation - continued I Condensate System Only one component wasidentified as having a negative time to Tmin. This was CD3ITEQ2DS, the downstream side of a 24x24x20 tee on the condensate header in the teed pump room. The CHECWORK$ prediction tor the I

downstream side of the tee has a small negative hrs relative to the remainder ot the components in the system and relaftive to the upstream side of the same tee. Other tees on the same header have been previously inspedtod and show no significant wear. The CHECWORKS model includes UT data up to the 1999 RFO. The inspections on this I

system performed in 2001 indicate minimal wear. Components CD3STEO2 and CD30SP04 were inspectjd iii 2004. This dala along with the 2001 inspection data will be input to CHECWORKS to better calibrate the imodel.

Moisture See.arator Drains & Heater Drain System I

No compontnts idehtified as having negative times to Tmin. No components wete selected icr inspection in 2001, 2002, or 2004 based on high susceptibility. However future operation under HWC will change dissoived oxygen in system.- A separate evaluation has been perforrmed and components were selected for inspection in 2002. See U

Section LD below, Extraction Steam System I Three components on this system with negative time to code min, wall: The piping is Chrome-Moiy. ES4ATEOI &

E§4AI"t02, 30inch diameter tees inside the condenser have neggtqi'e prediction (-3426Hrs.) for trriie to gti6 wl The rheqpve times to tmin may be conservative based on the modeling 1t00n[ques used. Relinement of th'4i 3rnpd#lt.i0t6is I

s iprogross. The tiegtttive titi tol*lhi if most likely.a function of ta6k of inspoction data Vsý mWtifis-- ut .at utex.-.otehal lagpqlng on tills pipirng and the lWoation inside the condonser, no:coniponoents are sdl&.e*!"f~t UT -irins';pOtin in 20b4 based 66t high susceptibility. HoWever, ar oqp*p~rn!ty to perforti an lnttrna~lvis*UAl; iitfftpýQ Of all the Extraotion .Steaiti lires Inside the condenser during planed [P turbine work in !Iie 20bo5AFb rka'y: *-r*

na I

itSdf. Sde Section LF below.

Note the short section of straight pipe on line 12"-ES-1A at the connection to the 36 inch A cross atountd is. azssum-ned to beAl 06 Gr. B carbon steel is not modeled in CHECWORKS. This componenlwas inspected in 2W004 by eAf6Wl I

UT anid an internal visual inspectic0 from the 36" cross around line.-

I I

I I

I Page 5 of 14 NEC037108 I

VY Piping FAC Inspection Program pp 7028 - 2005 Refueling Outage Inspection Location Worksheets I Methods and Reasons for Component Selection LC: Large Bore Ccmponents Identified by Industry Events/Experience.

Review of FAC related Large Bore Operating Experience (OE) and/or piping failures reported since April 2003 Deseri tion & Recommended Actions at VY 8/9/2004 Mihama3 - 0E193681OE18895: RuptiUre of Condensate line downstream of restriction orifice.

PWR PWR system highly susceptible to single phase FAC due to low DO. Similar region of system as 1986 Surry evant (5 fatalities). Based on info gathered by INPO/CHUGfFACnet the location was omitted from previous inspections due to clerical error, once discovered management missed opportunllyto inspect and deferred inspection until 9104. TOO late. Lesson: inake sure all h~ghly susceptible locations get inspected. PWR Condensateffeectwater piping is much more suscdptible to single phase FAC than BWR with 02 injectioh. GiVen that, previous inspection hlslory, and condensate CHtCWORKS modeling; inspect pipihg bs of all flow orifioes rn the higher temperature condensate system that have not been previous!y inspected in RFO25. Inspect CDO3rEQl I QO"ELl.I / CD130$P02 in Rr.2. (re-peat lnspe'tlon from 1090). Aiso, inspect C. Firtoi eI.$-2EL64 /

10117103 Duane Arnold - OE1.7300: Through wall leak In 4" diaireter ohrome-moly Heater(.Drain System BWR bypass line to the condenser. The line was a temporary inst-allation due-t delayed FWD heater installation. The cause of the leak ,ppears to be droplet:imping4nent erosion due to use of a bypass control vilvo. The equivalent litlesat VY ard the Heater Drain bypass lines to the condopser downstream 61i-'. It* high leavtl hetrl valxes Teselin hvdMD~attaehed'to -rttionftor 1e66k.00qe int1ih idrftdeer

• 0TPM sytem). ome ispe-iins have b~en p6tdrrhd4 ..thIf."es Conalder for "i4/03 South Texas OSPYŽY Rfin~g &in~ternal6 wear fond -ondisct-iarge ipin c5it Odntkosh* tt Proiet - PWR Polishing System. Pipe, ieaton Ste(elIlowwater t8emprature"(.O to.1 4*F) neutral pH, and velocity of 12.2 Ft/,sec Tortuoues floW.th and 'controtValves, we.r.may-be mpin.ement, PWR system Low. drssclvedoex#yg. Equivatfntsyste. at VY is Con0tiisate Lihminerafizer System which is to6W .temnp ansdt*reeIspe6t N.S.Ac P2L

______as not: *I letPA lte"'r'irNoGEa 3W.~~~~t 11/07103 .faldweod 2- 0.1,4 W*l:tnlr1st. d O.FD .. .-*ish"C#6nt - {6i *rnaqnt0 sytem heitl'sry has 10w 0.00. therefbdre mro4 Sp fe6~t~t~ i 1, tQ- l la0 pufip disohafge nozzies.and dewhsteýaýrnpip'ng hiaV. mut0irnW*A*-*S*Io-dat, No further ac*tions arAratqjt from this 6E.'

10/31/03 Clinton _BWR -0EZ17412fOElg947: Throu.(h-wall leak-sin2_ATB h--1&-r Ven-lihes t-5ble-he (lager bore lines assumed given description of backing rings in piping). Apparerit cause attributed to steam jet impingement from wei steam. Equivalent line at VY is common 4 inch feedwater heater vent line 101 No.4 FDW heaters. This line is included in the SSB database since Itconnects to (2) 2-1/2" lines. Inspection priority will be determined in the small bore ranki! andydoriization.

1T119/03 Hope Greek - OE17700; Pinhole leak aid wall thinning in 8e in carbon steel.Extraction Stearl BWR suppiy iine to Steam Seal Evaporator. Location of wear is downstrean of pressure safety valves. Apparent Cause of leak &wear is due to liqUid droplet inipinpeMeht due to high flows from failure 01 pressure safety relief valves, No e4tiivalent confiouration at VY.

1/24/04 LaSalle 1 - BWR 0E171991 OE18381: Tough-waii hOles in extraction steam piping Inside condenser.

Location of holes at inlet nozzles to No.2 FDW heaters located Inthe neck oj the condensers (2nd lowest stage). All 12 nozzle are CS. with A335-P1 1 upstrearn piping. VY has only the No.5 FDW heaters in the neck of the condenser. The No.

5 FDW heaters were replaced with Chromo-moly shells. ES piping Is A335-P1 1 or e uivalent which is FAC resistant. No further actions are anticinated from this OE, Page 601 14 NEC037109

I I

VY Piping FAC Inspection Program PP 7028 - 2005 Refueling Outage Inspection Location Worksheets / Methods and Reasons for Component Selection I LC: Large Bore Components Identified by Industry Events/Experlence - continLued I

Date Pfant -- T e Descn tion & Recommended AClions at VY 2/17/04 Peach Bottom 2 BWR OE18637: Online leak inlO inch main steam drainlille header to the condenser.

Hole was located directly below the connection of I" main steam lead drain. The header was replaced with 1-1/4 Chrome material approx. 5 years before the leak.

I Also, ROs in steam drains were modified. The cause was attributed to steam impingement 'Additional information to follow after next RFO. The oniy large bore drain COllectorat VY is the 8 inch diameter low point drain header, line S"MSD-9.

Flow Is through steam traps and ICVs vs. a continuous flow through a restriction I

orifice. This line is now part of the AST ALT boundary, Inspections of the entire 8/26/04 PaeVerde 3-bottom 01 this headerwere performed during RF024 with recommendations for retieat insoedtions In 2010.

OE20S86: Through wall leak found on a 10 Inch flashing tee cap on the IP I

PWR feedwater-heater drains. Problems with inspection of flashing tees in program. Only 14 QUi 011.53 susoeptible tocations have UT data at Palo Verde 1,2,3. There are no flashing tees 0.8. of LCVs on the heater drain system at VY. The onty flashing toes U

al VY are located on the FWD pump min flow lines at the condsnspr. in*dpexoin 01

-9124f04WL Palisades- PWR all.3.lines FP A 4 WE-F and) 6".F..WEPit§ s -schedduled for R F02,5.

OGiM94-: Wall thinning In carb6n steel Extraction steatn piping. Increased Iocali2ed wear downstream of Bleedertrip valve, Equivalent piping at VY Is

- I ExtractioR Steam piping downstream of the reverse c.urrent valves. ES piping at VY

-9/18/04' Catawaba 2 -

PWR is A.5-P i lwIdh is.AC rislanft N6futth'er &btidd is f.ei uifed IQt1tiE* GE.,

OEi9sO: W0a-lthinning~feuoTir diff'ie'tareiasorn FDW piping. Two areas-are not considered specific to Ctawba. 1)Area wher4 main fdedetafr by..s reg I

valves reonters the feedwator header and 2) dOWnstream 01 the main l1etN'ater reg valves. PWk feodwatfer system Ohemistryhas loW D.O. therefore me susceptible to wall loss due 10 single phase :FAC than BWR feedwater piping. At VY area 1) doses -not exist (bypass linros$ump to the cbndenseO 2) Inspections havbdNlen I

r*inc*." upptr*am of.both mainlfed rep, valves. hrft!a tiq'n of pnrfqtnied.. aq.do'?

r:iu*'si. doynstreaM a.re sdeai~ I.:ft:*6$.5 NbO foi*ther aittc5* re I

11/3/04 Ojane Arnold - 0'1t'Wfl fir tciTfiifiifoCWnstr&aIl orus Gooling Test Re4ti#:HeOT.Isdiation BWR Valve. Apparent cause was c8aviat!n erosion duo to lhrotting invalve &fhng!Jq4F'C I

& b10 ..teting. At VY, the-equiVaIeht valVes are V1O-34A &.348. Tre"d&.e-of oavltalien.presnt Is depafdent of the system design and may vary fiorm-Pl&ht to I

plant, Previous UT inspections were performed on valve bodies and ddwftýtreati 216105 Calvert Cliffs 1 -

reducers in early 90s. No significant wear was lound. Consider inspectfon of downstream pipi9 In RF026 if additional OE warraflts It.

OE2Qt27: Through-wal1 leak in 6 inch steam vent header for MSR rain tank. VY I

dlnc* nnt hauc4q £*. rnnficnuratinn. NO Mnistiire Senarator Re-heaters

-2117/0.5 PWR Clinton ,BWR doesnoth- Prri-6tonfinorntion NO Moisture Separator Re-heaters OE20246, Qatastrophlo failure 01 turbine extraction steam line bellows inside condenser, Found through-wall holes ES piping OS 01 bellows due 10 FAC.

I Apparent cause was attributed to the steam jet Irotin the holes inducing vibration of the expansion joint that led to high cycle latlgue lailure. At VY extraotion steam piping inside the condenser is A335.P1 1 or equivalent which is FAC resistant. No further actions are anticipated from this OE.

U Grand Gull f- Pip Hole-Leak in 4 inch-carbon steel elbUVV in KHN min tlow line. System has low BWR use at VY <<2% of time). (Perry also found thinning at elbow per C.Burton at CHUG meeting.) A review of VY drawings VYI.RHR.Pari 14 Sht.ill and VYI.RHR Part 15 I L Sht.1 It show elbows downstream of restriction orifices. Previous VY Inspections downstream 01 orilices on HPCIfand CS systems found no problems. Keep OE listed for future conslderali.oo.

I Page 7 of 14 N EC0371 10 I

VY Piping FAC Inspection Program PP 7028

• 2005 Refueling Outage Inspection Location Worksheets / Methods and Reasons for Component Selection LC: Large Bore Components Identified by Industry Events/Experience- c,onllnued Date .. PIanI*T e Descri ion & Recommended Actions at VY 9124102 IP2 -PWR Pin hole leak on 26 1/2N2 cross-under piping (HP 10 MSR) in vicinity of dog bones at expansion joint under location of weld overlay localized Wear under/around a previous weld overlay repair. VY has solid piping (no expansioo joints). Visual Inspections of 30' B CAR carbon steelip-pin. will be performed in 2085.

1/15/02 surry 1-PWK Leak in 8 inch Condenser drain header for 3 pl. FDW Heater vents. Also CHUG thinning in Gland Steam Pipidg inside the condenser and Ihel 2" Condensor Drain Meeting header from MS Drain trhp lines. The only large bore drain collector at VY is the 8 inch diameter low point drain header, line 8WMSD-9. This line is now part of the AST ALT boundary. Inspections 01 selected components on this line Were pertoimed during RF024 with recommendations for repeat inspections in 2010 (Section LB above). Given this line is part of the ALT Boundary Inspect approx. 2 fl. long sctAion at condentser wall durinLjq AO2f026 t 7' or RFO2Y f2O.

LD: Large Bore Components Selected to Calibrale CHECWORKS The CHECWORKS models have been upgraded 10 inciude the 96, 98, & 99 RFO inspection data. The 2001 and 20-2 ingp0otion data has been loaded however wear rate analyses have not been completed allhis time.

InZl001 corhpinents 011 the higher lemperature end of the Conden.sale System were Inspoced to calibrate !tie CNEOWOI3KS nki*4. The inspection data indicate minimal wear- and should reinloroe the assbesshie6t of low wear in the Con:4densateý ýýtem. Additional compon;ents selected for inlpedtfon in 2004 in Section LB above will be used 10 rliibiiateihe CHECWORKS modet HEatir.Dfnain.Moiftre Setarator Drains:

P6*ik[ 164t6*.22 REO there was limited inspection data for the Heater Drain system. The current CHECWORKS mfi~dt~i*Pe (b  ! tidtb. e Pass 2) indicate low wear rates, During 2002 a number 01 new inspections Weire p-tifr*d tn arbo steeopiping upstream of the level control valves (LCV) 40 obtain it baslirne:prWt6 6lh1* h operation otibydthe*p wAter chemistry. PipIng down slream 01 the LCVs is rAC resfstant material excfpt o0r ihlet l1 No.5 Feedvater" heaters. No additional components on the Healer Drain system will be inspected in 2005.

Eesdwater No inspeotions on line 18".FDW-'2 have been inspecled: Inspect FD12EL06 and FD12SP08US in 2005 Main.Steam Only 2 components In the Main Sleam system on line 18'MS-7A in Ihe drywell have been inspectedlO date. Inspect MS1 DEL07 and MS1DSP13US In 20105. (Nole this also addresses a license renewal consideration tor monitoring of Main Steam Piping).

Page 8 of 14 NEC0371 11

I VY Piping FAC Inspection Program PP 7028 - 2005 Refueling Outage Inspection Location Worksheets {Methods and Reasons for Component Selection I

I LE: Large BOTe Components subjected to off normal flow conditions Identified by turbine performance mpnitpring system (Systems Engineering Group). I The Systerhs Engineering Production Variance Reports for 2003 listed the "B1 and "C' !eadwates pump min flow valves as leaking into the condenser. There are sections on carbon steel piping at Ihe connection "0 the condenser on all three lines. As /li minimum inspect the '(8" and "C" lines In 2005. I There have been concems with cavitation at condensate min low valve FCV-4. An internal inspection 01 the valve petformed in RFO 24 showed some damage to the valve internals. However, due 10 a leaking isolation valve the c*.nhaect[ng piping was 110000d and an internal visual inspection coutd not be performed. UT Inspect the upstream I

and'do0wnstream piping during RF025, The valVe is operated during outages and startup at relatively low temperatures for FAC to occur. The piping is un-Insulated and close to the l100r. No insulation removal or scalolding will be rdquired. I Since startUp from 2004 (RF024), no other leaking valves or steam trapS have been identilled (to date) using Ihe Tubine Petformance Monitoring (rPM} system. However, it new data Indicates leaking valves then, additions to the Outage scope may be required. I I

LF: Englneetrng Judgment /Other I

Nine AME Soelion XI Class 1 Category 8-J welds are to be inspected by the FAC program per Code Case N-560.in I10.,j 01 a $ection XI volumetric weld inspection. The VY ISI Program Interval 4 schedule for inspedf6in of thlse Welds Is as tallows: I JusTnriF Outage Section Xl Description i-AU; Vrogram L;omponeflIs ISr PreIgtaI Weld 10 rw1e9-M3B upstr*afn.Pipe to tee "A" Feedwater on Sketch 010 I

FW¶9-F30 tas-to reducer FD19TEO1 I

M -,.,,

ftierval 4 FW1-F4 redtcer to pipe FD1lU0D0 P~eti;hd 1, FW21-F1 tee to pipe FD1OSP-4 Outage 1. F021SPOt Fall 2011 (RF029)

Interval 4 FWI8-3A FW20.3A upstream pipe to tee tee to reducer "8" Feedwater on Sketch 016 F018TE01 I

Period 3, FW20-F1 reducer 10 pipe FD20RD01 OUlage 6, FW20-F1B FW18.F4 horizontal pipe to pipe teo to hibe FD20SPOl F01.8SP64 I Continued I

I Page 9 of 14 I

NEC0371 12 I

VY Piping FAC Inspection Program PP 7028 - 2005 Refueling Outage Inspection Location Worksheets I Methods and Reasons for Component Selection LF: Engineering JUdgment! Other -continued Extended Power Uprate (EPU)

Eeedwater system:

EPU evaluation tor Feedwater System: The primary focus of Work to date (for PUSAR and RAIs) was on velocity changes given only-slight inoreases iAtemps and no chemistry changes. With all 3 FOW pumps running the 16 inch diameter lines to Ihe 24 inch FDW header have approx. [1.2(213) = 0.80120% reduction in velocity, Velocities In the remainder 01 the sysfem Increase approx. 20%. The highest velobites are at the 10 Inch reducers upstrkni and downstream of ihe FOW REG valves. The expander and downstream piping have mUllIple inspection datA.With FD07RD03/FDQ1SP03 last inspected in 2001 and FEI08RD03/FDOBSP02 last inspected in 1999. Both ot these se:gmerits should be re- inspected after soie lime of operation at EPU flows. Assuming EPU startilngeatly in 2005d, inspect components FDOSR[DO3 & FD08$P02 In 2005 to obtain an up to date pre-EPU measureltLit.

Inspect FD07RD031 FDO7SP03 In 2007 for a post EPU measurement.

Cond~ens~ae_ Sy~ttem*

Given the 8104 Mihama event: consider addiUonal component in lhe condensate system for inspection:

downstream of flow orifices & venturies:

FE-102-4 and downstream pipe on 24"C-8 venturi type (TB condensate pump 100m overhead) Given low oper(atihg te1'erg&tures and upstream of oxygen injection point, scope oui and evaluate for inspectio'i In RFb261 n 2007 FE-52-1"A to FE-SZI.E on Condensate De-mlnei(Rlzer Sy.tem (Restriction Orifices). Gven low opiratfng termeratur.s end upstream 01 d6,yfdn, injection point, scope out and evaluate tor inr$ptction in RF0.62 In 2;007 FE-i02-7and doWnsfrearh pipe on 14"C-21 venturi type TB Heater Bay E1237.5 Given low operiating tetfperatures and used for start-up, scope. out-And eVueV t.fOr it$sPectinh in RF026inn0 FE-1 02-2A 06. 0)-3"il0C 9-0ed10 thO ITYPPM pmlA (venturi typej) Previously 1b~fl inspected tn 19ý89 Petnet:FE aind doWntstream pipin.-1i1RF025

.-

(2- ort.:O0O-31,-V5c*[ in the lb 'FPR above FDW Pi1p 1 B (venturitype) No prevMous inspection data. Inspeot F* and dewn*stream pfping In *fO25 FE-1 02-2C on 20"C.32, located in the TB FPR above FDW pump 1C (venturi type) Previously inspected in 2001 All Extracflon Sleam piping is A335-P1 1, a. 1-114 chrome material, except for a short carbon steel stub piece in line 12'-ES-1A af Ihe connection to the 36" A cross around line. An inlernal visual inspection ollhis stub piece has' performed with the cross around inspection in RF024. Also an UT inspection of ESIASP01 was performed in RF024.

Extraction Steam piping in the condenser has external lagging which requires significant effort Ior removal when performing external UT inspections (piUS there are significant staging costs). The piping is A335.P1 1. However an

_ipportunity fo perform anlnternai visual inspecticn of all the Extraction Sleam lines inside the condenser during planed LP turbine work in the 2005 RFO may present itself.

Page 10 of 14 NEC0371 13

I VY Piping FAC Inspection Program PP 7028 - 2005 Refueling Outage I Inspection Location Worksheets I Methods and Reasons for Component Selection I

LG: Piping Identilled from EMPAC Work Orders (malfunctioning equip., leaking valves, etc.)

Word searches of open work orders on EMPAC were performed [or Ihe following keywords. trap, leak, valve, replace, repair, erosion, corrosion, sleam, FAC, wear, hole, drain, and inspect. No previ.0osly unidentified components or I

piping were identified as r*'quiring monitoring during the Fall 2005 RFO.

Note: the internal baffle piate in Condenser B Ior the AOG train tank return line to Ihe condenser is 10 be replaced in I RFO 25 (ER 04-1454/ ER 05-2321 ER 05-0274). Erosion on baffle plate is from condenser side (not piping side).

Internal visual inspection 01 LCV-103-3A-2 during RFO 24 indicated some type of casting flaw. The System Engtrteer suspects possible leaking by the normally closed valve. The downstream piping was last inspected in 1990. The line i

typiically has no flow. Re-evaluate using the Thermai Performance Monitoring System Data and cOIlsider inspection 01 downstream piping in RF026.

Through vwall leak in the steam seal header supply line ISSH4 discovered on 9/24/04 (CR-VTY-200402985). A i

temporary leak enclosure was Installed and a planned permanent repair is scheduled for RF025. The feaks are on the bott6im of un-insulated piping upstream of tha gland seal. Field Inspection of the leak location shows that the piping at the leak sloping doWn 10 the gland seal, not sloping up to the seal a shown on the design drawings. UT data on the top of the piping nearthe leak shows tull wall thickness. At this time, the exact mechanism which calased the I

leak is nol knoWn. Additional inspections t;o determine the extent ot condition on the 3 other gland seal Mteam supply lines are required I

Insp.et the 90 degrde elbow andapprOx. 2 ft. of downstream piping on lines 1SSH3; 1SSH4, IS.SH5, and 1S4'0"duri hg: R"-FO 0 .. . Also bas.d on Industry OE and sirmilar piping geometry, inspect 2 of the sPE lines (1 SPE3 And 1-ft5.dUring AFO 25. I I

I I

I I

I I

I Page 11 of 14 NEC037114 I

VY Piping FAC Inspection Program PP 7028 - 200S Refueling Outage Inspection Location Worksheets f Methods and Reasons for Component Selection Small Bore Piping SASusceptlble piping locations (groups of components) contained in the Small Bore Piping data base which have nol received an Initial inspection.

Locations on the continuous FDW heater vents to the condensetr on the No.3 heaters were inspected in 2002. The continuous vents on the No.4 heater were installed new in 1995. The start up vents operate less than 2% of operating tine. No wearwss found in irevious inspections on Heater Vent piping from the No.1 & 2 heaters. Given that and the lower pressure in the No.4, shells a complete Inspection of the remainder oi the No.4 heater vent piping can be deferred. The existing small bore date base and the piping susceptibility analysis is under revision. No additional compenents from Revision 1 of the data base will be inspected.

SB;Compothlets selected from measured or apparent wear found in previous inspection results.

Small Bore Point No. 20. 211/2. MSD-6 @ connection to condenserA at Nozzle 33 (Inspection No. 96-8B01 identified a low readingat weld on stub to condenser). Upstream valves are normally closedt TPM system does not indicate any abnormal flow. tnspect this piping In REO 26 Athrough wall leak in the turbine bypass valve chest 1ýt sealleak-olf line form the No. 1 bypass vales occurred in 20fl. (Vt Evenit Report 2o00-04:4). A Iteporar leak enclosure %W:.Slhst~led (TM,2003-002) to contain the leak}.

W:O. 0,3-2364 was writtin to inspecl/repaleffd ace/1ine. A localjzed iike-forlike (cparbon steel) replac~eteht of the leak locationwa8 perf&mad in .RFO 24. Additional inspections on this line rderftifiea 0Icawlized WUI loss and one adaifional like-for-like repair was performed. EngiMering Request ER 04.0963 was widttehlfo cotipletely replace this pipig With hiob-rn-oly piping. (Dresdan has already done this). The repltacdemj'ent (ER 04,0964) iS curiteilly stfibduled for,RFO 26. If this ad6ivi[y gqits 'de-ýsooped" thea, additionalin gpectlons Will be retlired to insute the piping is Ae**it"ble' for cni'tnuoed operatoan.

Page 12 of 1.4 NEC037115

I VY Piping FA,C inspection Program PP 7028 - 2005 flefueling Outage Inspection Location Worksheets f Methods and Reasons for Component Selection I

Small Bore Piping SC: Components identified by Industry events/experience via the Nuclear Network or I

through the EPRI CHUG.

I Date Plant-Type Description & Recommended Actions at VY 11/7/2003 Limerick 1, BWR 01217818: Through wall leak In 1 inch drain line back to condenser oft 12S piping at the conneclion to the large bore line. Normaliyno flow in line due to N.C.

valve. Piping downstream of valves to wndenser on all 3 lines was scheduled I

for replacernont. Location US of valve was thought not to be susceptible.

12$ piping at VY is FAC resistant A335P1 1 with no drains back to the condenser. Lesson from this event is any carbon steel line In a Wet steam system is susceptible & should be monitored. Also fuft line replacemenrt insures I

all susce ptLbQe Bjpqis r*oj U

__¢,__f1 1116/04 Clinton BWR 01217654: Potdbtial tend tor adverse equipment condition downstream of orifices. (Ref. Previous experience a Clinton with CRDpump min flow R08) anscect ORD oumD min flw orifices also Dinino 0S .01 RO-64-2 in frF025 12/08/04 V.C. Summer- 0E1C.9798: Coniplete failkre ola 1 indh E line at the l&6ationi of . pfeviciUsly PWR Installed Fe(manite clamp repair. Previous leak a,1 weld instaled in MAY 2004.

See presentatfon at January 2 0 0 5 CHUG meiting. (They did not do UT on the I

_..ie._ assuws~,tructyral integtv c.rior to instaifnwtheciann) 311105 McGui,e 2-PWR

-Thou'h-wa~ Id-akin-a-2inch *c-arboW t Vent fe on theMb- heating steam vent line. Caus..*.by FAC when tlashing'ocourrdd upstream of Rd (design J9 ogj) No. M$-.RS or ec uivel~nt I!cation at VY.

I 4/29799 Darlington lI - v..* l6*.it at s .Ae.n teooh ]*6h t:i "*kthreaded connectiOn. Equivyiat to PHWR HHS sytern at VY.(INPOK Event 931-990429-1) Threaded connections..typicahly oh 6onednsate side of HHSpiSipingi. Lower energy/co6sequence 61 166k.' Include HHSplpiripin FAC SuStibIfty.Fivlew, and in the Small Bore Datdba'ie.

I 6/14/99 9/1101 Datljrigtoli 2 -

PHWR PFath Bottom I LUak oh steaiir tf0 di_-tcoarg.p*lgo -pipa sv-9ten at VY. INPO Even**P

. -A hed cbnneOtion. Equival~iht tb HHS 9ý.614-11 Same ad abbve.

(From 1114102 CHUG .Metihg) ma. on I lflh :Sab 80Q.:1e.e**i, il Ga- Re-I 3-BWA ~ reýi~wd AiG oiib small botWte o pltocQqdenser Pefrmadtif 8ta~ ~ ~

incud rAki FAA Suscitihih Revi~w. Updatelfgfrt6 an con Voens ot failure.

I Vi /6#2 Hatch 1f2 -BWR Coode~nserin Loakdge due 1 th6:oubgh Well erosion' (ýteý i) of'011/2 iodh "slop" OHUG Mt1. drains lines insid. the endensee. Lines In each unit were cut.and eapped similar events at ByTrn Unit I (0C 1t2609) and Columbia (OE1214'S). Urnerick &

Dresden. VY slop drain lines Inside condenser were walked down during I

RF024. Some external erosion or, oi ina and SUODOrts was found.

1/15102 CHUG Mig.

Catawba 2 -

PWR Leak in HP lurbine pocket shell drain 1 inch dia. OEM showed pipe as P-11.

However, A106 Gr. B was installed. Inspections were be performed on Ihis line In 2004 to base line condition rnlor to HP turbine rotor re lacement.

I 1/15/02 Dresden 2 Thinning found in Bypass valve lqak-off line to the 7 stage extraction stean CHUG Mtg. BWR line, Line is 2" Soh, 80, GE B4A39B. Lowest reading was 0.070" found using PhOsphor Plate radiog'aphy. Line was replaced with A335 P-11. Same line as I

2003 VY through wall ieak. Partial GS replacement was. performed In RFK124.

Pi inQ Is scheduled to be replaced with A335-P1 1 In RFO2.S (ER 04-0965).

I I

Page 13 of 14 NEC037116 I

VY Piping FAC Inspection Program PP 7028 - 2005 Refueling Outage Inspection Location Worksheets / Methods and Reasons for Component Selection Small Bore Piping SD:Comrp~onents subjected to off normal flow conditions, as Indicated froanthe turbine performance monitoring system (Systems Engineering Group).

No small bore lines have been identified by Systems Engineering on or bejore 3/1 105.

SE: Engineering judgment Look at piping DS at orif-ices based on BWR OE Condensate: Given the 8104 Mihama event: consider additional component In the condensale system tor inspection downstream of floW orific4es & venturies.

FE-102-6 and downstream pipe on 21rT'C-43 venlurllype (TB healer bay elev. 230+/- Given low operating temperatures and upstream of oxygen injection point, scope oul and evaluate for Inspftfin iii RZ6 in 2007 SG: Piping Identffled from EMPAC Work Orders (malfunctioning equip., leaking valves, etc,)

See LG above, The EMPAC search performed in LG above is applicable 10 both Large and Smaii compdnehts.

Page 14 of 14 NEC037117

MEMORANDUM I Vermont Yankee Design Engineering To S.D.Goodwin Date -May 5 2005 From James Fitzpatrick File # VYM Z004!007a Subject Piping FAC Inspection Scope for the 2005 Relueiing Outage (Revision la)

REFERENCES (a) PP 7028 Piping Flow Acceleraled Corrosion Inspection Program, LPG 1, 12/6/2001.

(b) V.Y. Piping F.A.C. Inspection Program -1996 Relueling Outage Inspectfon Report, March 23,1999, (cl V.Y. Piping F.A.C. Inspection Program - 1998 Refueling Outage lnspection Report, April 2,1999.

I (d) V.Y. Piping FAG, Inspection Program -1999 Refueling Outage Inspection Reporl, February 11, 2000, (e) V.Y. Piping FAG. Inspection Program - 2001 Refueling Outage Inspection Report, August 11,2001.

(f) V.Y. Piping FAG. Inspection Program - 2002 Refueling Outaga Inspeotion Report, January 20, 2003, (g) V.Y. Piping FAC. Inspection Program- 2004 Refueling Outage Inspection Report, February 15, 2005 (h) DiSCUSSION Altached please lind the Piping FAC Inspeotion Scope for the 2005 Refueling Outage. The scope includes locations identified using: previous inspeotion resuits, theCHECWORKS models, industry and I plant operating experience, input from the Turbine Performance Monitoring System, the CHECWORKS study performed to postulate affects 01 Hydrogen Water Chemistry operation on FAC wear rates in plant piping, and engineering jUdgment. 3 The planned 2005 RFO inspection scope consists 0137 large bore components at 16 locations, internal ihspection 01 three legs 01 the turbine cross around piping, and 5 sections 01 small bore piping.

Also, any industry or plant events that occur in the interim may necessitate an increase in the planned scope.

Iwill be available to support planning and inspections as necessary. Ii you have any questions or need additional information please contact me.

(Revision 1 identifies Small Bore Inspections due to IndUstry OEI.

'(Revision la adds component Nos. to SSH & SPE piping & oorrects inor typos in Attachment) am $ . Fitzpatrick D . n Engineering Mechanioal/Structural Group ATTACHMENT: 2005 RFO FAC Inspection Scope 3111/05 (3 Pgs) Revlsed 515/05 CC LLukeos Code Programs Supervisor OoKIng (0SI)

T.I.OConnor (Design Engineering)

Neil Fales (Systems Engineering)

NEC0371 18I I

m-m-m -m--- m m m m m M M M - M M mM ATTACHMENT tv. ,YM 2004!007a VERMONT YANKEE PIPING FAC INSPECTION PROGRAM 2005 INSPECTION SCOPE (515105) Page 1 0 13 LARGE BORE PIPING: External UT Inspections Point Component ID location location Previous Reason /Comments I Notes No. Sketch Inspections 2005-01 FD14EL03 008 T.B. Htr. Ba Elev.267. 1999 1999 recommendation for repeat inspection.

2005-02 FD14SP03US 008 1999 2005-03 FD04RD01 017 T.B. Htr. Ba Elev.24S. 1999 Inspect per 1999 calculated wear rate.

It IF ii 2005-04 FD04TE01 017 1999 It i! II 2005-05 Gond Noz32A 017 1999 2005-06 FD05RD01 01' T.B. Htr. Ba Elev.245. 1993 TPM system indicated leakage by normally 2005-07 FDOS TE01 018 1993 closed valve.

2005-08 Gond Noz 328 018 1993 010 T Nr Qg In~nc'trzt r lqQ r~Icii~t~

l %Ah\-Ar rnff Al--_n

• II il ...... I ,* .* ["1 O-LC-L--- FI -j 111 -11 M k! I 7F IUZ7 1-- 1 rilvi byb~teill IIIUI~dLfUU Iý-d~d~ge uy IUIId 2005-11 Cond Noz32C 019 1999 closed valve.

2005-12 FD08RD03 011 T.B. FPR Elev.231 1999 EPU flows increase 2005.13 FD08SP02 011 1999 2005-14 FD12EL06 007 T.B. Htr. Ba E!ev.264. NO Ghecworks Mode! Calibration. Asbestos 2005-15 FD12SP08US 007 NO removal required.

2005-16 GD30FE01 037 T.B. FPR E!ev.241 1989 FE-102-2A (Millama Event) 2005-17 CD30ELl 1 037 above "A" FDW pump 1989 2005-18 CD30SP12 037 1989 NEC037119

ATTACHMENT t% vfM 2004/007a POint Component ID Location Location Previous Reason I Comments / Notes No. Sketoh Inspections 2005-19 CD31 FE01 038 T.B. FPA Elev. 241 NO FE-102-2B (Mlhama Event) 2005-20 CD3 EL04 038 above 1B" FDW pump NO Asbestos removal required.

I 2005-21 CD31SP04 038 NO 2005-22 CD21RD02 040 T.B. Htr, Ba Elev.230. NO Inspect piping upstream and downstream of 2005-23 CD21RD01 040 NO FCV-102-4 (piping is not insulated).

2005-24 1SSH3EL05 ° Turbine deck at packing NO LP Turbine Steam Seal supply lines due to 2005-25 1SSH3SP06US 3 Htr, Bay Efev, 254. through wall leak at elbow on nne I SSH4.

2005-26 1SSH4EL01 Turbine deck at packing NO 2005-27 ISSH4SP02US 4 Htr. Bay Elev, 254. 'See markup 01 Dwg. 5920-1239 n _______0_iq~;P Tiirhinp dlarL n] n~ir'kinri No____________________

2005-29 1SSH5SP02US ° 5 Htr. Bav Elev. 254.

2005-30 1SSH6EL06 Turbine deck at packing NO 2005-31 1SSH6SP08US 6 Htr. Bav Elev. 254.

2005-32 2SPE3EL01 ° Turbine deck at packing NO IP Turbine SteamPacking Exhaust at packing 3 2005-33 2SPE3SP01 US ° 3 HIr. Bay Elev. 254. and 5 due 10 Ihrough wall leak at elbow on line 2005-34 2SPE5EL01 + Turbine deck at packing NO 1SSH4.

2005-35 2SPE5SP01 US 5 Htr. Bay EIev, 254.

'See Markuo of Dwn. 5920-1239 2005-36 MS1DEIO7 080 AX Stm Tunnel Elev. NO EPU and LR data required for Main Steam lines 2005-37 MISIDSP13US 080 2b410UZU NU LARGE BORE UT NOTES,

1. Coordinate minimum extent of insulation to be removed with J.Fitzpatrick or T.M. O'Connorirom DE-MIS.

2. A 'No"' in the previous inspeotlon oolumn indicates asbestos abatement may be required.

SPage 2 of 3 NEC037120

- - - n-m m-m- - - m- - - m m-m

- - - - - - n- m - - m m - m m ATTACHMENT tk, VYM 20041007a LARGE BORE PIPING: Internal Visual Inspections (With supplemental UT as required In. etion Point No. Deserl ion 2005-38 36" CAR A (36 inch diameter Line A Turbine Cross Around under HP turbine) 2005-39 36" CAR C (36 inch diameter Line C Turbine Cross Around under HP turbine) 2005- 4 0 30"CAR B 30 inch diameter Line- B Turbine Cross Around uqpepr east side of heater ba SMALL BORE PIPING Small Bore S.B. System Description Location Drawings Reason lComments inspection Data Number Base

-No.-

05-SBOt 11 Condensate 1" piping OS of R.O. 64.2 T.R Heater Bay G191157Sht,1 InduslryOE17654 5920- FSI -17 05-SB02 128 CRD I" Piping D.S. 01 R.O.-3-24A Rx. SW Elev, 232,5 G1911701G191212 Industry OE17654 1 P3E-IA IG191215 05-S803 12 CRD 1" Piping D.S. of R.O,-3-25A Rx. SW Elev. 232.5 G191170 1G191212 IndustryOE17654 P38-IA IG191215 05-S804 130 CRD 1'" Piping D.S. of R,O,-3.24B Rx, SW Eisv. 232.5 G1911701 G191212 Industry 0E17654 P38-1B IG191215 05-8805 431 CRD 1' Piping D,S, of R.O.--3-25B Rx. 8W Elev, 232.5 G191170 1 G191212 IndustryOE17654

_ 1 P38'1B IG191215 Page 3 of 3 NEC037121

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TURBINE BULPD!N-FEED PUMP ROOM/HEATER .DAY REFERENCES, G1I1 157,GI91 82,G 191163 5"9S0"FS-124,5B920-FS-CS5 REVISION It 11124/91 VERMONT YANKxLr r'illut eROSION-CORROSION INSPECTION PROGRAM FEEDWATER lINE 4'-FDW-4 0%

COMPONENT LOCATION SKETCH No.017 4-App*cr.dbxA FP70ZS-digius Pa etof102 NEC037123

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AOQ~ VI'T-IViUl I YAINRE- IPrl*Itj rROStON-1 CORROSION INSPECTION PROGRAM FEEDWATIER LIW2 4*-FDW-S TUR5'NE BVULNG-FEED PUMP ROOM/HEATER BAY REFERENCES, Gi 15"7,219G)13!_,C41183 19 5950-F5-424,5920-F 5*-i25 COMPON-ENT LOCATION SKETCH "40.01 8

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trEATe 0A.j~ ) COeCNSER A REVISION 1111/25/91 VERMONT YANKEE PIPING EROSi0N-CORROSION INSPECTION PROGRAM TURGNE BEtONG-FEED PUMP ROOM/HýATER BAY FEEDWATER LINE 4'--DW-6 REFSRNCES* GI9J 157,G)9191MZ191 193 59S0-PS-124,5920-FS-125 COMPONENT LOCATION SKETCH No.019 AppendixA PP7028Original Pagc 24of 102

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"-D FEEDWATER LNEI B'-"Dw-B TURBNE BUILDNG-FEED PUMP ROOM REFERENCES G191157,G191 102,G19118,9o2O-F5-:24 COmPONtENTr LOCATION SKETCH NoaO1 I

-.1 -I Appendi, A PP 7028 UhOgnmJ Pzsz 16 of 102

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j 5-PD 2tLO7 REVISION I: 11/24/91

'Et NOZZLE. VERMONT YANKEE PIPING EROSION-I-EATBR E--JA FDI1SP2O CORROSION INSPECTION PROGRAM

-.- /_ FD 2*a~c EL 251'-4' REF PO 12W.9 FEEOWATER LINE IS' FZW-12 TURBNt4. BUJLDJNG-HEATER SAY -4 REFERENCES C91 157,O191 182,GI9'15S.s592D-FS-i25 COMPONENT LOCATION SKCETcH No.007 Appendix A PP 702-8 Original Pgc 12 of OZ NEC037127

EL 2411-G' (FLANGED SPOOL PE*CI FEFDWATI' PUMP 1/2

--

VERMONT YANKEE APING EROSION-CORROSION INSPECTION PROGRAM CONDENSATh UNE 2tr-C-30 (CONTINUED)

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Apptn~dixA PP7028OzinnzI Pa?&p42falf 2 NEC037128

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- TURBINE Ut(D1NG-F'i)WATER PUMP ROOM R-EFERENC"ESi G191l 157CCW*

I150 ,0 191 187,5920-,FS-3. IS COMPONENT LOCATION S<ETCH No. 03E 4 4 A~ppt:nd'x A PP 7028 Oyigý,nm Pagc 43 of 102 NEC037129

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%ObC -32 CORROSION INSPECTION PROGRAM MAIN STEAM UNE I8"-.M5-U1 & Y"-M$-7D A QRYWELL & STEAM TUNNEL EF ERENCE$i G 191167,G 1911862,5920-F5-lP COMPONENT LOCATION SKCETCH No.080 UnOinaJ Ftigefl85otiO

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NEU03/135 17 of 0 . . 10

VERMONT YANKEE 4-SCOPE MANAGEMENT REV[BW FORM Date: Tracking Nwnber: "_ I (Assigned by Work Scope Control Coordinator)

Work Order Number: 04- " __ kdcrTn.eccrejnni C r Initiator: AKiZowf 6Xtbthrt3cA T_

T_

I. Dept. MgT. I Location of Work to be Perfonned:

ADDITIONQ DELETION ° GEIANGEf Description Justification for Request I QA&S14 aci~ :g~4i~a&c m~

Review Process Additional Cost:

Dillation and Schedufling Impact:

Assigned DeptAqflan-t-lours to Complete:

Source of Manpower/Other Scope Impacted:

Dose, Chemistry, Safety Implication:

Engineenng Impact- Man-Hours/Engineering Dept.

Optional Ways to Address:

Approval Process justification tlease provide a brief Scope Review Committee RecommendatioW/Plannming Priority: ,,1 Priority "C"WO Responsible Depl Approval CGenral Manager, ____-_-___.-____"-

Plant Operations:

EMPAC Change Made for ent_

4eD&Pririt Coprocýisprv oe & Priori s App roy isapprove Date:

Date d

Log Updated: Q,,0 w--schdu hi -"

Go, ig;4,t,W ork Contro,*

VYPPF 7102.01 PP 7102 Rev. 2 Pagelofl NEC037136

Prepared By: James Fitzpatrick Date: 11/1/05 RFO 25 FAG Program inspections location nos. 2005-25 through 2005.35 References Work Order 04-004983-000, FAC Inspections Work Order 04-004983-010, Surface Preparation on SSH piping TM 04-031 Work Order 04.004884-006 ER-05-0190 CR-VTY-04-2985 CA3

Background:

CR-VTY-2004-02925 doouments a steam/water leak on the turbine steam seal piping, line 1 SSH4 to the No.4 packing. TM 2004-031 installed a temporary leak enclosure on this line.

Inspections on Turbine Steam Seal Piping were included in the scope of the FAG program for RFO 25 per CA3 of CA-vrY-2004-02925. The purpose of these inspections is to determine the extent of condition on the remaining steam seal piping.

Work pe These inspections require access to the SSH & SPE piping on elevation 272 of the Turbine Building. The piping is located under the IP turbine appearance lagging deck plates and requires removal of section of the plates to access the piping for surface preparation and inspection. It was intended that these inspections be performed along with restoration of Temp Mod 2004-031 (W.O.

2004-4884-006).

Discussion Restoration of TM 2004.031 was removed from the outage soope on 10/24/05 due 10 interference with critical path work planned on the LP lurbines. A detailed ralionale for delaying resloration of the TM from RF025 was deveioped by George Benedict on 9/98105 and is attaohed here. The same reasoning and technioal basis applies to these Inspections.

In addition these inspeotions are not programmatically required under PP 7028 (Piping FAG Inspection Program). The inspeotions were added to the RFO 25 scope to determine the condition of the piping at parallel and similar locations on the Steam Seal piping as the 2004 through wall leak.

The system is a low pressure system with piping Iooated in the heater bay or under the turbine deck plating. Deferral ot these inspections does not. pose a significant personal safety hazard as exposure to these lines during operation is minimal. The possibility of a leak at another location on the Steam Seai piping still exists_ However, the low operating pressures and the results of UT measuremenls made on Ihe 1SSH41ine at the location of the existing leak indicate that any failure would be a pinhole type leak vs, a catastrophic failure of the pipe.

NEC037137

I Aft Prepared By: G.13enedict

-2Ttergy l3ate: 9/28105 I

Replacement ofN4 Steam Supply Piping I Work Order 04-48S4-06 TM 2004-031 ER 05-0190 I

The steam seal supply line to TB4I-JA, N4 packing developed a leak from what appears to be I the result o.fpipe erosion on one ofthe pipe radiuses. Team Inc. was contacted 10 develop on-line repair options and determined that,the most appropriale long tenn repai.r would be to instill! a pre-fabricated clamping device. The cIamp was fabricated as recommended. and successfully installed per the above referenced Temporary Modification (TM 2004-03!).

The permanent repair for the N4 steamý seal supply line is currently scheduled to be implemented during RFO 25. The pipe clamp and the degraded section ofpipe will be removed and new piping will be field fit and installed. To facilitate this work, it will be necessary to re.move sections ofthe LP turbine appearance lagging deck plates to gain access to the piping. Use ofthe overhead crane will also be required to remOVe/install piping and dcck plates.

ItM -MmSal EiAvsp nzýdn

?x During RFO 25 a significant amount ofwork will be pcrfonned on thc LP turbines which are located in the immediate area of the degraded N4 steam seal supply line. The LP turbines will be completely dismantled to facilitate the installation of the new 8 stage diaphragms and to perform the required tcn year inspection. The location of the degraded steam seal line is directly betweevn both LP turbines and implementing the LP inspection in conjunction with the steam scal line repair will create personnel safety hazards, potential equipment damage, and logistical NomplicaCions.1 I

I NEC037138I

Prepared By; G. Benedict

- EDate: 9128105 The following represents the specific issues that will be present during the implementation of the N4 steam seal line replacement and the LP turbine inspection:

PeTsonnel Safety:

> Fall and drop halzards will be created by both work crews in proximity to both work areas. Open holes will exist on. the turbine deck appearance lagging deck plates and in the area between the LP inner casings and exhaust hoods. Although, personnel protection barriers and equipment will be utilized to mitigate fall and drop hazards, personnel awareness, focus, and goal will be on each individuals own task. The drop and fall hazards will be continually changing as each work activity progresses and although peisOnnel are required to communicate changes to safety hazards these types ofcbanges will be extremely difficult to manage due to the pace of the LP turbine inspection activity,

> The crew working on the stemn seal piping will continually be interrupted due to overhead hazards from materials being removed and returned to the LP turbine centerline. Once again dne to the pace of the LP Iturbine inspection and the fact that the steam seal piping replacement crew will be in and out ofthe work area which is not visible from the turbine floor only increases the potential to inadvertently transfer a load over the piping replacement crew.

Equipment Safety and Quality:

The removal and installation of the steam seal piping will involve welding and grinding activities. Shielding can and must be installed to prevent inadvertent weld flash, slag, and grinding dust, however, performing these types ofactivities in the vicinity ofopen bearing oil sumps, exposed shaft journals, and bearing babbitt surfaces increases the risk for accidental damage.

Schedule and Logistics The LP turbine work is the primary critical path activity for the Outage and any delays encouutered by the implementation ofthe N4 steam seal supply line repair will most likely result in an increase in duration. The repair of the steam seal line will require a moderate use of the turbine building crane to remove/install deck plates, piping, and appearance lagging. In addition, crane support will be required to remove damaged pipe...install and fit-up new pipe sections.. ,remove new section to perform non-field welds.,.and permanent installation, TIlere is zero turbine building crane availability during RFO 25, The open hole caused by the removal ofdeck plating will cause the "A" LP to be logistically separated from the "B" LP on the right side of the centerline which NEC037139

I aai Wer ('ri-1 Prepared 'By: G. Benedicl Dateý: 9/28/05 I I

,will create a delay in the txansfer oftooling and materials between LP "A" and I

> Asbestos coneem: There is a potential that the steam seal line being repaired contains asbestos insulation. Any asbestos insulation issues could shutdown work on the turbine deck. U

> Maintenance resources: Mamtellallce crewS assigned to the steam seal line repair have 7 shifts available to perform this repair- lfthere are any delays in perbrrmng the repair (e.g. coordination issues or emergent issues during the I

work), the maintenance crew would be required to leave the steam seal pipe repair and return to the refuel floor. U I

Team Inc. was contacted (o determine the feasibility ofoperating the unit fol' an additional cycle with the Team clamp in place, The response from Team IIIC. was very faverable with regard to operating an additional cycle with the clamp in place. According to Jim Savoy (Team Inc.

I District Manager) many commercial industrial facilities that have utilized clamps similar to the one installed on the N4 steam seal supply line have operated for extended periods much greater than the requested 18 months.

I The steam seal supply is approximately 2 - 5 lbs. of pressure with a maximum temperature of 255 degrees F. This is cotsidered very low in comparison to many ofthe applications that Team I Inc. has installed similar long term clamps on. Ifthe clamp is left installed for an additional operating cycle there is a risk that the clamp will teak once the plant is placed back online.

Although considered a low probability, the risk is due to the thermal cycling ofdissimilar I materials that are utilized in the clamping and sealing process. Ifa leak were to occurTeam Inc.

would re-inject the clamp with sealant which has been successfully perfOlmed at othcr locations.

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NEC037140 I

VERMONT YANKEE SCOPE MANAGEMENT REVIEW FORM Date: / 4)tc*ý Tracking Number:

(Assigned by Work Scope Control Coordinator)

\Vork Order Number: olV 4 -ac Reference Documen t M .ooAc-3t

- 1 (ER, MM, TI, 0028, etc,)

Initiator: IWo to ef ie t re Approved By:

Dept. 77gr.

L.ocation of Work to be Per-formed:'~ 3~c

-7 AA~T ION 0 DELETION 2 CBANGE t]

Description Justification fo, Request

$ge~ ~ ~~b aca mZe'. tA A'tke~t >a6 42a? tA jao-4 J1 Review Process Additional Co,t:__._.

Duration and Scheduling Impact:

Assigned Dept./Man-Hours to Complete: ------

Soumrce ofManpowerlOtler Scope Impacted:

Dose-, Chemistry, Safely implication:.

ngineening-mpac Tan-Hoursi..ngine..ng Dept. _=

Optional Ways to Address:

Approval Process Please prgrid. a b etss*ifl,*tion Scope Review Committee RecomnrnendationfPlanning Priority-,

Priority "C" ,,. nsib. Dept Approval__--.

Pi~ntManaj ~ JA ~ PA prv Disapprove- Date:.I9 2,4q- q EMPAC Cw-- or Even( Co{e Pfi riy ....

or~vertC-~Pri rity_____ 8CC Date Log Updated: -

CopieA to Work Control, Outage Scheduling.

VYPPF7I02.01 pp 7102 Rev. I Pagelofl LKPC ,15 NEC037141

i outa.RFO-25 Piping FAC Inspections Outage Scope Challenge Meeting 5/4/05 ..-.

Short or Cryrptic summary of what the roiect involves and why we need to complete the proeec' in RFO 25 e_., regulatory requirement. dsk to generation. program requirement. appropriate management of the assetI In response to USNRC Generic letter89-08, inspections of piping components susceptible to damage from Flow Accelerated Corrosion (FAC) are performed each refueling outage.

The planning, inspection, and evaluation activities are currently defined in program I

procedure PP 7028, "Piping Flow Accelerated Corrosion Inspection Program". Before the start of RF025, VY will transition to a new Entergy procedure "Flow Accelerated Corrosion Program", ENN-DC-315.

Description of the s of the ioec.L what it encompasses. o that have been considered

{identiy minima required m discretionary could be deferred s Otheo_.ag*., o that interlaces'with or can be included in this prole.tu Impact on others.

The scope of the inspections for each refueling outage is based on previous inspection results, predictive modeling, industry and plant operating experience, postulated power uprate effects, and engineering judgment. The scope for the Fall 2005 RFO is defined in Design Engineering-MIS Memo VYM 2004/007, Revision 1. The 2005 RFO Scope includes:

I External Ultrasonic Thickness (UT) Inspection of 37 large bore components at 16 locations.

Includes:

I

• 5 components recommended for repeat inspections based on prior UT data

% 2 components for CHECWORKS model calibration

• 6 components based on Operating Experience (Mihama Event)

I

% 6 components downstream of leaking N.C. valves (identified from TPM)

• 4 components based on increased EPU flows

% 2 components D.S of FCV -104-4 (suspected cavitation) i

, 12 components based on current through wall leak in SSH at LP turbines External Ultrasonic Thickness (UT) Inspection of 5 sections of small bore piping based on industry experience. Includes 4 sections of piping downstream of restriction orifices at the CRD pumps. i Internal Visual Inspection of two 36 inch CAR lines to assess changes in flows from HP turbine modifications installed in RFO 24. Internal Visual inspection of the only remaining carbon steel 30 inch diameter line 30"-8.

Pre-outage scoe and lg Lead time parts/contract that have been identified, None 3 Page 1 of3 I NEC037142

RFO-25 Piping FAC Inspections Outage Scope Challenge Meeting 5/4/05 Initiatives. creative opportunities, unique prnblems associated with the project.

None The inspecUon process used is the industry standard. Removal of insulation and surface preparation are required for the UT equipment. Remote methods which do not require insulation removal are still in the development stage, and do not currently have the accuracy required to trend low wear rates (EPRI CHUG). Phosphor Plate Radiography which is currently being adopted to screen small bore components without insulation removal is primarily applicable to PWR plants. limited use on BWRs, Design Engineering - MIS has minimized the number of Inspections performed each RFO.

VY has traditionally trended well below industry average number of components inspected each RFO. This is primarily due the original design of the plant and replacements with Chrome-Moly piping_ Recent trends in numbers of components inspected at other plants show reduced numbers of inspections based on piping replacements.

Identify additional organizational support requied, and specifically, management support 1_ces ar-V.

Inspections will be performed by the ISI personnel. Scheduling and staffing will be coordinated with other ISI activities. Inspections are performed using approved NDE procedures. Training on inspection procedures is performed under the ISI program, Grid marking per new ENN Standard ENN-EP-S-005 Primary DE-MIS interface is the ISI level ill and/or ISI Program Engineer for coordination in review and approval of inspection data. Interface with craft & other plant groups is normally through establIshed links in the ISI program. Unusual situations which require additional support will be raised to management level as required, Two DE-MIS engineers (J.Fitzpatrick & T.O'Connor) currently trained in evaluation procedures and have prior VY FAC Program Experience. Other DE-M/S engineers with pipe stress experience can be trained on shari notice. The number of inspections Is slightly higher than the last two outages, Coverage will be provided 7 days a week (or as required) to evaluate UT data.

The FAC Program Coordinator (J,Fitzpatrick) is responsible to insure that inspections are performed and the data is evaluated in accordance with the program requirements. Activities will be coordinated with the 151 coordinator (Dave King), Any problems that arise that can not be handled at the engineer level, will be elevated per outage management guidelines (30 minute rule, etc.),

Page 2 of3 NEC037143

. /,: I I

/"

RFO.25 Piping FAC Inspections Outage Scope Challenge Meeting 5/4/05 ldeii*f aDy preparation issues necessary Q meet upcoming otae_ milestones, I

" Coordination with LP Turbine work for inspection of SSH components (physical space)

I

" Coordination with LIP Turbine/Condenser work for ventilation path (opening) for the 30" B Cross Around Line and for a window to perform inspections (noise issue).

I

" ER for Design Engineering - Fluid Systems to develop a (paper) Design Change to reduce the piping design pressure in the Feedwater Pump Bypass Lines at the condenser. Current design pressure for the piping attached directly to the condenser is I 1900 PSI. Local sections of carbon stee! piping remain at the condenser. Leaking valves during past operation cycles may have resulted in increased wear in carbon steel section of line. I Identify if all necessar outage and re-outaae WO's for the proiecUprogram scope are generated. I Work Orders to for support activities and inspections (04-4983-000 series)

,1entiv if a J _4 opportunities to perfonn any Part of this scope could be completed pre-outage?

I The only components which are not high temperature and are in an accessible location I

during plant operation are 4 sections of smali bore piping downstream of restriction orifices at the CRD pumps. These may be inspected during operation. However, this is a high noise area.

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Page 3 of 3 I NEC037144 I

Engineering Standard Review & Approval Form

• ..... Engineering Standard Change Class]fication Tempora r N~qj evisedt E Cancel S Editorialt I CC Flow Accelerated Corrosion Component Scanning and ENN-EP-S-005 N/A Grkddinc Standard I nctplonaA7I,ý ur? 11 -nI11 Manua uth r Engineering Programs Jeffery Goldstein Ian Mew 6Si -

hnducn Reviews I AO 0 n

ECH lAP 0 GONS PN PS 0 WF3 WPO 0

6 Review T Yes No Date Technical Review (See Not. belom for DesIg_ Chnrge _Standards) James C F_-_z Independelli Design Verification (See Note below for Oesirmn Change Standards)

I0CFRS3059IProeess

{ . *-zte c¢ n inq Dn Applicability Review d, <.výu alio n d joc um nents zl) ED 0 J a r e s 0, F ýIz p a td , ".

Sa. Nole t'rlow ior Deii] n C.har c Sea-dardel Note: ReW ' o a/aChangeStan~dards ara CVoaaoented w/th/n IAe 91 Nu6 appiicabieER

• An ER Number Is Oa ured for jasi Chann e S(8,ed rds, on ,

Cross Discipline Reviews I O1, 1 !geiwr__e___t_ Reviewer Name L[Sfqnaturt e

I Ia

,*iK*WJ Changde {TCN Approval I Name,. [Signalure: IGl I Comments Section mments Made .- Cornmer~ts' trkpPd I flt... - --

I I .1 IL.1 L .I* dII 13 M11UIUl*U VtM CommentsiTCH Chalwe This standard replaces VY specific 'Component Gridding Guidelines" previously contained in Appendix A o0 VY NDE procedure NEtS,05.9. NE-80543 has bedn superseded.by ENN-NDE-9.05 All VY comments were resolved during development of Ehis standard.

Lb.

- ---------

NEC037145

/ 9 ENTERGY ENN ENGINEERING STANDARD ENN-Ep.$-005 Rev. 0 EffecOve Dale: JAFIWPO.9/1/04 PII- 61110$

IPEC.1 01.1104 Flow Ac£,celerated coQrfs~jln CO Ig - and Griddkog Standard Applicable Site(s):

11,10 11'2 [] 11'3 [ JAF L PNPS 9 V Safety Related: Yes x No Prepared by:

Approved by: ___ 6~ ~ 514 Dale:gif(

Engineering Gui~e Owrngr NEC037146

'Th(S Engineering Standard Review & Approval Form YM4id 16pil-

-

0 n- Ineer[n Standard Chan e Classification New Iz Revised I 0 CancenPl I 0 Editorial -L 0 Temporary I 0 (TCNI En Ineerin Standard Title Doc, 0. RevNo. TeN No.

Pipe Wall Thinning Structural Evaluation Functional Diec lite En lresri Standard Owner En ineerln andard Preparer lR. Penny H. Y. Chang Site Cond~uctin Reviews JAF I PNP$$tW H Review T . Yes. No RevlewerNa elSi ore Date Techinical Revielw0 (See Note below for Peslgnr Change Standards)

• 0 James C._Fitzpa.rick Indopenden, Design Verificatlon I15 0 flzptro" --

(See Note below ic Design Change Standarde James .t 100RSOS~lrocssAppFicabWiity R-eview (attach screening and evaluation doroumen, 2t 0l-,atic k,

.,e Note telow for Desr Chr~e Standards) James C.

NcAea Ior<eslnR Change Staadsrds rarOocunnted wdhin the

.eviews am)-fiable ER. RNunbe-r e

. An Cl rvxww- s 1w t ,e ¢rnanj p&a ardos jIC Cross Discipline ReviewsI - IK Site Enaineerina Standard Charnnlon Scott D. Goodwin Editorial Change ITCN Approval IName: ISignature: _ JDate~ ____

I I

{C~UI~fWUflFT Comments Sectlon l VAt~ l II I TCN Effective(Expiratlon Date I CommentsfrCN Change All VY comments resolved during development of this standard.

NEC037147

Pagelofl Fitzpatrick, Jim "Tom: Fitzpatrick, Jim

,ent' Tuesday, September 27, 2005 11:45 AM To: VTYEngineering-Mechanical Structural; VTY EFINDL

Subject:

FW: Communication of Approved Engineering Standard I

This is a new fleet standard for evaluation ot thinned wall piping components which will replace ENN-DC-133 ENN-I)C-133 will be superseded, VY Department Procedure DP ENN-DC-315 0072, "Structural Evaluation of Thinned Wall Piping Components will be revised or I

superseded as required when is adopted.

Entiy Conditions for this Standard will be in ENN -DC-315 "Flow Accelerated Corrosion Program" and ENN-OC-185 "Through wall leaks in ASME Section X) Class 3 Moderate Energy Piping Systems". WPO has the responsibility to revise the references to ENN-DC-133 in these procedures. 3 Qualifications(frainillq; At present there is no ENN QUAL CARD for use of this Engineering Standard, Calculations performed using standard are documented per ENN-004-26. Based on the scope of this standard, only Design Engineering- Civil! Structural personnel and the Mechanical types in EFIN with previous pipe stress experience have the charter and background to apply this standard.

Summary ot Changes from ENN-OC-133 as applicable to VY:

More tormalized ties to ENN-OC-315, Wear rate determination for FAC program inspections is the responsibility oj the FAC Program Engineer

• Calculation of componellt Wear, Wear Rate and Predicted Thickness is consistent the same as OP0072. The only change Jrom OPDOn is a reduction on the Safety Factor (SF) from i .2 to 1.1.

• The methods used to calculate the code required thickness for pressure and moment loads are consistent with OPO072, but presented in a differentformat.

• No significant changes to application oJ ASME Code Case N-513 for though wall leaks.

Added attachment for guidance in oatculation of component wear rates.

Excel spreacdaheel templates are available to facilitate calculations.

I From: Ettlinger, Alan I sent: Monday, September 26, 2005 9:33 AM To: Casella, Richard; Fitzpatrick, Jim; LO, Kai; Pace, Raymond Cc: Unsal, Ahmet

Subject:

Communication of Approved Engineering Standard In accordance with EN-DC-146, as the Site Procedure Champion (SPC) at your site, please inJorm and communicate to applicable site personnel, the issuance of the following fleet NMM Engineering Standard. I ENN-CS-S-008, revision 0 Pipe Wall Thinning Structural Evaluation ThiS standard supersedes ENN.DC-133. The standard can be accessed in IDEAS on the Citrix server.

The standard becomes effective, and will be posted on September 28, 2005.

Ifyou have any questions, please give me a call, I

10122f2005 NEC037148

,NEC-UW_21 Second victim dies of burns from power plant explosion Milwaukee Sentinel, Mar 9, 1995 by BETSY THATCHER

• E-mail

" Print

" Link A second victim of the Feb. 12 steam explosion at Wisconsin Electric Power Co.'s Pleasant Prairie plant died Tuesday.

, Ad Feedback WEPCO employee Gregory A. Schultz, of Waterford, died at St. Mary's Hospital in Milwaukee, where the 37-year-old operating supervisor was being treated for severe bums after a steam pipe ruptured at the Kenosha County power plant.

Schultz and another operating supervisor, Steven Baker, were performing a routine inspection of the piant when a 12-inch pipe that carries hot, pressurized water into the boiler of Unit 1 ruptured. Baker, 38, of Kenosha, died at the plant.

Schultz, who had worked for the company since 1978, received second- and third-degree burns over 60% of his body.

Related Results "Words can't express the sorrow and regret we feel," WEPCO President Richard Grigg said in a statement. "We are remembering the Schultz and Baker families in our thoughts and prayers."

An investigation into the cause of the rupture is expected to be completed by the end of the week, company spokesmen said.

Preliminary results indicate there was substantial thinning of the pipe wall, which resulted in a break, a company statement said.

Employees of the Pleasant Prairie plant plan to buy a granite marker to place near a flagpole outside the plant in memory of the men.

Copyright 1995 Provided by ProQuest Information and Learning Company. All rights Reserved.

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NE-C-UW_22 Index I Site Map JFAQ I Facility Info I Reading Rm I New I Help I Glossary I Contact Us  : GooiC7. =S Abu-R uclearý R-eactors;; Nuclear Mat reiint!W MaeralWst Radioactive t :Nuclear:

Secursit Publ'6icMeetin~s I Home > Electronic Reading Room > Document Collections > NUREG-Series Publications > Staff Reports > NUREG-0933 > ISSUE 139:

THINNING OF CARBON STEEL PIPING IN LWRs (REV. 1)

I ISSUE 139: THINNING OF CARBON STEEL PIPING IN LWRs (REV. 1)

DESCRIPTION I Historical Background This issue was raised 1 0 8 9 as a result of a pipe rupture in the main feedwater (MFW) system at the Surry Unit 2 nuclear power plant on December 9, 1986. The MFW pipe rupture followed a reactor trip from full power shortly after the unit returned to operation on December 8, 1986, following a scheduled refueling outage. The staff presented briefings on the incident to the Commission on February 25, 1987, and to the ACRS at its 322nd Meeting on February 5, 1987.

The Surry pipe rupture was in the 18-inch "A" MFW pump suction line immediately downstream of a compound 90 elbow and T-section connecting the 18-inch pipe to the 24-inch condensate header. The rupture was a catastrophic, 360 circumferential break. A piece of the ruptured pipe (approximately 4 feet by 2 feet in size) was blown some distance from the break point.

i The piping still attached to the pump suction rotated away from the break point and came to rest against a portion of the "B" MFW pump discharge piping. No significant damage to the "B" MFW pump was noted.

The failed 18-inch suction line was fabricated from ASTM A-106 Grade B carbon steel and ASTM A-234 Grade WPB carbon steel wrought fittings with a nominal wall thickness of 0.5 inches. Visual inspections of the inside surface of the elbow revealed a dimpled surface and general pipe wall thinness as small as 0.05 inches. Ultrasonic thickness measurements indicated the wall-thinning to be a gradual change over most of the elbow fitting. The licensee concluded that the pipe ruptured because of the thinned wall and that the thinning was a result of erosion/corrosion.

I On January 15, 1987, the Honorable Edward Markey (U.S. House of Representatives) requested the GAO to assess NRC actions following the Surry event and several other technical problems at nuclear power plants. The GAO assessment 1 0 9 0 of actions taken related to the Surry event and similar piping deteriorations detected at other LWRs was issued in March 1988.

The major GAO conclusions and recommendations are provided in the conclusion of this analysis.

A similar pipe rupture occurred at the Trojan plant following a reactor/turbine trip on March 9, 1985 (See LER 85002, Docket No. 5000344). The pipe rupture at the Trojan plant was in the 14-inch heater drain pump discharge line immediately downstream of a globe valve leading to the condensate header and MFW suction side. The piping was the same ASTM A-106 Grade B material with a required minimum wall thickness of 0.375 inches. The wall thickness in the region of the rupture was thinned to approximately 0.1 inches and the cause was attributed to wall-thinning by erosion/corrosion.

In both events, the fluid medium was single-phase, subcooled water at nominally 350F and 450 psi. Water velocities were in the range of 20 to 40 fps and the flow in the ruptured locations was subject to turbulence induced by piping and fitting configurations, with pressure increases resulting from automatic MFW isolation.

Historically, erosion/corrosion in nuclear and fossil plants has occurred primarily in wet steam (two-phase) lines and has not I been reported in dry steam lines (EPRI NP-5410). 1 0 9 2 The erosion/corrosion in single-phase (water) systems was not expected and differs in the mechanisms contributing to the process, being a complex phenomenon dependent on many variables such as alloy content, temperature, Ph, and flow velocities and perturbations caused by piping and fitting i configurations.

Following the Surry event, the staff issued a series of Information Notices informing the industry of the Surry pipe rupture.

On July 9, 1987, the staff issued NRC Bulletin No. 87-011093 requesting licensees to submit information concerning their I programs for monitoring the thickness of pipe walls in high-energy, single- and two-phase, carbon steel piping systems.

Staff review of the licensees' responses to Bulletin 87-011093 were reported in SECY-88-50 1 0 9 4 and Information Notice No.

88-17.1095 A staff report on the status of the industry erosion/corrosion program was provided in SECY-88-50A. 1 0 9 6 For two-phase, high-energy, carbon steel piping systems, responses indicated that licensees had programs at all plants for I of 7 4/28/2008 1:11 AM

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inspecting pipe wall-thinning. However, because the guidelines were not required to be implemented, the scope and extent of the programs varied significantly from plant to plant.

For single-phase piping systems such as in the feedwater/condensate lines, a limited number of inspections were conducted following the Surry event. Based on the Bulletin 10 9 3 responses up to the time this issue was evaluated in November 1988, 23 out of a total of 110 units had not established an inspection program for the single-phase lines. Of these units, 17 were operating plants and 6 were under construction.

I The staff review1 0 9 1 showed that wall-thinning in the feedwater/condensate systems was more prevalent in PWRs than in BWRs. The review indicated that licensees of 27 PWRs and 6 BWRs identified various degrees of wall-thinning in feedwater I piping and fittings. The pipe wall-thinning problem was widespread for single- and two-phase, high-energy, carbon steel piping systems in PWR and BWR plants. Since the problem was more prevalent in PWRs, this analysis focused on PWR plants. However, due to the nature of the problem, the resolution indicated that the issue related to all LWRs.

Safety Significance There were no requirements for the industry to have an inspection program for monitoring and examining the ASME minimum wall thickness for carbon steel piping. Therefore, even though a pipe break is a design basis event for which plants i

are designed, the potential frequency of such breaks was higher than previously anticipated. Lacking inspection requirements to provide assurance of the defense-in-depth against catastrophic pipe ruptures in the secondary power conversion systems (and specially the feedwater/condensate systems), plants may not have adequate assurance that they meet the design basis life.

The higher pipe rupture frequencies could also introduce additional challenges to safe plant shutdown from potential systems interactions of the high-energy steam/water releases that may damage, or affect, other systems (see "Systems Interactions from Pipe Ruptures" below). Thus, risks from design basis pipe ruptures that did not account for erosion/corrosion wall-thinning in the secondary piping systems may be greater than previously evaluated.

Possible Solution The staff was continuing its review of pipe wall-thinning and was expected to assess the results obtained from inspections to be performed during the 1988 Spring refueling outages. 1 0 9 4 1 0 9 6 This assessment included visiting up to ten plants to review their inspection methods and results. The staff anticipated that its review would be completed by December 1988 and I

could, if necessary, provide the basis for new requirements1094,1096 in single- and two-phase carbon steel piping systems.

A possible solution for the single-phase piping systems, which unlike the two-phase systems that have existing monitoring programs, mnight include inspections to be conducted at each refueling outage. However, for the long term solution, the staff planned to continue working with NUMARC and EPRI to arrive at an implementation program and schedule for the resolution of pipe wall-thinning in both single- and two-phase carbon steel piping systems.

PRIORITY DETERMINATION Pipe ruptures from erosion/corrosion-induced wall-thinning of carbon steel piping had not been reported prevalent in dry-steam lines 1 0 9 2 such as the main steam lines. Two-phase piping lines, such as the turbine crossover/under piping and steam extraction lines, had experienced erosion/corrosion wall-thinning and ruptures even though licensees had monitoring and inspection methods (though not required) in place to various degrees for some time. This indicated that improvements were needed in the existing inspection programs to provide timely detection of the piping degradations. I Single-phase carbon steel piping runs, which were not believed to be susceptible to erosion/corrosion wall-thinning, were not in general (prior to the Surry event) monitored or inspected for potential wall-thinning. The single-phase systems in the secondary power conversion systems which had been found to be susceptible to wall-thinning were the feedwater/condensate systems and the high pressure feedwater heater drain pump discharge piping lines. These I

single-phase lines transport water at a nominal temperature of 350 0 F and water velocities ranging from 20 to 40 fps. Both of these conditions tend to exacerbate the erosion/corrosion phenomenon in carbon steel piping systems carrying single-phase fluid I

(water).

AFW piping lines that typically draw water at lower temperatures from the condensate storage tank, and do not experience continuous flow during power production, had not been reported to be susceptible to erosion/corrosion wall-thinning.

1 Because it was difficult to determine the effectiveness of the two-phase piping systems inspections, lacking information on previous repairs and replacements resulting from the inspections, the two-phase rupture frequency was assumed equivalent to the single-phase carbon steel piping rupture frequency estimated below. Without existing inspections, the two-phase 2 of 7 4/28/2008 1:11 AM

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piping systems would be expected to have a higher rupture frequency.

As stated above, this analysis focused on evaluating the carbon steel wall-thinning pipe ruptures in single-phase piping systems and the wall-thinning ruptures in two-phase piping systems of PWR power conversion systems. Based on existing inspection results, BWRs appeared to have a similar problem, but to a lesser degree. Therefore, this analysis bounded the issue for all LWRs.

Recovery of Power Conversion Systems I The power conversion systems feed into one another through various piping configurations, including straight lines or headers and various valving or fitting arrangements. Therefore, a rupture in either the single- or two-phase piping systems could disable the PWR power conversion systems to various degrees. Thus, the probability of recovering the power I conversion systems was uncertain. Therefore, it was conservatively estimated that the probability of non-recovery of the power conversion systems (PCSNR) was 0.5, given a rupture in the secondary systems.

Carbon Steel Pipe Rupture Frequency I he data on erosion/corrosion-induced wall-thinning resulting in ruptures of carbon steel piping carrying single-phase fluid was limited to the Surry and Trojan events described earlier. This limited data was used to estimate upper and lower bounds of the subject pipe rupture frequency.

I For the upper bound estimate, the plant-specific experiences of Surry and Trojan were used. At the Trojan plant, the pipe rupture occurred after approximately 9 years of operation. At the Surry plant, the pipe rupture occurred after approximately 14 years of operation. This data yielded an upper bound rupture frequency of 9 x 10- 2 /RY. For the lower bound estimate, I the two pipe ruptures were ratioed over the total number of PWR reactor-years of operation (approximately 600 RY). This yielded a lower bound estimate of 3.3 x 10- 3 /RY.

The rupture frequency was approximated by a log normal distribution with an error factor of five and the upper and lower I bounds were assumed as two symmetrically located percentiles (0.05 to 0.95) of a log normal distribution. The calculated mean rupture frequency was 3 x 10-2/RY. As stated earlier, it was assumed that the rupture frequency of 3 x 10-2/RY was applicableto the secondary side carbon steel piping systems identified herein.

I Most of the pipe ruptures that might occur in the non-safety-related portions of the secondary systems are likely to be outside of containment because most (90%) of the secondary side piping is located outside containment. Pipe ruptures in the safety-related portion of the MFW piping inside containment can result in the secondary side of the affected steam i generator blowing down to the containment atmosphere. For these lower frequency ruptures, (0.1)(3 x 10-2) = 3 x 10- 3 /RY, isolation of AFW to the affected steam generator will reduce the chance of containment overpressurization from continued long-term steaming due to decay heat from the reactor core. Automatic AFW isolation is necessary to ensure that the containment design pressure will not be exceeded. This event, like other ruptures that may occur in the PWR power I conversion systems, was treated as a total loss of main feedwater. This sequence was bounded by the TMLU rupture event sequence described below. However, pipe ruptures inside containment are less likely and will not likely induce the negative systems interaction problems that can result from pipe ruptures outside containment.

i Systems Interactions from Pipe Ruptures Communication Systems Failures: During the MFW pipe rupture at the Surry plant, the Cardox and Halon fire suppression systems were actuated by steam/water intrusion into their control panels. The security repeater which was located approximately five feet from a Cardox discharge nozzle failed and was later found to be covered with a thick layer of ice. As a result, security communications were temporarily limited to the non-repeater hand-held radios. Therefore, actuation of the Surry fire protection system (FPS) resulted in loss of a train of the communication systems.

Given that loss of one train of plant communications occurred in one of the two pipe rupture events, the probability that failure of this train of communication can occur as a result of pipe ruptures in the secondary systems outside containment was estimated to be 0.5.

To estimate the probability of loss of the backup hand-held communication radios, the following.were assumed: probability of battery failure = 0.1; probability of operator error in not replacing the batteries = 0.1; and probability that other units are not readily available = 0.1. The probability of loss of both communication systems, given a pipe rupture in the secondary systems outside containment, was estimated to be 5 x 10-4.

To estimate the impact of the loss of plant communication systems, it was assumed that loss of communications would increase operator errors in the four event sequences affected by the pipe rupture. Based on an examination of the fault trees 5 4 for the four sequences, and adjusting the operator errors to account for loss of communications, the percentage increase in core-melt frequency for each sequence was estimated as follows:

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Loss of Communications Sequence  % Increase In Sequence Core-Melt Frequency TMQH 7 TMKU negligible TML(PCSNR)U 7 TMQD 2 Actuation of FPS: Within minutes of the MFW pipe rupture at Surry, 62 sprinkler heads opened in the immediate area of the rupture. As a result of the sprinkler water and the feedwater discharge, the Cardox and Halon suppression systems control panels were affected by intrusion of steam/water. The intrusion caused the time limit, battery charger, and the dual zone modules to short. Thus, the manual remote actuation circuit located in the control room was affected. I In Issue 57, the effects of actuation of the FPS actuation and the potential increases to core-melt frequency were estimated; the sequence evaluated was the TMLU sequence and the safety system evaluated was the AFW system. Because one of the two pipe rupture events (Surry and Trojan) affected the FPS manual remote control, the estimates in Issue 57 were adjusted by assigning a probability of 0.5 to failure of the FPS manual control. With this adjustment, the increase in unavailability of I

the AFW system, given actuation of the FPS water deluge system, was estimated to be2 x 10-5. Assuming typical AFW unavailability of 5 x 10-5 (discussed later), the combined AFW unavailability, given actuation of the FPS, was 7 x 10-5.

Using the same 2 x 10-5 increased unavailability for other safety systems in the event sequences of this issue, no significant effect was found because the other safety systems were less sensitive to the 2 x 10-5 estimate. This conclusion was consistent with the Issue 57 assessment.

Electric Door Lock Failures: At tl'ie time of the Surry pipe rupture event, water and steam saturated a security card-reader located approximately 50 feet from the break point. As a result, key-cards would not open plant doors. The control room doors were opened to provide access to the control room and security personnel were assigned to the control room to provide the access security. Onel operator was temporarily trapped in a stairway due to the card-reader failure. At the time I

of this evaluation, the Surry plant was considering installing electric override switches to remedy this problem.

In Issue 81, the impact of the electric lock (card-reader) failure at Surry was evaluated. The results from Issue 81 indicated that failure of electric locks, without override protection, may contribute approximately 2 % to core-melt accidents from pipe ruptures outside containment.

Frequency Estimate To estimate the core-melt frequency from ruptures in PWR secondary systems, an example PRA 5 4 was used together with additional information provided in NUREG/CR-2800. 6 4 The pertinent accident sequences were then adjusted to account for pipe ruptures in the secondary side of PWR plants. The accident sequences used in this analysis were TMQD,'TMKU, TMQH, I

and TML(PCSNR)L where:

TM - a loss of power conversion system (PCS) transient caused by other than loss-of-offsite power. For this analysis, TM corresponds to the secondary system pipe rupture frequency (3 x 10-- 2 /RY) resulting in loss of the main feedwater I

system (M = 1);

Q- the pressurizer safety/relief valve demanded opens (0.01) and any pressurizer safety/relief Ivalve fails to re-close

( 0.J0 5);

D - failure to provide sufficient ECCS injection (10-3);

K- failure of the RPS (2.6 x 10-5); I H- failure of the ECCS recirculation system (7 x 10-3 PCSNR failure to recover the PCS (0.5, as discussed earlier); I U- failure of the operator to start high pressure injection, or feed-and-bleed is initiated, but is unsuccessful. For this analysis, U = 0.2 was assumed; L- failure of the AFW system. For 3-train AFW system plants, a typical AFW unavailability was 1.8 x 10-5 /demand.

For 2-train AFW system plants, the goal of Issue 124 was to upgrade the AFW systems to 10-4/demand. Therefore, a typical value of 5 x 10- 5 /demand was used in this analysis.

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Table 3.139-1 includes the sequences with and without the effects of systems interactions from pipe ruptures in the secondary systems outside of containment.

Examination of the results indicate that collectively the systems interactions may increase the core-melt frequency from pipe ruptures in the secondary systems outside containment by approximately 20% (9 x 10- 8 /RY). The total core-melt frequency, with the systems interactions (SI) effects included, was estimated to be 5 x 10- 7 /RY.

TABLE 3.139-1

[Sequence Without (SI) I Communications (SI) IFPS (SI) Locked Doors (SI) TOTAL I

TMQD 1.50 x 10-8 3.00 x 10-10 ,neg. 3.0 x 10-10 1.56 x 10-8 TMQD 1.50 x 10. 7 ineg. neg. [3.0x 10-9 1.53x 10-7 ITMQH 1.05 x 10-7 7.40 x 10- 9 ;neg. 2.1 x 10-9 1.15 x 10-7 1.0 111 x 0 TML f 8 6 io 130x1f 1.50 x 10f 7 1.05 x 10-8 x 10-8 13.0 x 10-9 2.24 x 10-7 IIS-8 14.20 x 10i7 1.80 x 10-8 F 6 x 8 x 1u-- x 0 7' Consequence Estimate The core-melt sequences under consideration involve no large breaks initially in the reactor coolant system pressure boundary. The reactor is likely to be at high pressure until the core melts through the lower vessel head with a steady discharge of steam and gases through the PORV(s). These are conditions that may produce significant H2 generation and I combustion.

F~or these sequences, a 3% probability of containment failure due to H2 burn and a 1% probability of containment isolation i failure were used. If the containment does not fail by H2 burn or isolation failure, it was assumed to fail by basemat melt-through.

The conditional releases for these containment failure modes had a weighted average core-melt release of 1.7 x 105 I man-rem. The calculated releases were based on a core inventory typical of a 1120 MWe plant, a uniform population density of 340 persons per square mile from an exclusion area of one-half mile out to a 50-mile radius from the plant, no evacuation of people, no injestion pathways, and meteorology typical of a midwest site.

I The annual public risk from secondary side piping ruptures due to wall-thinning was the product of the core-melt frequency (5 x 10- 7 /RY) and the weighted average release (1.7 x 10-5 man-rem). Therefore, the public risk was 8.5 x 10-2 man-rem /RY. Assuming a remaining plant life of 30 years, the cumulative public risk was 3 man-rem/reactor.

Cost Estimate I Industry Cost: A possible solution for early detection of wall-thinning in carbon steel piping in the secondary systems was to implement and conduct inspection programs for these systems during each refueling outage. A report was prepared by EPR1 10 9 2 to provide guidance to the industry for conducting NDE of ferritic piping systems for wall-thinning caused by erosion/corrosion in nuclear and fossil power plants. The EPRI report contained the results of investigations of various NDE methods that may be applicable to the detection of erosion/corrosion effects. EPRI reported that virtually all plants used manual ultrasonic thickness measurements. Four utilities had performed automated ultrasonic thickness measurements from the outside surface of the piping. One EPRI source reported that an automated examination would cost approximately

$50,000 and take one week, whereas a manual team of two operators could perform the examination in one afternoon.

Therefore, the cost of the manual inspection was estimated to be $10,000 per outage.

The difference noted by EPRI was that the manual team would acquire data on a 4-inch grid pattern and the automated I system could acquire data continuously over the entire surface. Additional setup time was also required for the automated system. Therefore, the above $10,000 cost for the manual inspection could have been overestimated.

An additional cost associated with the inspections was the removal and disposal of asbestos insulation and re-insulation.

I These costs were reported to range from $300,000 to $750,000 per outage. In some plants, asbestos insulation was programatically being removed due to strict state and local guidelines associated with health hazards to workers from E of7 4/28/2008 1:11 AM

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Approximately half (44) of the 92 plants contacted in the EPRI survey had asbestos insulation. Thirty-two of the forty-four had at least partially replaced asbestos with other insulation, or were planning to remove the asbestos, and the remaining twelve plants were undecided.

Based on the above, any NRC requirement to conduct NDE inspections at each refueling outage could provide an additional incentive for the 12 plants (13% of all plants) to remove and replace the asbestos insulation with other types of insulation.

Therefore, on an average, the industry costs to remove and dispose of the asbestos insulation to facilitate NDE inspections was estimated to be a one-time cost of (0.13)($750,000 + $300,000)/2 = $68,000/plant. However, the argument could be made that the cost of asbestos removal could be driven by the state and local requirements, and not by NRC inspection I

requirements.

Assuming a remaining plant life of 30 years and a typical time between refueling outages of 1.5 years, the cumulative number of inspections that may be conducted during each refueling outage for each plant was 20. The annual cost over 30 years was (20)($10,000)/30 = $6,700/plant. The present value of the NDE annual costs over 30 years, considering a 5%

discount rate, was approximately $100,000/plant. The combined one-time costs for asbestos insulation removal and disposal and the sent value NDE cost over 30 years is $168,000/plant. I NRC Cost: It was estimated that one man-year of effort may be needed to reach a staff position on this issue and an additional man-year of effort to develop a Regulatory Guide or SRP Section. Assuming $100,000/man-year, the NRC costs were estimated to be $200,000. When distributed over approximately 100 plants, this cost was $2,000/plant.

Total Cost: The combined industry and NRC cost for the possible solution was estimated to be $170,000/plant.

Value/Impact Assessment Based on the estimated risk reduction of 3 man-rem/reactor and implementation costs of $170,000/plant for the possible solution (NDE examinations at each plant refueling outage), the value/impact score was given by:

3 rna- rem

$0.17M

= 17.6mma- rem/$M Other Considerations Accident Avoidance Cost: The present value of onsite property damage conditional on a core-melt for a remaining plant life of 30 years, assuming a 5% discount rate, was $20 b'illion. For a core-melt frequency of 5 x 10- 7 /R'Y attributed to pipe ruptures in the secondary systems, the accident avoidance cost by eliminating or significantly reducing the probability of pipe ruptures was $10,000/plant.

Industry Rupture Avoidance Cost: The rupture avoidance costs are the plant costs estimated to result from a pipe rupture in the secondary systems, assuming the plant responds as designed and no core-melt from potential equipment failures ensues. For a pipe rupture frequency of 3 x 10- 2 /RY, the chance of a pipe rupture in the secondary side can approach unity over the life of a plant.

To estimate the costs of plant repairs after a forced outage from a pipe rupture in the secondary system, historical plant operational data indicates that a best estimate repair cost from forced outages for a typical nuclear power plant is approximately $1,000/hour.1 0 8 2 The Trojan plant outage time following a pipe rupture in the secondary system was 6 days, whereas the Surry plant outage time lasted approximately 90 days. Based on the above, the plant repair costs from these I

two events was estimated to range from $140,000 to $2M. The replacement power costs resulting from the forced outages of 6 days for the Trojan plant and 90 days for the Surry plant were $3M and $45M, respectively; the cost of replacement power was estimated at $500,000/day.

It was assumed that the above cost estimates reflected lower and upper bound costs that could be represented by a log normal distribution with an error factor of 4. The combined repair costs and replacement power costs, adapted to a log normal distribution, yielded an estimated value of $17M as the mean plant costs resulting from a pipe rupture in the secondary systems.

I The $10,000/plant accident (core-melt) avoidance costs were small compared to the estimated rupture avoidance costs of 6 of 7 4/28/2008 1:11 AM

NRC: A Prioritization of Generic Safety Issues (NUREG-0933) - ISSUE... http://www.nrc.gov/reading-rm/doc-collections/nuregs/staff/sr0933/sec3...

$17M/plant. The low core-melt frequency of 5 x 10- 7 /RY drove down the accident avoidance costs. However, based on the estimated pipe rupture frequency of 3 x 10- 2 /RY, the chance of a pipe rupture in the secondary systems over the life of a plant approaches unity. Thus, the rupture avoidance costs dominated the combined accident and rupture avoidance costs.

When the implementation cost ($170,000/plant) is offset by the accident and rupture avoidance costs (a $17M/plant cost savings), the denominator of S becomes negative. The negative denominator of approximately $17M/plant indicates a substantial potential cost savings (industry incentive) by avoiding piping ruptures in the secondary systems.

I Occupational Safety: Erosion/corrosion-induced ruptures in high energy carbon steel piping lines described in this analysis resulted in injury and fatalities to plant personnel and contractor employees working in the area of the ruptures. At the time of the Surry pipe rupture, 8 contractor employees were working in the area of the pipe rupture; 6 of these individuals were I hospitilized for treatment of severe burns and 2 were treated at a clinic and released. Four of the severely burned individuals died and the other two were in serious to critical condition. One of the two remained in serious condition for more than a month after the accident. Following the pipe rupture at the Trojan plant, one member of the plant operating staff received first and second degree burns and was treated at a local hospital over a three-week period.

I CONCLUSION The estimated core-melt frequency of 5 x 10- 7 /Ry and the potential risk reduction of 3 man-rem/reactor indicated that pipe ruptures in the PWR secondary systems from erosion/corrosion-induced wall-thinning is of low safety significance to the public. Since inspection results indicated that erosion/corrosion wall-thinning of carbon steel piping is less prevalent in BWR plants, the above PWR risk estimates should be bounding. Therefore, as a generic safety issue, this issue would have been i given a low priority ranking. However, the erosion/corrosion-induced wall-thinning of carbon steel piping in secondary systems was not expected to be a significant cause of pipe ruptures. Pipe ruptures were more generalized as limiting faults:

postulated, but not expected to occur. Thus, knowledge and an understanding of this phenomena was limited. This analysis I indicated that, without adequate defensive methods or measures, pipe rupture induced by wall-thinning can be expected within the lifetime of a plant: an infrequent event with a higher frequency than the limiting fault (postulated) pipe ruptures.

The GAO concluded that the Surry accident initiated a new era of understanding regarding erosion/corrosion at nuclear power plants and demonstrated that unchecked erosion/corrosion can lead to a fatal accident. The GAO also concluded that NRC needed a mechanism to ensure that utilities periodically assess the integrity of piping systems to reduce the risk of future injury to plant personnel or damage to equipment caused by erosion/corrosion. The GAO recommended that NRC require utilities to:

(1) inspect all nuclearlplants to develop data regarding the extent that erosion/corrosion existed in piping systems, including straight sections of pipe; (2) replace piping that did not meet the industry's minimum allowable thickness standards; and I (3) periodically monitor piping systems and use the data developed during these inspections to monitor the spread of erosion/corrosion in the plants.

Based on the potential low public risk, the NRC need (References 1090, 1094, 1096) to establish a new position or I requirement on the previously unexpected phenomena, and a significant industry cost incentive to address and resolve the issue, this issue was classified as a Regulatory Impact issue by RES consistent with the ongoing levels of staff and. industry actions described. in SECY-88-501 0 9 4 and SECY-88-50A. 10 9 6 However, NRR considered the issue to be resolved based on:

(1) guidelines on erosion/corrosion in single-phase piping, as developed by NUMARC and found acceptable by the staff; (2)

.participation in a timely way by all 113 operating LWR plants; (3) acceptable analytical procedures for the evaluation and selection of piping to be inspected; (4) replacement of components as needed; and (5) a long-term as well as a short-term 1132 program for continuing evaluation and inspection of both single-phase and two-phase piping.

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UNITED STATES NUCLEAR REGULATORY COMMISSION ATOMIC SAFETY AND LICENSING BOARD Before Administrative Judges:

Alex S. Karlin, Chairman Dr. Richard E. Wardwell Dr. William H. Reed In the Matter of )

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ENTERGY NUCLEAR VERMONT YANKEE, LLC ) Docket No. 50-271-LR and ENTERGY NUCLEAR OPERATIONS, INC. ) ASLBP No. 06-849-03-LR

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(Vermont Yankee Nuclear Power Station) )

NEW ENGLAND COALITION, INC.

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NEC-JH 26 NUREG/CR-6909 ANL-06/08 Effect of LWR Coolant Environments on the Fatigue Life of Reactor Materials Final Report Argonne National Laboratory U.S. Nuclear Regulatory Commission Office of Nuclear Regulatory Research Washington, DC 20555-0001

I I

NUREG/CR-6909 ANL-06/08 I

Effect of LWR Coolant Environments on the I Fatigue Life of 3 Reactor Materials I Final Report I Manuscript Completed: November 2006 Date Published: February 2007 Prepared by O. K. Chopra and W. J. Shack Argonne National Laboratory 9700 South Cass Avenue Argonne, IL 60439 H. J. Gonzalez, NRC Project Manager Prepared for Division of Fuel, Engineering and Radiological Research Office of Nuclear Regulatory Research U.S. Nuclear Regulatory Commission Washington, DC 20555-0001 NRC Job Code N6187 41 I

I I

Abstract The ASME Boiler and Pressure Vessel Code provides rules for the design of Class I components of nuclear power plants. Figures 1-9.1 through 1-9.6 of Appendix I to Section III of the Code specify design curves for applicable structural materials. However, the effects of light water reactor (LWR) coolant environments are not explicitly addressed by the Code design curves. The existing fatigue strain-vs.-life (c-N) data illustrate potentially significant effects of LWR coolant environments on the fatigue resistance of pressure vessel and piping steels. Under certain environmental and loading conditions, fatigue lives in water relative to those in air can be a factor of z12 lower for austenitic stainless steels, z3 lower for Ni-Cr-Fe alloys, and ýl 7 lower for carbon and low-alloy steels. This report summarizes the work performed at Argonne National Laboratory on the fatigue of piping and pressure vessel steels in LWR environments.

The existing fatigue c-N data have been evaluated to identify the various material, environmental, and loading parameters that influence fatigue crack initiation, and to establish the effects of key parameters on the fatigue life of these steels. Fatigue life models are presented for estimating fatigue life as a function of material, loading, and environmental conditions. The environmental fatigue correction factor for incorporating the effects of LWR environments into ASME Section III fatigue evaluations is described.

The report also presents a critical review of the ASME Code fatigue design margins of 2 on stress (or strain) and 20 on life and assesses the possible conservatism in the current choice of design margins.

iii

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iv

Foreword This report summarizes, reviews, and quantifies the effects of the light-water reactor (LWR) environment on the fatigue life of reactor materials, including carbon steels, low-alloy steels, nickel-chromium-iron (Ni-Cr-Fe) alloys, and austenitic stainless steels. The primary purpose of this report is to provide the background and technical bases to support Regulatory Guide 1.207, "Guidelines for Evaluating Fatigue Analyses Incorporating the Life Reduction of Metal Components Due to the Effects of the Light-Water Reactor Environment for New Reactors."

Previously published related reports include NUREG/CR-5704, "Effects of LWR Coolant Environments on Fatigue Design Curves of Austenitic Stainless Steels," issued April 1999; NUREG/CR-6717, "Environmental Effects on Fatigue Crack Initiation in Piping and Pressure Vessel Steels," issued May 2001; NUREG/CR-6787, "Mechanism and Estimation of Fatigue Crack Initiation in Austenitic Stainless Steels in LWR Environments," issued August 2002; NUREG/CR-6815, "Review of the Margins for ASME Code Fatigue Design Curve - Effects of Surface Roughness and Material Variability," issued September 2003; and NUREG/CR-6583, "Effects of LWR Coolant Environments on Fatigue Design Curves of Carbon and Low-Alloy Steels," issued February 1998. This report provides a review of the existing fatigue &-Ndata for carbon steels, low-alloy steels, Ni-Cr-Fe alloys, and austenitic stainless steels to define the potential effects of key material, loading, and environmental parameters on the fatigue life of the steels. By drawing upon a larger database than was used in earlier published reports, the U.S. Nuclear Regulatory Commission (NRC) has been able to update the Argonne National Laboratory (ANL) fatigue life models used to estimate the fatigue curves as a function of those parameters. In addition, this report presents a procedure for incorporating environmental effects into fatigue evaluations. The database described in this report (and its predecessors) reinforces the position espoused by the NRC that a guideline for incorporating the LWR environmental effects in the fatigue life evaluations should be developed and that the design curves for the fatigue life of pressure boundary and internal components fabricated from stainless steel should be revised. Toward that end, this report proposes a method for establishing reference curves and environmental correction factors for use in evaluating the fatigue life of reactor components exposed to LWR coolants and operational experience.

Data described in this review have been used to define fatigue design curves in air that are consistent with the existing fatigue data. Specifically, the published data indicate that the existing code curves are nonconservative for austenitic stainless steels (e.g., Types 304, 316, and 316NG). Regulatory Guide 1.207 endorses the new stainless steel fatigue design curves presented herein for incorporation in fatigue analyses for new reactors. However, because of significant conservatism in quantifying other plant-related variables (such as cyclic behavior, including stress and loading rates) involved in cumulative fatigue life calculations, the design of the current fleet of reactors is satisfactory.

Brian W. Sheron, Director Office of Nuclear Regulatory Research U.S. Nuclear Regulatory Commission v

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vi

Contents Abstract ............................................................................................................... iii Foreword .............................................................................................................. v Executive Summary................................................................................................. xv Abbreviations ........................................................................................................ xvii Acknowledgments .................................................................................................. xix

1. Fatigue Analysis ..............................................................................................

2. Fatigue Life ..................................................................................................... 7

3. Fatigue Strain vs. Life Data................................................................................... 9 4 Carbon and Low-Alloy Steels ................................................................................ I11 4.1 Air Environment...........I............................................................................ 11 4.1.1 Experimental Data........................................................................ 11 4.1.2 Temperature ............................................................................... 12 4.1.3 Strain Rate................................................................................. 12 4.1.4 Sulfide Morphology ...................................................................... 13 4.1.5 Cyclic Strain Hardening Behavior...................................................... 13 4.1.6 Surface Finish............................................................................. 14 4.1 .7 Heat-to-Heat Variability ................................................................. 15 4.1.8 Fatigue Life Model ....................................................................... 17 4.1.9 Extension of the Best-Fit Mean Curve from 106 to 10 11 Cycles ................... 18 4.1.10 Fatigue Design Curve................................................................... 19 4.2 LWR Environment................................................................................... 21 4.2.1 Experimental Data........................................................................ 21 4.2.2 Strain Rate.................................................................................. 22 4.2.3 Strain Amplitude.......................................................................... 23 vii

4.2.4 Temperature............................................................................... 26 4.2.5 Dissolved Oxygen ........................................................................ 29 4.2.6 Water Conductivity....................................................................... 30 4.2.7 Sulfur Content in Steel................................................................... 30 4.2.8 Tensile Hold Period ...................................................................... 31 4.2.9 Flow Rate.................................................................................. 33 4.2.10 Surface Finish ........................................................................... 34 4.2.11 Heat-to-Heat Variability................................................................ 35 4.2.12 Fatigue Life Model ..................................................................... 36 4.2.13 Environmental Fatigue Correction Factor ........................................... 38 4.2.14 Modified Rate Approach............................................................... 38 5 Austenitic Stainless Steels..................................................................................... 41 5.1 Air Environment ...................................................................................... 41 5.1.1 Experimental Data........................................................................ 41 5.1.2 Specimen Geometry ...................................................................... 43 5.1.3 Temperature............................................................................... 43 5.1.4 Cyclic Strain Hardening Behavior ..................................................... 44 5.1.5 Surface Finish............................................................................. 45 5.1.6 Heat-to-Heat Variability ................................................................. 45 5.1 .7 Fatigue Life Model....................................................................... 46 5.1.8 New Fatigue Design Curve.............................................................. 48 5.2 LWR Environment.............................................. ****".....

...... 49 5.2.1 Experimental Data........................................................................ 49 5.2.2 Strain Amplitude.......................................................................... 51 5.2.3 Hold-Time Effects........................................................................ 52 viii

5.2.4 Strain Rate ................................................................................. 53 5.2.5 Dissolved Oxygen ........................................................................ 54 5.2.6 Water Conductivity....................................................................... 54 5.2.7 Temperature............................................................................... 55 5.2.8 Material Heat Treatment................................................................. 56 5.2.9 Flow Rate.................................................................................. 57 5.2.10 Surface Finish ........................................................................... 58 5.2.1 1 Heat-to-Heat Variability................................................................ 58 5.2.12 Cast Stainless Steels .................................................................... 60 5.2.13 Fatigue Life Model...................................................................... 61 5.2.14 Environmental Correction Factor..................................................... 63 6 Ni-Cr-Fe Alloys and Welds ................................................................................... 65 6.1 Air Environment ...................................................................................... 65 6.1.1 Experimental Data........................................................................ 65 6.1.2 Fatigue Life Model........................................................................ 66 6.2 LWR Environment ................................................................................... 67 6.2.1 Experimental Data........................................................................ 67 6.2.2 Effects of Key Parameters................................................ I............... 68 6.2.3 Environmental Correction Factor ...................................................... 68 7 Margins in A SME Code Fatigue Design Curves........................................................... 71 7.1 Material Variability and Data Scatter .............................................................. 73 7.2 Size and Geometry.................................................................................... 73

7. 3 Surface Finish ......................................................................................... 74 7.4 Loading Sequence..................................................................................... 74 7.5 Fatigue Design Curve Margins Summarized...................................................... 75 ix

8 Sum mary ............................................................................................................................................... 79 References ................................................................................................................................................... 83 APPENDIX A ............................................................................................................................................. A.I

Figures

1. Schematic illustration of growth of short cracks in smooth specimens as a function of fatigue life fraction and crack velocity as a function of crack depth ........................................................ 7

2. Crack growth rates plotted as a function of crack depth for A533-Gr B low-alloy steel and Type 304 SS in air and LW R environm ents .................................................................................. 8

3. Fatigue strain vs. life data for carbon and low-alloy steels in air at room temperature .............. II

4. Fatigue strain vs. life data for carbon and low-alloy steels in air at 288°C ................................. 12

5. Effect of strain rate and temperature on cyclic stress of carbon and low-alloy steels ................ 13

6. Effect of surface finish on the fatigue life of A106-Gr B carbon steel in air at 289°C ............... 14

7. Estimated cumulative distribution of constant A in the ANL models for fatigue life for heats of carbon steels and low -alloy steels in air .................................................................................... 16

8. Experimental and predicted fatigue lives of carbon steels and low-alloy steels in air ................ 18

9. Fatigue design curve for carbon steels in air .................................................................................. 20

10. Fatigue design curve for low -alloy steels in air .............................................................................. 20

11. Strain amplitude vs. fatigue life data for A533-Gr B and AI06-Gr B steels in air and high-dissolved-oxygen w ater at 288°C .................................................................................................. . 21

12. Dependence of fatigue life of carbon and low-alloy steels on strain rate .................................... 23

13. Fatigue life of A106-Gr B carbon steel at 288°C and 0.75% strain range in air and water environments under different loading waveforms ........................................................................ 24

14. Fatigue life of carbon and low-alloy steels tested with loading waveforms where slow strain rate is applied during a fraction of tensile loading cycle ............................................................... 25

15. Experimental values of fatigue life and those predicted from the modified rate approach w ithout consideration of a threshold strain ..................................................................................... 26

16. Change in fatigue life of A333-Gr 6 carbon steel with temperature and DO .................. 26

17. Dependence of fatigue life on temperature for carbon and low-alloy steels in water. ................ 27

18. Waveforms for change in temperature during exploratory fatigue tests ....................................... 28

19. Fatigue life of A333-Gr 6 carbon steel tube specimens under varying temperature, indicated by ho rizo ntal bars ............................................................................................................................. 28

20. Dependence on DO of fatigue life of carbon steel in high-purity water ....................................... 29 xi

21. Effect of strain rate on fatigue life of low-alloy steels with different S contents ........................ 30

22. Effect of strain rate on the fatigue life of A333-Gr 6 carbon steels with different S contents... 31

23. Fatigue life of AI06-Gr B steel in air and water environments at 288°C, 0.78% strain range, and hold period at peak tensile strain .............................................................................................. 32

24. Effect of water flow rate on fatigue life of A333-Gr 6 carbon steel at 289°C and strain amplitude and strain rates of 0.3% and 0.01%/s and 0.6% and 0.001%/s .................................... 33

25. Effect of flow rate on low-cycle fatigue of carbon steel tube bends in high-purity water at 2 4 00 C ............................................................................................................................................ 34

26. Effect of surface roughness on fatigue life of A106-Gr B carbon steel and A533 low-alloy steel in air and high-purity water at 289°C .................................................................................. 34

27. Estimated cumulative distribution of parameter A in the ANL models for fatigue life for heats of carbon and low-alloy steels in LWR environments ........................................................ 35

28. Experimental and predicted fatigue lives of carbon steels and low-alloy steels in LWR en v iro nm en ts . ................................................................................................................................... 37

29. Application of the modified rate approach to determine the environmental fatigue correction factor Fen during a transient ............................................................................................................. 39

30. Fatigue c-N behavior for Types 304, 316, and 316NG austenitic stainless steels in air at vario us tem peratures ........................................................................................................................ 41

31. Influence of specimen geometry on fatigue life of Types 304 and 316 stainless steel ................ 43

32. Influence of temperature on fatigue life of Types 304 and 316 stainless steel in air ................... 43

33. Effect of strain amplitude, temperature, and strain rate on cyclic strain-hardening behavior of Types 304 and 316N G SS in air ..................................................................................................... . 44

34. Effect of surface roughness on fatigue life of Type 316NG and Type 304 SSs in air ................. 45

35. Estimated cumulative distribution of constant A in the ANL model for fatigue life for heats of austenitic SS in air ....................................................................................................................... 46

36. Experimental and predicted fatigue lives of austenitic SSs in air ................................................. 47

37. Fatigue design curve for austenitic stainless steels in air ............................................................. 48

38. Strain amplitude vs. fatigue life data for Type 304 and Type 316NG SS in water at 288'C ...... 49

39. Higher-magnification photomicrographs of oxide films that formed on Type 316NG stainless steel in simulated PWR water and high-DO water ......................................................... 50 xii

40. Schematic of the corrosion oxide film formed on austenitic stainless steels in LWR environments .............................................. ............... 50

41. Effects of environment on formation of fatigue cracks in Type 316NG SS in air and low-DO water at 2 8 8'C .................................................................................................................................. 51

42. Results of strain rate change tests on Type 316 SS in low-DO water at 325°C .......................... 52

43. Fatigue life of Type 304 stainless steel tested in high-DO water at 260-288°C with trapezoidal or triangular waveform ................................................................................................. . 52

44. Dependence of fatigue lives of austenitic stainless steels on strain rate in low-DO water ......... 53

45. Dependence of fatigue life of Types 304 and 316NG stainless steel on strain rate in high-and low -D O water at 288°C ............................................................................................................ 53

46. Effects of conductivity of water and soaking period on fatigue life of Type 304 SS in high-D O water ........................................................................................................................................... 55

47. Change in fatigue lives of austenitic stainless steels in low-DO water with temperature .......... 55

48. Fatigue life of Type 316 stainless steel under constant and varying test temperature ................. 56

49. The effect of material heat treatment on fatigue life of Type 304 stainless steel in air, BWR and PWR environments at 289°C, z0.38% strain amplitude, sawtooth waveform, and 0.004% /s tensile strain rate .............................................................................................................. 57

50. Effect of water flow rate on the fatigue life of austenitic SSs in high-purity water at 289°C .... 57

51. Effect of surface roughness on fatigue life of Type 316NG and Type 304 stainless steels in air and high-purity w ater at 289°C ................................................................................................. 58

52. Estimated cumulative distribution of constant A in the ANL model for fatigue life for heats of austenitic S Ss in water ................................................................................................................. 59

53. Dependence of fatigue lives of CF-8M cast SSs on strain rate in low-DO water at various strain am p litud es ............................................................................................................................... 60

54. Estimated cumulative distribution of constant A in the ANL model for fatigue life of wrought and cast austenitic stainless steels in air and water environments ............................................... 61

55. Experimental and predicted values of fatigue lives of austenitic SSs in LWR environments ... 63

56. Fatigue s-N behavior for Alloys 600 and 690 in air at temperatures between room tem perature and 3 15'C .................................................................................................................... 65

57. Fatigue s-N behavior for Alloys 82, 182, 132, and 152 welds in air at various temperatures... 66

58. Fatigue s-N behavior for Alloy 600 and its weld alloys in simulated BWR water at z289°C... 67 xiii

59. Fatigue c-N behavior for Alloys 600 and 690 and their weld alloys in simulated PWR water at 3 15 or 32 5°C ................................................................................................................................. 67

60. Dependence of fatigue lives of Alloys 690 and 600 and their weld alloys in PWR water at 325°C and Alloy 600 in BW R water at 289°C ............................................................................... 68

61. The experimental and estimated fatigue lives of various Ni alloys in BWR and PWR env ironm ents .................................................................................................................................... 69

62. Fatigue data for carbon and low-alloy steel and-Type 304 stainless steel components .............. 72

63. Estimated cumulative distribution of parameter A in the ANL models that represent the fatigue life of test specimens and actual components in air ........... ................... 77 Tables

1. Sources of the fatigue c-N data on reactor structural materials in air and water environments. 9

2. Values of parameter A in the ANL fatigue life model for carbon steels in air and the margins on life as a function of confidence level and percentage of population bounded ........................ 16

3. Values of parameter A in the ANL fatigue life model for low-alloy steels in air and the margins on life as a function of confidence level and percentage of population bounded .......... 17

4. Fatigue design curves for carbon and low-alloy steels and proposed extension to 1011 cycles. 20

5. Fatigue data for STS410 steel at 289°C in water with I ppm DO and trapezoidal waveform.... 33

6. Values of parameter A in the ANL fatigue life model for carbon steels in water and the margins on life as a function of confidence level and percentage of population bounded......... 36

7. Values of parameter A in the ANL fatigue life model for low-alloy steels in water and the margins on life as a function of confidence level and percentage of population bounded .......... 36

8. Values of parameter A in the ANL fatigue life model and the margins on life for austenitic SSs in air as a function of confidence level and percentage of population bounded>. .................. 46

9. The new and current Code fatigue design curves for austenitic stainless steels in air ................ 48

10. Values of parameter A in the ANLI. fatigue life model and the margins on life for austenitic SSs in water as a function of confidence level and percentage of population bounded .............. 59

11. The median value of A and standard deviation for the various fatigue E-N data sets used to evaluate m aterial variability and data scatter ................................................................................. 73

12. Factors on life applied to mean fatigue c-N curve to account for the effects of various material, loading, and environmental parameters .................................... 76

13. Margin applied to the mean values of fatigue life to bound 95% of the population ............ 77 xiv

Executive Summary Section 1I1,Subsection NB, of the ASME Boiler and Pressure Vessel Code contains rules for the design of Class I components of nuclear power plants. Figures 1-9.1 through 1-9.6 of Appendix I to Section III specify the Code design fatigue curves for applicable structural materials. However, Section II1, Subsection NB-3121 of the Code states that the effects of the coolant environment on fatigue resistance of a material were not intended to be addressed in these design curves. Therefore, the effects of environment on the fatigue resistance of materials used in operating pressurized water reactor (PWR) and boiling water reactor (BWR) plants, whose primary-coolant pressure boundary components were designed in accordance with the Code, are uncertain.

The current Section-Ill design fatigue curves of the ASME Code were based primarily on strain-controlled fatigue tests of small polished specimens at room temperature in air. Best-fit curves to the experimental test data were first adjusted to account for the effects of mean stress and then lowered by a factor of 2 on stress and 20 on cycles (whichever was more conservative) to obtain the design fatigue curves. These factors are not safety margins but rather adjustment factors that must be applied to experimental data to obtain estimates of the lives of components. Recent fatigue-strain-vs.-life (F-N) data obtained in the U.S. and Japan demonstrate that light water reactor (LWR) environments can have potentially significant effects on the fatigue resistance of materials. Specimen lives obtained from tests in simulated LWR environments can be much shorter than those obtained from corresponding tests in air.

This report reviews the existing fatigue s-N data for carbon and low-alloy steels, wrought and cast austenitic stainless steels (SSs), and nickel-chromium-iron (Ni-Cr-Fe) alloys in air and LWR environments. The effects of various material, loading, and environmental parameters on the fatigue lives of these steels are summarized. The results indicate that in air, the ASME mean curve for low-alloy steels is in good agreement with the available experimental data, and the curve for carbon steels is somewhat conservative. However, in air, the ASME mean curve for SSs is not consistent with the experimental data at strain amplitudes <0.5% or stress amplitudes <975 MPa (<141 ksi); the ASME mean curve is nonconservative. The results also indicate that the fatigue data for Ni-Cr-Fe alloys are not consistent with the current ASME Code mean curve for austenitic SSs.

The fatigue lives of carbon and low-alloy steels, austenitic SSs, and Ni-Cr-Fe alloys are decreased in LWR environments. The reduction depends on some key material, loading, and environmental parameters. The fatigue data are consistent with the much larger database on enhancement of crack growth rates in these materials in LWR environments. The key parameters that influence fatigue life in these environments, e.g., temperature, dissolved-oxygen (DO) level in water, strain rate, strain (or stress) amplitude, and, for carbon and low-alloy steels, S content of the steel, have been identified. Also, the range of the values of these parameters within which environmental effects are significant has been clearly defined. If these critical loading and environmental conditions exist during reactor operation, then environmental effects will be significant and need to be included in the ASME Code fatigue evaluations.

Fatigue life models developed earlier to predict fatigue lives of small smooth specimens of carbon and low-alloy steels, wrought and cast austenitic SSs, and Ni-Cr-Fe alloys as a function of material, loading, and environmental parameters have been updated/revised by drawing upon a larger fatigue s-N database. The functional form and bounding values of these parameters were based on experimental observations and data trends. An approach that can be used to incorporate the effects of LWR coolant environments into the ASME Code fatigue evaluations, based on the environmental fatigue correction factor, Fen, is discussed. The fatigue usage for a specific stress cycle of load set pair based on the Code fatigue design curves is multiplied by the correction factor to account for environmental effects.

xv

The report also presents a critical review of the ASME Code fatigue design margins of 2 on stress and 20 on life and assesses the possible conservatism in the current choice of design margins. Although these factors were intended to be somewhat conservative, they should not be considered safety margins.

These factors cover the effects of variables that can influence fatigue life but were not investigated in the experimental data that were used to obtain the fatigue design curves. Data available in the literature have been reviewed to evaluate the margins on cycles and stress that are needed to account for such differencesI and uncertainties. Monte Carlo simulations were performed to determine the margin on cycles needed to obtain a fatigue design curve that would provide a somewhat conservative estimate of the number of cycles to initiate a fatigue crack in reactor components. The results suggest that for both carbon and low-I alloy steels and austenitic SSs, the current ASME Code requirements of a factor of 20 on cycle to account for the effects of material variability and data scatter, as well as size, surface finish, and loading history in low cycle fatigue, contain at least a factor of 1.7 conservatism. Thus, to reduce this conservatism, fatigue design curves have been developed from the ANL fatigue life model by first correcting for mean stress effects, and then reducing the mean-stress adjusted curve by a factor of 2 on stress or 12 on cycles, whichever is more conservative. These design curves are consistent with the existing fatigue c-N data.

A detailed procedure for incorporating environmental effects into fatigue evaluations is presented.

xvI

Abbreviations ANL Argonne National Laboratory ANN Artificial Neural Network ASME American Society of Mechanical Engineers BWR Boiling Water Reactor CGR Crack Growth Rate CUF Cumulative Usage Factor DO Dissolved Oxygen EAC Environmentally Assisted Cracking ECP Electrochemical Potential EPR Electrochemical Potentiodynamic Reactivation EPRI Electric Power Research Institute GE General Electric Co.

IHI Ishikawajima-Harima Heavy Industries KWU Kraftwerk Union Laboratories LWR Light Water Reactor MA Mill Annealed MEA Materials Engineering Associates MHI Mitsubishi Heavy Industries MPA Materialprufungsanstalt MSC Microstructurally Small Crack NRC Nuclear Regulatory Commission ORNL Oak Ridge National Laboratory PVRC Pressure Vessel Research Council PWR Pressurized Water Reactor RCS Reactor Coolant System RT Room Temperature SCC Stress Corrosion Cracking SICC Strain Induced Corrosion Cracking SS Stainless Steel UTS Ultimate Tensile Strength WRC Welding Research Council xvii

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xviii

Acknowledgments The authors thank W. H. Cullen, Jr., and J. Fair for their helpful comments. This work is sponsored by the Office of Nuclear Regulatory Research, U.S. Nuclear Regulatory Commission, under NRC Job Code N6187; Project Manager: H. J. Gonzalez.

xix

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xx

1. Fatigue Analysis The American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel Code Section III, Subsection NB, which contains rules for the design of Class I components for nuclear power plants, recognizes fatigue as a possible mode of failure in pressure vessel steels and piping materials.

Fatigue has been a major consideration in the design of rotating machinery and aircraft, where the components are subjected to a very large number of cycles (e.g., high-cycle fatigue) and the primary concern is the endurance limit, i.e., the stress that can be applied an infinite number of times without failure. However, cyclic loadings on a reactor pressure boundary component occur because of changes in mechanical and thermal loadings as the system goes from one load, set (e.g., pressure, temperature, moment, and force loading) to another. The number of cycles applied during the design life of the component seldom exceeds 105 and is typically less then a few thousand (e.g., low-cycle fatigue). The main difference between high-cycle and low-cycle fatigue is that the former involves little or no plastic strain, whereas the latter involves strains in excess of the yield strain. Therefore, design curves for low-cycle fatigue are based on tests in which strain rather than stress is the controlled variable.

The ASME Code fatigue evaluation procedures are described in NB-3200, "Design by Analysis,"

and NB-3600, "Piping Design." For each stress cycle or load set pair, an individual fatigue usage factor is determined by the ratio of the number of cycles anticipated during the lifetime of the component to the allowable cycles. Figures 1-9.1 through 1-9.6 of the mandatory Appendix I to Section III of the ASME Boiler and Pressure Vessel Code specify fatigue design curves that define the allowable number of cycles as a function of applied stress amplitude. The cumulative usage factor (CUF) is the sum of the individual usage factors, and ASME Code Section III requires that at each location the CUF, calculated on the basis of Miner's rule, must not exceed 1.

The ASME Code fatigue design curves, given in Appendix I of Section III, are based on strain-controlled tests of small polished specimens at room temperature in air. The design curves have been developed from the best-fit curves to the experimental fatigue-strain-vs.-life (c-N) data, which are expressed in terms of the Langer equation' of the form ea = Al(N)-n1 +A2, (,)

where Ea is the applied strain amplitude, N is the fatigue life, and Al, A2, and nI are coefficients of the model. Equation 1 may be written in terms of stress amplitude Sa instead of Fa. The stress amplitude is the product of Ea and elastic modulus E, i.e., Sa = E-a (stress amplitude is one-half the applied stress 2

range). The current ASME Code best-fit or mean curve described in the Section III criteria document for various steels is given by E 100 )B a -4 -Nf 0 - +B , (2) where E is the elastic modulus, Nf is the number of cycles to failure, and Af and Bf are constants related to reduction in area in a tensile test and endurance limit of the material at 107 cycles, respectively. The current Code mean curve for carbon steel is expressed as 5

Sa = 59,734 (Nf)-0 + 149.2, (3)

I

I for low-alloy steel, as Sa = 49,222 (Nf)-0 5

+ 265.4, (4) m and for austenitic SSs, as Sa = 58,020 (Nf)-0-5 + 299.9. (5)

Note that because most of the data used to develop the Code mean curve were obtained on specimens that were tested to failure, in the Section III criteria document, fatigue life is defined as cycles to failure.

I Accordingly, the ASME Code fatigue design curves are generally considered to represent allowable number of cycles to failure. However, in Appendix I to Section III of the Code the design curves are simply described as stress amplitude (Sa) vs. number of cycles (N).

I In the fatigue tests performed during the last three decades, fatigue life is defined in terms of the number of cycles for tensile stress to decrease 25% from its peak or steady-state value. For typical cylindrical specimens used in these studies, this corresponds to the number of cycles needed to produce an f 3-mm-deep crack in the test specimen. Thus, the fatigue life of a material is actually being described in terms of three parameters, viz., strain or stress, cycles, and crack depth. The best-fit curve to the existing fatigue c-N data describes, for given strain or stress amplitude, the number of cycles needed to develop a 3-mm deep crack. The fatigue c-N data are typically expressed by rewriting Eq. 1 as ln(N) = A - B ln(ca - C), (6) where A, B, and C are constants; C represents the fatigue limit of the material; and B is the slope of the log-log plot of fatigue c-N data. The ASME Code mean-data curves (i.e., Eqs. 3-5) may be expressed in terms of Eq. 6 as follows. The fatigue life of carbon steels is given by ln(N) = 6.726 - 2.0 ln(ca - 0.072), (7) for low-alloy steels, by ln(N) = 6.339 - 2.0 ln(Ea - 0.1 2 8 ), (8) and, for austenitic SSs, by ln(N) = 6.954 - 2.0 ln(ca - 0.167). (9)

The Code fatigue design curves have been obtained from-the best-fit (or mean-data) curves by first adjusting for the effects of mean stress using the modified Goodman relationship given by S =Sa!u YI for Sa < Oy, (10) and S' S, for Sa > Gy, (11) 2 I

where S, is the adjusted value of stress amplitude, and Gy and au are yield and ultimate strengths of the material, respectively. Equations 10 and 11 assume the maximum possible mean stress and typically give a conservative adjustment for mean stress. The fatigue design curves are then obtained by reducing the fatigue life at each point on the adjusted best-fit curve by a factor of 2 on strain (or stress) or 20 on cycles, whichever is more conservative.

The factors of 2 and 20 are not safety margins but rather adjustment factors that should be applied to the small-specimen data to obtain reasonable estimates of the lives of actual reactor components. As described in the Section III criteria document, 2 these factors were intended to account for data scatter (including material variability) and differences in surface condition and size between the test specimens and actual components. In comments about the initial scope and intent of the Section III fatigue design procedures Cooper 3 states that the factor of 20 on life was regarded as the product of three subfactors:

Scatter of data (minimum to mean) 2.0 Size effect 2.5 Surface finish, atmosphere, etc. 4.0 Although the Section III criteria document 2 states that these factors were intended to cover such effects as environment, Cooper3 further states that the term "atmosphere" was intended to reflect the effects of an industrial atmosphere in comparison with an air-conditioned laboratory, not the effects of a specific coolant environment. Subsection NB-3 121 of Section III of the Code explicitly notes that the data used to develop the fatigue design curves (Figs. 1-9.1 through 1-9.6 of Appendix I to Section 11) did not include tests in the presence of corrosive environments that might accelerate fatigue failure. Article B-2131 in Appendix B to Section III states that the owner's design specifications should provide infornation about any reduction to fatigue design curves that is necessitated by environmental conditions.

Existing fatigue c-N data illustrate potentially significant effects of light water reactor (LWR) coolant environments on the fatigue resistance of carbon and low-alloy steels and wrought and cast austenitic SSs.445 Laboratory data indicate that under certain reactor operating conditions, fatigue lives of carbon and low-alloy steels can be a factor of 17 lower in the coolant environment than in air.

Therefore, the margins in the ASME Code may be less conservative than originally intended.

The fatigue c-N data are consistent with the much larger database on enhancement of crack growth rates (CGRs) in these materials in simulated LWR environments. The key parameters that influence fatigue life in these environments, e.g., temperature, dissolved-oxygen (DO) level in water, strain rate, strain (or stress) amplitude, and, for carbon and low-alloy steels, S content of the steel, have been identified. Also, the range of the values of these parameters within which environmental effects are significant has been clearly defined. If these critical loading and environmental conditions exist during reactor operation, then environmental effects will be significant and need to be included in the ASME Code fatigue evaluations. Experience with nuclear power plants worldwide indicates that the critical range of loading and environmental conditions that leads to environmental effects on fatigue crack 45 6 1 initiation can occur during plant operation. -

Many failures of reactor components have been attributed to fatigue; examples include piping, nozzles, valves, and pumps. 46 -53 The mechanism of cracking in feedwater nozzles and piping has been attributed to corrosion fatigue or strain-induced corrosion cracking (SICC). 5 4-56 A review of significant occurrences of corrosion fatigue damage and failures in various nuclear power plant systems has been presented in an Electric Power Research Institute (EPRI) report. 45 In piping components, several failures were associated with thermal loading due to thermal stratification and striping. Thermal stratification is 3

I caused by the injection of low-flow, relatively cold feedwater during plant startup, hot standby, or variations below 20% of full power, whereas thermal striping is caused by rapid, localized fluctuations of the interface between hot and cold feedwater. Significant cracking has also occurred in nonisolable piping connected to a PWR reactor coolant system (RCS). In most cases, thermal cycling was caused by interaction of hot RCS fluid from turbulent penetration at the top of the pipe, and cold valve leakage fluid that had stratified at the bottom of the pipe. Lenz et al. 55 have shown that in feedwater lines, strain rates are 10- 3 5%/s due to thermal stratification and 10-1%/s due to thermal shock. They also have reported I

that thermal stratification is the primary cause of crack initiation due to SICC. Full-scale mock-up tests to generate thermal stratification in a pipe in a laboratory have confirmed the applicability of laboratory data to component behavior.4 4,62 A study conducted on SS pipe bend specimens in simulated PWR I

primary water at 240'C concluded that reactor coolant environment can have a significant effect on the fatigue life of SSs. 63 Relative to the fatigue life in an inert environment, life in the PWR environment at ai strain amplitude of 0.52% was decreased by factor of 5.8 and 2.8 at strain rates of 0.0005%/s and n 0.01%/s, respectively. These values show excellent agreement with the values predicted from the correlations presented in Section 5.2.14 of this report. 3 Thermal loading due to flow stratification or mixing was not included in the original design basis analyses. Regulatory evaluation has indicated that thermal-stratification cycling can occur in all PWR surge lines. 64 In PWRs, the pressurizer water is heated to z227°C. The hot water, flowing at a very low rate from the pressurizer through the surge line to the hot-leg piping, rides on a cooler water layer. The I

thermal gradients between the upper and lower parts of the pipe can be as high as 149°C.

Two approaches have been proposed for incorporating the environmental effects into ASME Section 1II fatigue evaluations for primary pressure boundary components in operating nuclear power plants: (a) develop new fatigue design curves for LWR applications, or (b) use an environmental fatigue correction factor to account for the effects of the coolant environment.

In the first approach, following the same procedures used to develop the current fatigue design curves of the ASME Code, environmentally adjusted fatigue design curves are developed from fits to experimental data obtained in LWR environments. ) Interim fatigue design curves that address I

environmental effects on the fatigue life of carbon and low-alloy steels and austenitic SSs were first proposed by Majumdar et al. 65 Fatigue design curves based on a more rigorous statistical analysis of experimental data were developed by Keisler et al. 6 6 These design curves have subsequently been revised on the basis of updated ANL models. 4 ,6,3 8,39 However, because, in LWR environments, the fatigue life of carbon and low-alloy steels, nickel-chromium-iron (Ni-Cr-Fe) alloys, and austenitic SSs depends on several loading and environmental parameters, such an approach would require developing several design curves to cover all possible conditions encountered during plant operation. Defining the number of these design curves or the loading and environmental conditions for the curves is not easy.

The second approach, proposed by Higuchi and Jida, 13 considers the effects of reactor coolant environments on fatigue life in terms of an environmental fatigue correction factor, Fen, which is the ratio of fatigue life in air at room temperature to that in water under reactor operating conditions. To incorporate environmental effects into fatigue evaluations, the fatigue usage factor for a specific stress cycle or load set pair, based on the ASME Code design curves, is multiplied by the environmental fatigue correction factor. Specific expressions for Fen, based on the Argonne National Laboratory (ANL) fatigue life models, have been developed. 39 Such an approach is relatively simple and is recommended in this report.

I I

This report presents an overview of the existing fatigue s-N data for carbon and low-alloy steels, Ni-Cr-Fe alloys, and wrought and cast austenitic SSs in air and LWR environments. The data are evaluated to (a) identify the various material, environmental, and loading parameters that influence fatigue crack initiation and (b) establish the effects of key parameters on the fatigue life of these steels.

Fatigue life models, presented in earlier reports, for estimating fatigue life as a function of material, loading, and environmental conditions have been updated using a larger database. The Fen approach for incorporating, effects of LWR environments into ASME Section III fatigue evaluations is described. The report also presents a critical review of the ASME Code fatigue design margins of 2 on stress (or strain) and 20 on life and assesses the possible conservatism in the current choice of design margins.

5

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6

2. Fatigue Life The formation of surface cracks and their growth to an engineering size (3-mm deep) constitute the fatigue life of a material, which is represented by the fatigue e-N curves. Fatigue life has conventionally been divided into two stages: initiation, expressed as the number of cycles required to form microcracks on the surface; and propagation, expressed as cycles required to propagate the surface cracks to engineering size. During cyclic loading of smooth test specimens, surface cracks 10 ptm or longer form early in life (i.e., <10% of life) at surface irregularities either already in existence or produced by slip bands, grain boundaries, second-phase particles, etc. 4,5 Thus, fatigue life may be considered to constitute propagation of cracks from 10 to 3000 Ltm long.

Mechanically Small Crack /

SMicrostructurall* . A(53

/-_ Small Crack /

/  ! _ 7 G Mechanically Small Crack AG2 " .

'a (Stage IITensile Crack) /) >- -- ~.

" I,/

,- ." " (1) --

>U - \

AG A 51 1 * *] , Linear-elastic or

',/Small Crack (MSC) Mirsrcrly. AG*

S(StageIShearCrack).A 2 > AGI CJ:roacksng i fracture mechanics 3 > A( 2 > AG 1 I I i I i i i [ i 00.2 0.4 0.6 0.8.1 Life Fraction Crack Depth (a) (b)

Figure 1. Schematic illustration of (a) growth of short cracks in smooth specimens as a function of fatigue life fraction and (b) crack velocity as a function of crack depth.

A schematic illustration of the initiation and propagation stages of fatigue life is shown in Fig. 1.

The initiation stage involves growth of "microstructurally small cracks" (MSCs), characterized by decelerating crack growth (Region AB in Fig. Ia). The propagation stage involves growth of "mechanically small cracks," characterized by accelerating crack growth (Region BC in Fig. la). The growth of the MSCs is very sensitive to microstructure. 5 Fatigue cracks greater than a critical depth show little or no influence ofmicrostructure and are considered mechanically small cracks. Mechanically small cracks correspond to Stage Ia (tensile) cracks, which are characterized by striated crack growth, with the fracture surface normal to the maximum principal stress. Various criteria, summarized in Section 5.4.1 of Ref. 6, have been used to define the crack depth for transition from microstructurally to mechanically small crack. The transition crack depth is a function of applied stress () and epicrostructure of the material; actual values may range from 150 to 250 pgm. At low enough stress levels (Aa 1 ), the transition from MSC growth to accelerating crack growth does not occur. This circumstance represents the fatigue limit for the smooth specimen. Although cracks can form below the fatigue limit, they can grow to engineering size only at stresses greater than the fatigue limit. The fatigue limit for a material is applicable only for constant loading condtonins. Under variable loading conditions, MSCs can grow at high stresses (Aa 3 ) to depths larger than the transition crack depth and then can continue to grow at stress levels below the fatigue limit (Adrj).

7

I Studies on the formation and growth characteristics of short cracks in smooth fatigue specimens in LWR environments indicate that the decrease in fatigue life in LWR environments is caused primarily by the effects of the environment on the growth of MSCs (i.e., cracks <200 jim deep) and, to a lesser extent, on the growth of mechanically small cracks. 4 ,7 Crack growth rates measured in smooth cylindrical fatigue specimens of A533-Gr B low-alloy steel and austenitic Type 304 SSs in LWR environments and air are shown in Fig. 2. The results indicate that in LWR environments, the period spent in the growth of I MSCs (region ABC in Fig. Ia) is decreased. For the A533-Gr B steel, only 30-50 cycles are needed to form a 100-mm crack in high-DO water, whereas z450 cycles are required to form a 100-mm crack in low-DO water and more than 3000 cycles in air. These values correspond to average growth rates of zi2.5, 0.22, and 0.033 jLm/cycle in high-DO water, low-DO water, and air, respectively. Relative to air, CGRs for A533-Gr B steel in high-DO water are nearly two orders of magnitude higher for crack sizes

<100 pm, and one order of magnitude higher for crack sizes >100 im. I ti I i i I i, ' i , I ' i i' i iI 102 A533 Gr. B Low-Alloy Steel 288°C 102 Type 304 SS 288°C Strain Range: 0.75%

Strain Range: 0.80%I Strain Rate: 0.004%/s . Strain Rate: 0.004%/s 101 101 II

(-3

_

E .T S10° , ( / " 10° "

S100 o- ' - 1 C) C High-Dissolved Oxygen Water 10"2 - - PWR

.-

10-2 --

- - ---- High-Dissolved Oxygen Water - - - High-DO Water SAir Air (Estimated)

Ii i i i i i r Ir l I , , ' , l I l l 100 1000 100 1000 Crack Depth (gm) Crack Depth (rim)

(a) (b)

Figure 2. Crack growth rates plotted as a function of crack depth for (a) A533-Gr B low-alloy steel and (b) Type 304 SS in air and LWR environments.

The fatigue a-N data for carbon and low-alloy steels in air and LWR environments have been i examined from the standpoint of fracture mechanics and CGR data. 67 ,68 Fatigue life is considered to consist of an initiation stage, composed of the growth of microstructurally small cracks, and a propagation stage, composed of the growth of mechanically small cracks. The growth of the latter has been characterized in terms of the J-integral range AJ and crack growth rate data in air and LWR environments.

I The estimated values show good agreement with the experimental &-N data for test specimens in air and water environments.

8

3. Fatigue Strain vs. Life Data The existing fatigue s_-N data developed at various establishments and research laboratories worldwide have been compiled by the Pressure Vessel Research Council (PVRC), Working Group on F-N Curve and Data Analysis. The database used in the ANL studies is an updated version of the PVRC database. A summary of the sources included in the updated PVRC database, as categorized by material type and test environment, is presented in Table 1.

Unless otherwise mentioned, smooth cylindrical gauge specimens were tested under strain control with a fully reversed loading, i.e., strain ratio of-1. Tests on notched specimens or at values of strain ratio other than -1 were excluded from the fatigue c-N data analysis. For the tests performed at ANL, the estimated uncertainty in the strain measurements is about 4% of the reported value. For the data obtained in other laboratories, the uncertainty in the reported values of strain is unlikely to be large enough to significantly affect the results.

In nearly all tests, fatigue life is defined as the number of cycles, N2 5, necessary for tensile stress to drop 25% from its peak or steady-state value. For the specimen size used in these studies, e.g., 5.1-9.5 mm (0.2-0.375 in.) diameter cylindrical specimens, this corresponds to a z3-mm-deep crack. Some of the earlier tests in air were carried out to complete failure of the specimen, and life in some tests is defined as the number of cycles for peak tensile stress to decrease by 1-5%. Also, in fatigue tests that were performed using tube specimens, life was represented by the number of cycles to develop a leak.

Table 1. Sources of the fatigue c-N data on reactor structural materials in air and water environments.

Source Material Environment Reference General Electric Co. Carbon steel, Type 304 SS Air and BWR water 8-11 Japan; including Ishikawajima- Carbon and low-alloy Air, BWR, and PWR JNUFAD* database, Harima Heavy Industries (IHI) steel, wrought and water 12-33 Co., Mitsubishi Heavy cast austenitic SS, Industries (MHI) Ltd., Hitachi Ni-Cr-Fe alloys Research Laboratory Argonne National Laboratory Carbon and low-alloy Air, BWR, and PWR 4-7, 34-40 steel, wrought and cast water austenitic SS Materials Engineering Carbon steel, austenitic SS Air and PWR water 41-43 Associates (MEA) Inc.

Germany; including MPA Carbon steel 44-45 France; including studies Austenitic SS Air and PWR water 69-71 sponsored by Electricite de France (EdF)

Jaske and O'Donnell Austenitic SS, Air 72 Ni-Cr-Fe alloys Others Austenitic SS, Air 73-78 Ni-Cr-Fe alloys Private communication from M. Higuchi, Ishikawajima-Hariina i-eavy Industries Co. Japan, to M. Prager of the Pressure Vessel Research Council, 1992. The old database "Fadal" has been revised and renamed "JNUFAD."

9

I I

For the tests where fatigue life was defined by a criterion other than 25% drop in peak tensile stress (e.g., 5% decrease in peak tensile stress or complete failure), fatigue lives were normalized to the 25%

drop values before performing the fatigue data analysis. 4 The estimated uncertainty in fatigue life determined by this procedure is about 2%.

An analysis of the existing fatigue E-N data and the procedures for incorporating environmental effects into the Code fatigue evaluations has been presented in several review articles 7 9 - 9 0 and ANL topical reports. 4 ,6,7,38-4 0 The key material, loading, and environmental parameters that influence the fatigue lives of carbon and low-alloy steels and austenitic stainless steels have been identified, and the I range of these key parameters where environmental effects are significant has been defined.

How various material, loading, and environmental parameters affect fatigue life and how these effects are incorporated into the ASME Code fatigue evaluations are discussed in detail for carbon and low-alloy steels, wrought and cast SSs, and Ni-Cr-Fe alloys in Sections 4, 5, and 6, respectively.

1 I

I I

I I

I I

I I

I 10 I

4 Carbon and Low-Alloy Steels The primary sources of relevant c-N data for carbon and low-alloy steels are the tests performed by General Electric Co. (GE) in a test loop at the Dresden 1 reactor; 8,9 work sponsored by EPRI at GE; 10,1 I the work of Terrell at Mechanical Engineering Associates (MEA); 4 1- 4 3 the work at ANL on fatigue of pressure vessel and piping steels; 4-7,3 4 4 0 the large JNUFAD database for "Fatigue Strength of Nuclear Plant Component" and studies at Ishikawajima-Harima Heavy Industries (IHI), Hitachi, and Mitsubishi Heavy Industries (MHI) in Japan; 12- 30 and the studies at Kraftwerk Union Laboratories (KWU) and Materialprufungsanstalt (MPA) in Germany. 44- 45 The database is composed of z1400 tests; z60% were obtained in the water environment and the remaining in air. Carbon steels include z12 heats of A333-Grade 6, A106-Grade B, A516-Grade 70, and A508-Class I steel, while the low-alloy steels include z16 heats of A533-Grade B, A302-Gr B, and A508-Class 2 and 3 steels.

4.1 Air Environment 4.1.1 Experimental Data In air, the fatigue lives of carbon and low-alloy steels depend on steel type, temperature, and for some compositions, applied strain rate and sulfide morphology. Fatigue c-N data from various investigations on carbon and low-alloy steels are shown in Fig. 3. The best-fit curves based on the ANL models (Eqs. 15 and 16 from Section 4.1.8) and the ASME Section III mean-data curves (at room temperature) are also included in the figures. The results indicate that, although significant scatter is apparent due to material variability, the fatigue lives of these steels are comparable at less than 5 x 105 cycles, and those of low-alloy steels are greater than carbon steels for >5 x 105 cycles. Also, the fatigue limit of low-alloy steels is higher than that of carbon steels.

Carbon Steel Low-Alloy Steel i Room Temperature Airi - -"Room Temperature Ai 1.0

41. CA333-6 1.0 0 Aaoa-3 I Best-Fit Air E

0.

Best-Fit Air ANL Model 0 --- - _ - ------SE fMea Cd ANILModel Cu:`rv U) 0.1 0.1 ASME Code<

MeanCarve 102 103 104 105 106 107 108 102 103 104 105 106 107 108 Fatigue Life (Cycles) Fatigue Life (Cycles)

Figure 3. Fatigue strain vs. life data for carbon and low-alloy steels in air at room temperature (JNUFAD database and Refs. 4,12,13,41).

The existing fatigue c_-N data for low-alloy steels are in good agreement with the ASME mean data curve. The existing data for carbon steels are consistent with the ASME mean data curve for fatigue life

S5 x 105 cycles and are above the mean curve at longer lives. Thus, above 5 x 105 cycles, the Code mean curve is conservative with respect to the existing fatigue c-N data.

I1

I 9 The current Code mean data curves are either consistent with the existing fatigue E-N data or are somewhat conservative under some conditions. g 4.1.2 Temperature In air, the fatigue life of both carbon and low-alloy steels decreases with increasing temperature; however, the effect is relatively small (less than a factor of 1.5). Fatigue a-N data from the JNUFAD database and other investigations in air at 286-300'C are shown in Fig. 4. For each grade of steel, the data represent several heats of material. The best-fit curves for carbon and low-alloy steels at room temperature (Eqs. 15 and 16 from Section 4.1.8) and at 289°C (Eqs. 13 and 14 from Section 4.1.8) are I

also included in the figures. The results indicate a factor of z1.5 decrease in fatigue life of both carbon and low-alloy steels as the temperature is increased from room temperature to 300°C. As discussed later in Section 4.1.7, the greater-than-predicted difference between the best-fit air curve at room temperature and the data for A106-Gr B steel at 289°C is due to heat-to-heat variability and not temperature effects.

& The effect of temperature is not explicitly considered in the mean data curve used for obtaining the I fatigue design curves; variations in fatigue life due to temperature are accountedfor in the subfactorfor "datascatter and material variability."

" Carbon

. Steel i I I Low-Alloy Steel 288-300C Air' 286-300°C Air Best-Fit Air Bet-i Air.

ANL Model II \\ / ANL Model Room Temperature Room Temperature C) A333 i 0 A508-3

._

1 0 l(*

. .. I qi Ii I

Best Fit Ai

• o',-Fi C r Si. -* .... i Beatq_FitoO~

Air,

,-N Moe /. -.-. 01 ._

• ANL- o9 deI!

01 NL 28est . .. 01o-del _ _ _ _ _ _ 1_

3 4 5 7 102 10 10 10 1 06 10 108 102 103 104 105 106 107 108 Fatigue Life (Cycles) Fatigue Life (Cycles)

Figure 4. Fatigue strain vs. life data for carbon and low-alloy steels in air at 288°C (JNUFAD database, and Refs. 4,12,13,42,43).

4.1.3 Strain Rate I

The effect of strain rate on the fatigue life of carbon and low-alloy steels in air appears to depend on the material composition. The existing data indicate that in the temperature range of dynamic strain aging (200-370'C), some heats of carbon and low-alloy steel are sensitive to strain rate; with decreasing strain rate, the fatigue life in air may be either unaffected, 4 decrease for some heats, 9 1 or increase for others. 92 The C and N contents in the steel are considered to be important. Inhomogeneous plastic deformation can result in localized plastic strains. This localization retards blunting of propagating cracks I

9 that is usually expected when plastic deformation occurs and can result in higher crack growth rates. 1 The increases in fatigue life have been attributed to retardation of CGRs due to crack branching and 92 suppression of the plastic zone. Formation of cracks is easy in the presence of dynamic strain aging.

• Variations in fatigue life due to the effects of strain rate are not explicitly considered in the fatigue design curves, they are accountedfor in the subfactorfor "data scatter and material variability."

12

1100 1100 - I"! ' ''" ' ' " ' ' "

A106-Gr B Carbon Steel A533-Gr B Low-Alloy Steel i

-- A~t =0.775 %-

5t o/07 1000 A %* 0.7 t '" I..*+: 1000 CL 900 to 0~ 900 0 D 0

-I o 0

< 800 ouu e5 0t 0O 700 A'A~ Co 0* 700 Stra__Ri ______

Strain Rate %Is) at 600 Open Sym~bp s 0~4A 600 Open S" ymuors*

tO Closed Symaols: 0.004 Al U) Closed Symbols: 0.004 500 . R. a A... ...... ....... ... 500 288 C Air _

L+ Ro om-Temperature Air' A Room Temperature Air

'+u 4UU 40U 100 lot 102 101 102 103 104 Number of Cycles Number of Cycles Figure 5. Effect of strain rate and temperature on cyclic stress of carbon and low-alloy steels.

4.1.4 Sulfide Morphology Some high-S steels exhibit very poor fatigue properties in certain orientations because of structural factors such as the distribution and morphology of sulfides in the steel. For example, fatigue tests on a high-S heat of A302-Gr. B steel in three orientations* in air at 288°C indicate that the fatigue life and fatigue limit in the T2 orientation are lower than those in the R and TI orientations. 4 At low strain rates, fatigue life in the T2 orientation is nearly one order of magnitude lower than in the R orientation. In the orientation with poor fatigue resistance, crack propagation is preferentially along the sulfide stringers and is facilitated by sulfide cracking.

• Variations in fatigue life due to differences in sulfide morphology are accountedfor in the subfactor for "datascatter and material variability."

4.1.5 Cyclic Strain Hardening Behavior The cyclic stress-strain response of carbon and low-alloy steels varies with steel type, temperature, and strain rate. In general, these steels show initial cyclic hardening, followed by cyclic softening or a saturation stage at all strain rates. The carbon steels, with a pearlite and ferrite structure and low yield stress, exhibit significant initial hardening. The low-alloy steels, with a tempered bainite and ferrite structure and a relatively high yield stress, show little or no initial hardening and may exhibit cyclic softening with continued cycling. For both steels, maximum stress increases as applied strain increases and generally decreases as temperature increases. However, at 200-370'C, these steels exhibit dynamic strain aging, which results in enhanced cyclic hardening, a secondary hardening stage, and negative strain rate sensitivity. 9 1,9 2 The temperature range and extent of dynamic strain aging vary with composition and structure.

The effect of strain rate and temperature on the cyclic stress response of AI 06-Gr B carbon steel and A533-Gr B low-alloy steel is shown in Fig. 5. For both steels, cyclic stresses are higher at 288°C than at room temperature. At 288°C, all steels exhibit greater cyclic and secondary hardening because of dynamic strain aging. The extent of hardening increases as the applied strain rate decreases.

Both transverse (T) and radial (R) directions are perpendicular to the rolling direction, but the fracture plane is across the thickness of the plate in the transverse orientation and parallel to the plate surface in the radial orientation.

13

I The cyclic strain hardening behavior is likely to influence the fatigue limit of the material; variations in fatigue life due to the effects of strain hardening are not explicitly considered in the fatigue design curves, they are accountedfor in the subfactorfor "datascatter and materialvariability."

4.1.6 Surface Finish 3 The effect of surface finish must be considered to account for the difference in fatigue life expected in an actual component with industrial-grade surface finish, compared with the smooth polished surface of a test specimen. Fatigue life is sensitive to surface finish; cracks can initiate at surface irregularities that are normal to the stress axis. The height, spacing, shape, and distribution of surface irregularities are I

important for crack initiation. The most common measure of roughness is average surface roughness Ra, which is a measure of the height of the irregularities. Investigations of the effects of surface roughness on the low-cycle fatigue of Type 304 SS in air at 593°C indicate that fatigue life decreases as surface roughness increases. 9 3,94 The effect of roughness on crack initiation Ni(R) is given by Ni(Rq) = 1012 Rq 021

, (12) i where the root-mean-square (RMS) value of surface roughness Rq is in pim. Typical values of Ra for surfaces finished by different metalworking processes in the automotive industry 95 indicate that an Ra of 3 jim (or an Rq of 4 jim) represents the maximum surface roughness for drawing/extrusion, grinding, honing, and polishing processes and a mean value for the roughness range for milling or turning processes. For carbon steel or low-alloy steel, an Rq of 4 ltm in Eq.9312 (the Rq of a smooth polished specimen is z0.0075 jim) would decrease fatigue life by a factor of z3.

I Fatigue test has been conducted on a A106-Gr B carbon steel specimen that was intentionally roughened in a lathe, under controlled conditions, with 50-grit sandpaper to produce circumferential scratches with an average roughness of 1.2 [Lm and an Rq of 1.6 jim (z62 micro in.). 39 The results for I

smooth and roughened specimens are shown in Fig. 6. In air, the fatigue life of a roughened A106-Gr B specimen is a factor of z3 lower than that of smooth specimens. Another study of the effect of surface finish on the fatigue life of carbon steel in room-temperature air showed a factor of 2 decrease in life when Ra was increased from 0.3 to 5.3 gm.9 6 These results are consistent with Eq. 12. Thus, a factor of 2-3 on cycles may be used to account for the effects of surface finish on the fatigue life of carbon and low-alloy steels.

I

~i At06 Gr B Carbon Steel O Air. 0.0041,.s 289'C A Air., 0,01 1.0

-.Figure 6.

it u-*,,*,,

'- Effect of surface finish on the fatigue life of E-.. Al 06-Gr B carbon steel in air at 289eC.

< ASME Code A -

c Design Curve U) 0.1 Open Symbols: Smooth Specimens Closed Symbols: Rough Surface, 50 grt paper 3

102 10 104 105 106 Fatigue Life (Cycles) 14

• The effect of surface finish was not investigated in the mean data curve used to develop the Code fatigue design curves; it is included as part of the subfactor that is applied to the mean data curve to accountfor "surfacefinish and environment."

4.1.7 Heat-to-Heat Variability Several factors, such as small differences in the material composition and structure, can change the tensile and fatigue properties of the material. The effect of interstitial element content on dynamic strain aging and the effect of sulfide morphology on fatigue life have been discussed in Sections 4.1.3 and 4.1.4, respectively. The effect of tensile strength on the fatigue life has been included in the expression for the mean data curve described in the Section III criteria document, i.e., constant Af in Eq. 2. Also, the fatigue limit of a material has been correlated with its tensile strength, e.g., the fatigue limit increases with 97 increasing tensile yield stress.

The effects of material variability and data scatter must be included to ensure that the design curves not only describe the available test data well, but also adequately describe the fatigue lives of the much larger number of heats of material that are found in the field. The effects of material variability and data scatter are often evaluated by comparing the experimental data to a specific model for fatigue crack initiation, e.g., the best-fit (in some sense) to the data. The adequacy of the evaluation will then depend on the sample of data used in the analysis. For example, if most of the data have been obtained from a heat of material that has poor resistance to fatigue damage or under loading conditions that show significant environmental effects, the results may be conservative for most of the materials or service conditions of interest. Conversely, if most data are from a heat of material with a high resistance to fatigue damage, the results could be nonconservative for many heats in service.

Another method to assess the effect of material variability and data scatter is by considering the best-fit curves determined from tests on individual heats of materials or loading conditions as samples of the much larger population of heats of materials and service conditions of interest. The fatigue behavior of each of the heats or loading conditions is characterized by the value of the constant A in Eq. 6. The values of A for the various data sets are ordered, and median ranks are used to estimate the cumulative distribution of A for the population. 9 8,99 The distributions were fit to lognormal curves. No rigorous statistical evaluation was performed, but the fits seem reasonable and describe the observed variability adequately. Results for carbon and low-alloy steels in air are shown in Fig. 7. The data were normalized to room-temperature values using Eqs. 13 and 14 (section 4.1.8). The median value of the constant A is 6.583 and 6.449, respectively, for the fatigue life of carbon steels and low-alloy steels in room-temperature air. Note that the two heats of A 106-Gr B carbon steel are in the 10-25 percentile of the data, i.e., the fatigue lives of these heats are much lower than the average value for carbon steels.

The A values that describe the 5th percentile of these distributions give fatigue c-N curves that are expected to bound the fatigue lives of 95% of the heats of the material. The cumulative distributions in Fig. 7 contain two potential sources of error. The mean and standard deviation of the population must be estimated from the mean and standard deviation of the sample, 10 0 and confidence bounds can then be obtained on the population mean and standard deviation in terms of the sample mean and standard deviation. Secondly, even this condition does not fully address the uncertainty in the distribution because of the large uncertainties in the sample values themselves, i.e., the "horizontal" uncertainty in the actual value of A for a heat of material, as indicated by the error bars in Fig. 7. A Monte Carlo analysis was performed to address both sources of uncertainty. The results for the median value and standard deviation of the constant A from the Monte Carlo analysis did not differ significantly from those determined directly from the experimental values.

15

I 1 .0 . . . . . . . .. . . . . I .

' ' 'J. .. ... 1 .0 . . . . . i . . . . . -% , -l . . . . .

-:Carbon Steel-.. . ___j .... .Low-Alloy Steel---

Air -Air

0. . 75th Percentile .........

. ....r . .. ......

.. -r. ...... . * . = .. . .* ............... I. ..... ... 0.8 75th Percentile o_ I / ,I 0.6 -!0.6

.t Median 6.583 CI. ~ ~.. ~ ~~~~~~.. i .. . ..... ,-.

....... ..... I.

r-li Median 6.449 .... *:- 358 19 Data Points Heats -

-- [ 153 Data Points I - - ->

5 0.4 .7" 1 8 Heats . V A302-B

. . . - --- E ......L ( A533-8 A

.I.. .... .. " A5 3-B ,l,

.. . 0 0 AS33- 51

.......

2 t ....

e51hn le  : .! . . .......

.......

................... *i

• l~6 *,. *. Percen ile ~

..... *v A 0 - 1 P02l

....

.3.(" 0.2 A508-3(1)

......

J .. .. Ll......l......... A, A333-6 (2) 0< A508- (7) 3 0 A 333-6( ) A508 0M) 556 .5 7A A333-6(5 A333-6 (7) A 15MnN63 0.000 0 5 5.5 6 6.5 7 7.5 8 8.5 5 5.5 6 6.5 7 7.5 8 8.5 Constant A (a)

Constant A (b)

I Figure 7. Estimated cumulative distribution of constant A in the ANL models for fatigue life for heats of (a) carbon steels and (b) low-alloy steels in air.

The results for carbon and low-alloy steels are summarized in Tables 2 and 3, respectively, in terms of values for A that provide bounds for the portion of the population and the confidence that is desired in the estimates of the bounds. In air, the 5th percentile value of Parameter A at a 95% confidence level is 5.559 for carbon steels and 5.689 for low-alloy steels. From Fig. 7, the median value of A for the sample is 6.583 for carbon steels and 6.449 for low-alloy steels. Thus, the 95/95 value of the margin to account for material variability and data scatter is 2.8 and 2.1 on life for carbon steels and low-alloy steels, respectively. These margins are needed to provide 95% confidence that the resultant life will be greater than that observed for 95% of the materials of interest. The margin is higher for carbon steels because the analysis is based on a smaller number of data sets, i.e., 19 for carbon steels and 32 for low-alloy steels.

• The mean data curve used to develop the Code fatigue design curves represents the average behavior; heat-to-heat variability is included in the subfactor that is applied to the mean data curve to accountfor "datascatter and material variability."

Table 2. Values of parameter A in the ANL fatigue life model for carbon steels in air and the margins on life as a function of confidence level and percentage of population bounded. I Confidence Percentage of Population Bounded (Percentile Distribution of A)

Level 50 95(5) 5.798 90(10) 5.971 75(25)

Values of Parameter A 6.261 67(33) 6.373 50(50) 6583 I

75 95 5.700 5.559 5.883 5.756 6.183 6.069 Margins on Life 6.295 6.183 6.500 6.381 I

50 75 95 2.2 2.4 2.8 1.8 2.0 2.3 1.4 1.5 1.7 1.2 1.3 1.5 1.0 1.1 1.2 I

16 I I

Table 3. Values of parameter A in the ANL fatigue life model for low-alloy steels in air and the margins on life as a function of confidence level and percentage of population bounded.

Confidence Percentage of Population Bounded (Percentile Distribution of A)

Level 95 (5) 90(10) 75(25) 67(33) 50 (50)

Values of Parameter A 50 5.832 5.968 6.196 6.284 6.449 75 5.774 5.916 6.150 6.239 6.403 95 5.689 5.840 6.085 6.175 6.337 Margins on Life 50 1.9 1.6 1.3 1.2 1.0 75 2.0 1.7 1.3 1.2 1.0 95 2.1 1.8 1.4 1.3 1.1 4.1.8 Fatigue Life Model Fatigue life models for estimating the fatigue lives of these steels in air based on the existing fatigue c-N data have been developed at ANL as best-fits of a Langer curve to the data. 4 ,3 9 The fatigue life, N, of carbon steels is represented by In(N) = 6.614 - 0.00124 T - 1.975 ln(Ea- 0.113), (13) and that of low-alloy steels, by ln(N) = 6.480 - 0.00124 T - 1.808 ln(Ea- 0.151), (14) where Fa is applied strain amplitude (%), and T is the test temperature (°C). Thus, in room-tetnperature air, the fatigue life of carbon steels is expressed as In(N) = 6.583 - 1.975 ln(aa - 0.113), (15) and that of low-alloy steels, by In(N) = 6.449- 1.808 ln(Ea- 0.151). (16)

Note that these equations have been updated based on the analysis presented in Section 4.1.7; constant A in the equations is different from the value reported earlier in NUREG/CR-6583 and 6815.

Relative to the earlier model, the fatigue lives predicted by the updated model are z2% higher for carbon steel and -16% lower for low-alloy steels. The experimental values of fatigue life and those predicted by Eqs. 15 and 16 for carbon and low-alloy steels in air are plotted in Fig. 8. The predicted fatigue lives show good agreement with the experimental values; the experimental and predicted values are within a factor of 3.

9 The fatigue life models represent mean values offatigue life of specimens tested under fully reversed strain-controlled loading. The effects of parameters (such as mean stress, surface finish, size and geometry, and loading history) that are known to influence fatigue life are not explicitly considered in the model; such effects are accountedfor in the several subfactors that are applied to the mean data curve to obtain the Codefatigue design curve.

17

I CarbonTemperature

--Room Steels ----

---

-- , .* .. Carbon 100-290°ClSteelsrAr................. .. "- ;i....,

0e

V

.j o1

..

C o

¢ zý o.- 6 0~ a 10 ..... . 101 .I......

02---- - - -- 1024 6 .G . .

,, " !0 A333-GG. '  ; - 0 A333-Gr 5 6 5 101 to0 102 103 104 1o 10 10t101 102 103 104 10 106 Observed Life (Cycles) Observed Life (Cycles)

(a) (b)I Low-Alloy Steels Low-Alloy Steels (b)

Room Temperature Air 50-35 °C Air 6 0

• ~~~0 -.7--..----[ 104 ' .- 7.."

S--

0~0-

,-- - .

I t1o, 10 1 1 103 10 I , . . .

-? .'a I 0 0! I 102 .... % *:- ';. . .. a;-

? *552 . * ... 102 I'.. . . 0 A500-2 t le 1 10' I

4 lot 102 102 to 105 10e 10' 102 102 i04 10 106 Observed Life (Cycles) Observed Life (Cycles)

(c) (d)

Figure 8. Experimental and predicted fatigue lives of (a, b) carbon steels and (c, d) low-alloy steels in air.

4.1.9 Extension of the Best-Fit Mean Curve from 106 to 1011 Cycles The experimental fatigue c-N curves that were used to develop the current Code fatigue design I curve for carbon and low-alloy steels were based on low-cycle fatigue data (less than 2 x 105 cycles). The design curves proposed in this report are developed from a larger database that includes fatigue lives up to 108 cycles. Both the ASME mean curves and the ANL models in this report use the modified Langer equation to express the best-fit mean curves and are not recommended for estimating lives beyond the range of the experimental data, i.e., in the high-cycle fatigue regime.

I An extension of the current high-cycle fatigue design curves in Section II and Section VIII, I Division 2, of the ASME Code for carbon and low-alloy steels from 106 to 101 cycles has been proposed by W. J. O'Donnell for the ASME Subgroup on Fatigue Strength.! In the high-cycle regime, at temperatures not exceeding 371 'C (700'F), the stress amplitude vs. life relationship is expressed as 3 Sa = EFa = CIN- 0 0 5, (17)

• W. J. O'Donnell, "Proposed Extension of ASME Code Fatigue Design Curves for Carbon and Low-Alloy Steels from Temperatures not Exceeding 700°F," presented to ASME Subgroup on Fatigue Strength December 4, 1996.

10 6

to l10l Cycles for I

-18

where Ea is applied strain amplitude, E is the elastic modulus, N is the fatigue life, and C1 is a constant. A fatigue life exponent of -0.05 was selected based on the fatigue stress range vs. fatigue life data on plain plates, notched plates, and typical welded structures given in Welding Research Council (WRC) Bulletin 398.101 Because these data were obtained from load-controlled tests with a load ratio R = 0, they take into account the effect of maximum mean stresses and, may over estimate the effect of mean stress under strain-controlled loading conditions. Also, the fatigue data presented in Bulletin 398 extend only up to 5 x 106 cycles; extrapolation of the results to 1011 cycles using a fatigue life exponent of -0.05 may yield conservative estimates of fatigue life.

Manjoine and Johnson 9 7 have developed fatigue design curves up to 1011 cycles for carbon steels and austenitic SSs from inelastic and elastic strain relationships, which can be correlated with ultimate tensile strength. The log-log plots of the elastic strain amplitudes vs. fatigue life data are represented by a bilinear curve. In the high-cycle regime, the elastic-strain-vs.-life curve has a small negative slope instead of a fatigue limit. 9 7 For carbon steel data at room temperature and 371°C and fatigue lives extending up to 4 x 107 cycles, Manjoine and Johnson obtained an exponent of -0.01. The fatigue E-N data from the present study at room temperature and with fatigue lives up to 108 cycles yield a fatigue life exponent of approximately -0.007 for both carbon and low-alloy steels. Because the data are limited, the more conservative value obtained by Manjoine and Johnson 97 is used. Thus, in the high-cycle regime, the applied stress amplitude is given by the relationship S, = EEa = C 2 N- 0 0 1 . (18)

The high-cycle curve (i.e., Eq. 18) can be used to extend the best-fit mean curves beyond 106 cycles; the mean curves will exhibit a small negative slope instead of the fatigue limit predicted in the modified Langer equation. The constant C 2 is determined from the value of strain amplitude at 108 cycles obtained from Eq. 15 for carbon steels and from Eq. 16 for low-alloy steels.

4.1.10 Fatigue Design Curve Although the two mean curves for carbon and low-alloy steels (i.e., Eqs. 7 and 9) are significantly different, because the mean stress correction is much larger for the low-alloy steels, the differences between the curves is much smaller when mean stress corrections are considered. Thus, the ASME Code provides a common curve for both carbon and low-alloy steels. Fatigue design curves for carbon steels and low-alloy steels based on the ANL fatigue life models can be obtained from Eqs. 15 and 18, and Eqs.

16 and 18, respectively.

The best-fit curves are first corrected for mean stress effects by using the modified Goodman relationship, and the mean-stress adjusted curve is reduced by a factor of 2 on stress or 12 on cycles, whichever is more conservative. The discussions presented later in Section 7.5 indicate that the current Code requirement of a factor of 20 on cycles, to account for the effects of material variability and data scatter, specimen size, surface finish, and loading history, is conservative by at least a factor of 1.7. Thus, to reduce this conservatism, fatigue design curves based on the ANL model for carbon and low-alloy steels have been developed using factors of 12 on life and 2 on stress. These design curves are shown in Figs. 9 and 10, respectively. The current Code design curve for carbon and low-alloy steels with ultimate tensile strength (UTS) <552 MPa (<80 ksi) and the extension of the design curve to 1011 cycles proposed by W. J. O'Donnell are also included in the figures. The values of stress amplitude (Sa) vs. cycles for the ASME Code curve with O'Donnell's extension, and the design curve based on the updated ANL fatigue life model (i.e., Eqs. 15 and 18 for carbon steel and, 16 and 18 for low-alloy steel) are listed in Table 4.

19

m

• For low-alloy steels, the current Code fatigue design curve for carbon and low-alloy steels with I

ultimate tensile strength <552 MPa (<80 ksi) is either consistent. or conservative with respect to the existing fatigue E-N data. Also, discussions presented in Section 7.5 indicate that the current Code requirement of a factor of 20 on life is conservative by at least a factor of 1.7. Fatigue design curves have been developedfrom the ANL model usingfactors of 12 on life and 2 on stress.

Carbon Steels I UTS :552 MPa (-*80 ksi)

Air up to 371'C (7007F) 103 E = 206.8 Ga Figure 9.

..... -ASME Code Curve

. ANL Model& Eq. 17 Fatigue design curve for U)

____ANL Model & Eq 18 carbon steels in air. The E curve developed from the ANL model is based on I 102 factors of 12 on life and 2 Carbon Steels u = 551.6 VMPa oy = 275,8 MPa

.z"T lp7

.

on stress.

m 101 102 10 3

104 105 106 10 Number of Cycles N 7

108 log 1010 1011 m

IL 103 Low-Alloy Steels UTS <552 MPa (-<80 ksi)

Air up to 371'C (700TF)

E:= 206.8 GPa ASMOE Code Curve Figure 10.

I I

IANL ModeI & Eq.1 Fatigue design curve for CL

___.... ANL Model & Eq 18 low-alloy steels in air. The curve developed from the

~0 ANL model is based on 2

UO 102 Low-Alloy Steels o = 689.5 MPa oy = 482.6 MPa i i I factors of 12 on life and 2 on stress. I

.101

._......

102 103 104

... JI 10 5

I 106 I

107 Number of Cycles N I ..

108 I

109 1010 1011 I Table 4. Fatigue design curves for carbon and low-alloy steels and proposed extension to 1011 cycles.

Stress Amplitude (MPa/ksi) Stress Amplitude (MPa/ksi) m Cycles IE+O1 2 E+01 ASME Code Curve 3999 (580) 2827 (410)

Eqs. 15 & 18 5355 (777) 3830 (556)

Eqs. 16 & 18 Carbon Steel Low-Alloy Steel 5467 (793) 3880 (563)

Cycles 2 E+05 5 E+05 ASME Code Curve 114(16.5) 93 (13.5)

Eqs. 15 & 18 176 (25.5) 154 (22.3)

Eqs. 16 & 18 Carbon Steel Low-Alloy Steel 141 (20.5) 116(16.8)

I 5 E+01 1 E+02 2E+02 5 E+02 1896 (275) 1413 (205) 1069(155) 724 (105) 2510 (364) 1820 (264) 1355(197) 935 (136) 2438 (354) 1760 (255) 1300(189) 900 (131) 1 E+06 2 E+06 5E+06 1 E+07 86(12.5) 76.5 (11.1) 142 (20.6) 130 (18.9) 120(17.4) 115(16.7) 106 (15.4) 98(14.2) 94(13.6) 91 (13.2) m 1 E+03 2 E+03 5 E+03 1 E+04 572 (83) 441 (64) 331 (48) 262 (38) 733 (106) 584 (84.7) 451 (65.4) 373 (54.1) 720 (104) 576 (83.5) 432 (62.7) 342 (49.6) 2 E+07 5 E+07 1 E+08 1 E+09 68.3 (9.9) 60.7 (8.8) 110(16.0) 107 (15.5) 105 (15.2) 102 (14.8) 90(13.1) 88(12.8) 87(12.6) 83 (12.0) m I

2 E+04 214(31) 305(44.2) 276(40.0) 1 E+010 54.5(7.9) 97(14.1) 80(11.6) 5 E+04 159(23) 238(34.5) 210(30.5) 1 E+01 1 48.3 (7.0) 94(13.6) 77(11.2) 1 E+05 138 (20.0) 201 (29.2) 172 (24.9) 20

4.2 LWR Environments 4.2.1 Experimental Data Fatigue a-N data on carbon and low-alloy steels in air and high-DO water at 288°C are shown in Fig. 11. The curves based on the ANL models (Eqs. 20 and 21 in Section 4.2.12) are also included in the figures. The fatigue data in LWR environments indicate a significant decrease in fatigue life of carbon and low-alloy steels when four key threshold conditions are satisfied simultaneously, viz., applied strain range, service temperature, and DO in the water are above a minimum threshold level, and the loading strain rate is below a threshold value. The S content of the steel is also an important parameter for environmental effects on fatigue life. Although the microstructures and cyclic-hardening behavior of carbon steels and low-alloy steels are significantly different, environmental degradation of fatigue life of these steels is identical. For both steels, environmental effects on fatigue life are moderate (i.e., it is a factor ofz2 lower) if any one of the key threshold conditions is not satisfied.

A533-Gr. B Low-Alloy Steel Strain Rate (%/s) Al 06-Gr. B Low-Alloy Stool Strain Rate (%/s) 288°C 40.4 288°C .- 0 4;0A 0,.004/0.4 0, 0.004'0 4

- ANL Model 17 0.4/0004 - I ANLModel V 0.4/0 004 1.0 _* ..... * * * ] i - 288°C-... .Air.  ; - ......................... . . ... 1 .0

-

. ...---

-.--- *-..- - - ,. . .. .. ... . .

"

i 288CAir ANLModel A

. .,o -..

4) ANLModel Ia _I 289°C High-DO Water

. 289°C High-DO Water

-- 0 004%/s Strain Rate - .., .0004 %Is Strain Rate

"-*i-.. e0.*

_, I- - i. I E V- E

• ,b->- ~ d-,>e 0 .1 ......... 0 .1 Open Symbols: Air Open Symbols: I Closed Symbols: High DO Water Closed Symbols: High DO Water t I . . .. I. I....II - I . .I 3 4 3 5 102 10 10 105 106 107 102 10 104 10 106 107 Fatigue Life (Cycles) Fatigue Life (Cycles)

(a) (b)

Figure 11. Strain amplitude vs. fatigue life data for (a) A533-Gr B and (b) A106-Gr B steels in air and high-dissolved-oxygen water at 2880C (Ref. 4).

The existing fatigue data indicate that a slow strain rate applied during the tensile-loading cycle is primarily responsible for environmentally assisted reduction in- fatigue life of these steels. 4 The mechanism of environmentally assisted reduction in fatigue life of carbon and low-alloy steels has been termed strain-induced corrosion cracking (SICC). 4 8 ,55 ,5 6 A slow strain rate applied during both the tensile-load and compressive-load portion of the cycle (i.e., slow/slow strain rate test) does not further decrease the fatigue life, e.g., see solid diamonds and square in Fig. I lb for A 106-Gr 3 carbon steel.

Limited data from fast/slow tests indicate that a slow strain rate during the compressive load cycle also decreases fatigue life. However, the decrease in life is relatively small; for fast/slow strain rate tests, the major contribution of environment most likely occurs during slow compressive loading near peak tensile load. For example, the fatigue life of A533-Gr B low-alloy steel at 288°C, 0.7 ppm DO, and Z0.5%

strain range decreased by factors of 5, 8, and 35 for the fast/fast, fast/slow, and slow/fast tests, respectively, i.e., see solid circles, diamonds, and inverted triangles in Fig. 1 la. Similar results have been observed for A333-Gr 6 carbon steel; 1 7 relative to the fast/fast test, fatigue life for slow/fast and fast/slow tests at 288°C, 8 ppm DO, and 1.2% strain range decreased by factors of 7.4 and 3.4, respectively.

The environmental effects on the fatigue life of carbon and low-alloy steels are consistent with the slip oxidation/dissolution mechanism for crack propagation. 102 , 103 A critical concentration of sulfide 21

(S2 -) or hydrosulfide (HS-) ions, which is produced by the dissolution of sulfide inclusions in the steel, is required at the crack tip for environmental effects to occur. The requirements of this mechanism are that a protective oxide film is thermodynamically stable to ensure that the crack will propagate with a high aspect ratio without degrading into a blunt pit, and that a strain increment occurs to rupture that oxide film and thereby expose the underlying matrix to the environment. Once the passive oxide film is ruptured, crack extension is controlled by dissolution of freshly exposed surface and by the oxidation characteristics. The effect of the environment increases with decreasing strain rate. The mechanism assumes that environmental effects do not occur during the compressive load cycle, because during that period water does not have access to the crack tip.

A model for the initiation or cessation of environmentally assisted cracking (EAC) of these steels in low-DO PWR environments has also been proposed.1 04 Initiation of EAC requires a critical concentration of sulfide ions at the crack tip, which is supplied with the sulfide ions as the advancing crack intersects the sulfide inclusions, and the inclusions dissolve in the high-temperature water. Sulfide ions are removed from the crack tip by one or more of the following processes: (a) diffusion due to the concentration gradient, (b) ion transport due to differences in the electrochemical potential (ECP), and (c) fluid flow induced within the crack due to flow of coolant outside the crack. Thus, environmentally enhanced CGRs are controlled by the synergistic effects of S content, environmental conditions, and flow rate. The EAC initiation/cessation model has been used to determine the minimum crack extension and CGRs that are required to maintain the critical sulfide ion concentration at the crack tip and sustained environmental enhancement of growth rates.

• A L WR environment has a significant effect on the fatigue life of carbon and low-alloy steels; such effects are not considered in the current Code design curve. Environmental effects may be incorporated into the Code fatigue evaluation using the Fen approach describedin Section 4.2.13.

4.2.2 Strain Rate The effects of strain rate on fatigue life of carbon and low-alloy steels in LWR environments are significant when other key threshold conditions, e.g., strain amplitude, temperature, and DO content, are satisfied. When any one of the threshold conditions is not satisfied, e.g., low-DO PWR environment or temperature <150'C, the effects of strain rate are consistent with those observed in air.

When all thresholdconditions are satisfied, the fatigue life of carbon and low-alloy steels decreases logarithmically with decreasing strain rate below 1%/s. The fatigue lives of A106-Gr B and A333-Gr 6 carbon steels and A533-Gr B low-alloy steel 4, 17 are plotted as a function of strain rate in Fig. 12. Only a moderate decrease in fatigue life is observed in simulated (low-DO) PWR water, e.g., at DO levels of

<0.05 ppm. For the heats of A 106-Gr B carbon steel and A533-Gr B low-alloy steel, the effect of strain rate on fatigue life saturates at zO.001%/s strain rate. Although the data for A333-Gr 6 carbon steel at 250'C and 8 ppm DO do not show an apparent saturation at zO.001%/s strain rate, the results are comparable to those for the other two steels.

In L WR environments, the effect of strain rate on the fatigue life of carbon and low-alloy steels is explicitly considered in Fen given in Eqs. 27 and 28 (Section 4.2.13). Also, guidance is provided for defining the strainratefor a specific stress cycle or load set pair.

I 22!

A106-Gr B Carbon Steel A333-Gr 6 Carbon Steel 104 288 °C. E! =0.4% 104 250 °C, Eý=0.6%

ýp 103 gj"103 A x

a) a)

LL U_

102 . .. r....-. 102 - A 0Simulated P0,,R 0 Samulated PWR

[ i Z* 0.7 ppm DO z* 8 ppm DO 10-5 10-4 10-3 10-2 10-1 100 10-5 10-4 10-3 10-2 10-1 100 1 .-' . . . . . .. ! . . . . . . .. 102......

Strain Rate (%/s) Strain Rate (%/s)

(a) (b)

A533-Gr B Low-Alloy Steel 104 288°C, =0.4%. .

U) 0

... Figure 12.

3

.2 10 ,.-.*Dependence of fatigue life of carbon and low-alloy a) steels on strain rate (Refs. 4, 17).

LL, 102 .... ... Ar C sndwU~d P/P/n A 60.7ppm DO

.I . . ... . . . ... t . . .. .. i ,. . . ... .I , , ,,

3 10-5 10-4 10- 10-2 10-1 100 Strain Rate (%/s)

(c) 4.2.3 Strain Amplitude A minimum threshold strain range is required for environmentally assisted decrease in fatigue life, i.e., the LWR coolant environments have no effect on the fatigue life of these steels at strain ranges below the threshold value. The fatigue lives of A533-Gr B and A 106-Gr B steels in high-DO water at 288°C and various strain rates 4 are shown in Fig. 11. Fatigue tests at low strain amplitudes are rather limited.

Because environmental effects on fatigue life increase with decreasing strain rate, fatigue tests at low strain amplitudes and strain rates that would result in significant environmental effects are restrictively time consuming. For the limited data that are available, the threshold strain amplitude (one-half the threshold strain range) appears to be slightly above the fatigue limit of these steels.

Exploratory fatigue tests with changing strain rate have been conducted to determine the threshold strain range beyond which environmental effects are significant during a fatigue cycle. The tests are performed with waveforms in which the slow strain rate is applied during only a fraction of the tensile loading cycle. 4 ,18 The results for AI06-Gr B steel tested in air and low- and high-DO environments at 288°C and z0.78% strain range are summarized in Fig. 13. The waveforms consist of segments of loading and unloading at fast and slow strain rates. The variation in fatigue life of two heats of carbon steel and one heat of low-alloy steel 4, 18 is plotted as a function of the fraction of loading strain at slow strain rate in Fig. 14. Open symbols indicate tests where the slow portions occurred near the maximum tensile strain, and closed symbols indicate tests where the slow portions occurred near the maximum compressive strain. In Fig. 14, if the relative damage was the same at all strain levels, fatigue life should decrease linearly from A to C along the chain-dot line. Instead, the results indicate that during a strain 23

I cycle, the relative damage due to slow strain rate occurs only after the strain level exceeds a threshold value. The threshold strain range for these steels is 0.32-0.36%.

Loading histories with slow strain rate applied near the maximum tensile strain (i.e., waveforms C, D, E, or F in Fig. 13) show continuous decreases in life (line AB in Fig. 14) and then saturation when a portion of the slow strain rate occurs at strain levels below the threshold value (line BC in Fig. 14). In contrast, loading histories with slow strain rate applied near maximum compressive strain (i.e., waveforms G, H, or I in Fig. 13) produce no damage (line AD in Fig. 14a) until the fraction of the strain is sufficiently large that slow strain rates are occurring for strain levels greater than the threshold value. However, tests with such loading histories often show lower fatigue lives than the predicted I

values, e.g., solid inverted triangle or solid diamond in Fig. 14a.

Similar strain-rate-change tests on austenitic SSs in PWR environments have also showed the i existence of a strain threshold below which the material is insensitive to environmental effects. 29 The threshold strain range Acth appears to be independent of material type (weld metal or base metal) and temperature in the range of 250-325°C, but it tends to decrease as the strain range is decreased. The n threshold strain range has been expressed -in terms of the applied strain range Ac by the equation Acth/Ac = - 0.22 Ae + 0.65. (19)

This expression may also be used for carbon and low-alloy steels.

Fmcbonof st inat slow tot0 Ft-rof tniin ar srole: IFraction of strainat slowtate: 0170 A B C Air: 3,253; 3,753 Air: 3,721; 3,424; 6,275 Air: 4,122 PWR: 2,230; 1,525 PWR: 2,141 PWR: -

Hi DO: 2,077; 1,756 Hi DO: 303; 469 Hi DO: 888 Fractronof strainatslow ale: 0.347 F action of strain a so, rate:0666 F'ractonof strain at slow rate: 083

_~ 7 i, D E F Air: 5,139 Air: 5,261 Air: 3,893 PWR: - PWR: - PWR: -

Hi DO: 615; 553 Iii DO: 545 Hi DO: 340 Fraction of strain at slow rate: 0.167 Fraction of strain at slo rate. 0.334 Fraction of strain at slow rate: 065 V , i VV G

Air: 4,087 PWR: -

H Air:

PWR:

-

-

I Air:

PWR:

4,356

-

I Hli DO: 1,649; 2,080 Hi DO: 1,935 Hi DO: 615 Figure 13. Fatigue life of A106-Gr B carbon steel at 2880C and 0.75% strain range in air and water environments under different loading waveforms (Ref. 4). I 24

Al 06-&r. B Carbon Steel STS410 Carbon Steel 289"C 288°C Strain Range nO.78% StnainRange -1.2%

101 Average Life in Air 10 Average Life in Air AvcLife in PVWR.Water A!

A 103 AV.Life in PWR Water i

1 O0 AW 8 at* =036%° c!

A ptt 0O5Pt 00 2%.........C 0 -_0t.X 1 13 D r,, -0.32%

Slewnte appliedneat SOtotrate appliednear Open Sytbo[s peaktesle Se~vtrain Open Symbols:peak tensile Wttain Ceaed Symnols peat tempt. ,w .iv t, , .in , , Closed Spybols: peakeampteaine strain 0,0 0.2 0.4 0.6 0.8 1.0 0.0 0.2 0.4 0.6 0.8 1.0 Fraction of Strain at Slow Strain Rate Fraction of Strain at Slow Strain Rate (a) (b)

A533-Gr. B Low-AlloySteel 289*C Strain Range =0.78%

4 10 Average Life in Air Av. Life in PWR WNater S!

f- Figure 14.

Fatigue life of carbon and low-alloy steels tested with loading waveforms where slow strain rate is 101 applied during a fraction of tensile loading cycle (Refs. 4, 18).

A PWP C 0 .8 pp ot 0 B h 0.32%

Slo tateapplied near Open Symbols: peak tensile niean CIosedSymbols:peak conptnese train ln2

] .,,it,,,,,i,,,',,,

0.0 0.2 0.4 0.6 0.8 1.0 Fraction of Strain at Slow Strain Rate (c)

The modified rate approach, described in Section 4.2.14, has been used to predict the results from tests on four heats of carbon and low-alloy steels conducted with changing strain rate in low- and high-DO water at 289'C. 18 The results indicate that the modified rate approach, without the consideration of a strain threshold, gives the best estimates of life (Fig. 15). Most of the scatter in the data is due to heat-to-heat variation rather than any inaccuracy in estimation of fatigue life; for the same loading conditions, the fatigue lives of Heat #2 of STS410 steel are a factor of z5 lower than those of Heat #1. The estimated fatigue lives are within a factor of 3 of the experimental values.

• In L WR coolant environments, the procedurefor calculating Fen, defined in Eqs. 27 and 28 (Section 4.2.13), includes a threshold strain range below which environment has no effect on fatigue life, i.e., Fen

= 1. However, while using the damage rate approach to determine Fenfor a stress cycle or load set pair, including a thresholdstrain (Eq. 31 in Section 4.2.14) may yield nonconservative estimates of life.

25

I I

Carbon & Low-Alloy Steels /

10~4 V

(, STS410 Ili STS4t0 #2 AlSo-b

/ I Factor ..

Figure 15.

Experimental values of fatigue life and those predicted from the modified rate approach without I

consideration of a threshold strain (Ref. 18).

A- I

' Modified Rate Approach 0 ,.

I without strain threshold 102 Slow rate measured from peak Open Symbols: tensile strain Closed Symbols: compressive strain I . . . I. . . .. I . . ./

102 103 Experimental Life (Cycles)

I 4.2.4 Temperature The change in fatigue life of two heats of A333-Gr 6 carbon steel 12 , 13 , 16 with test temperature at U

different levels of DO is shown in Fig. 16. Other parameters, e.g., strain amplitude and strain rate, were kept constant; the applied strain amplitude was above and strain rate was below the critical threshold.

In air, the two heats have a fatigue life of -3300 cycles. The results indicate a threshold temperature of I

150'C, above which environment decreases fatigue life if DO in water is also above the critical level.

In the temperature range of 150-320'C, the logarithm of fatigue life decreases linearly with temperature; the decrease in life is greater at high temperatures and DO levels. Only a moderate decrease in fatigue life is observed in water at temperatures below the threshold value of 150°C or at DO levels <0.05 ppm.

I Under these conditions, fatigue life in water is a factor of z2 lower than in air; Fig. 16 shows an average life of zz2000 cycles for the heat with 0.015 wt.% S, and z1200 cycles for the 0.012 wt.% S steel.

3-Gr 6 Carbon Steel 0.6%, S = 0.012 wt.%-

- -- -

a) - ---

J* 103 -

a) *to

- 10 IO5 li t I, U- 33 102 50 100 150 200 250 300 350 0 50 100 150 200 250 300 350 0 Temperature (°C) Temperature (°C)

Figure 16. Change in fatigue life of A333-Gr 6 carbon steel with temperature and DO.

I I

26 I I

An artificial neural network (ANN) has also been used to find patterns and identify the threshold temperature below which environmental effects are moderate.10 5 The main benefits of the ANN approach are that estimates of life are based purely on the data and not on preconceptions, and by learning trends, the network can interpolate effects where data are not present. The factors that affect fatigue life can have synergistic effects on one another. A neural network can detect and utilize these effects in its predictions. A neural network, consisting of two hidden layers with the first containing ten nodes and the second containing six nodes, was trained six times; each training was based on the same data set, but the order in which the data were presented to the ANN for training was varied, and the initial ANN weights were randomized to guard against overtraining and to ensure that the network did not arrive at a solution that was a local minimum. The effect of temperature on the fatigue life of carbon steels and low-alloy steels estimated from ANN is shown in Fig. 17 as dashed or dotted lines. The solid line represents estimates based on the ANL model, and the open circles represent the experimental data. The results indicate that at high strain rate (0.4%/s), fatigue life is relatively insensitive to temperature. At low strain rate (0.004%/s), fatigue life decreases with an increase in temperature beyond a threshold value of z1 50'C. The precision of the data indicates that this trend is present in the data used to train the ANN.

Nearly all of the fatigue c-N data have been obtained under loading histories with constant strain rate, temperature, and strain amplitude. The actual loading histories encountered during service of nuclear power plants involve variable loading and environmental conditions. Fatigue tests have been conducted in Japan on tube specimens (1- or 3-mm wall thickness) of A333-Gr 6 carbon steel in oxygenated water under combined mechanical and thermal cycling.1 5 Triangular waveforms were used for both strain and temperature cycling. Two sequences were selected for temperature cycling (Fig. 18):

104 10-Carbon Steels Carbon Steels 0.012 wt.% Sulfur 0.012 wt.% Sulfur En 0 aI)

Li)

(-2 II)

-u 3 103' **-, """-Lo 10 3.

C-L

.0 76 Dissolved Oxygen = 0.2 ppm Dissolved Oxygen 0.2 ppm LL Strain Rate 0.4 %/s Strain Rate 0.004 %/s E =0.6% = 0.6%

102 102 0 50 100 150 200 250 300 350 0 50 100 150 200 250 300 350 Temperature (°C) Temperature (°C) 104 104 , ,I -

Low-Alloy Steels Low-Alloy Steels 0.012 wt.% Sulfur U) 0.012 wt.% Sulfur ci)

=0.6% >4

=0.6%

0 ci)

CU 3 ci) 10 3 -J 10 -~

ci) 0)

(S U-Dissolved Oxygen = 0.2 ppm Dissolved Oxygen 0.2 ppm Strain Rate = 0.4 %/s Strain Rate 0.004 %/s 102 . 1. I . . " . . I 1U-0 50 100 150 200 250 300 350 0 50 100 150 200 250 300 350 Temperature (°C) Temperature (°C)

Figure 17. Dependence of fatigue life on temperature for carbon and low-alloy steels in water.

27

I an in-phase sequence in which temperature cycling was synchronized with mechanical strain cycling, and I

another sequence in which temperature and strain were out of phase, i.e., maximum temperature occurred at minimum strain level and vice versa. Three temperature ranges, 50-290'C, 50-200'C, and 200-290'C, were selected for the tests. The results are shown in Fig. 19; an average temperature is used to I plot the thermal cycling tests. Because environmental effects on fatigue life are moderate and independent of temperature below 150'C, the temperature for tests cycled in the range of 50-290'C or 50-200'C was determined from the average of 150"C and the maximum temperature. The results in Fig. 19 indicate that load cycles involving variable temperature conditions may be represented by an I

average temperature, e.g., the fatigue lives from variable-temperature tests are comparable with those from constant-temperature tests. I 0.6 High 0.6 High I C

._) CL C I ci, E E I-Low 4)

I--

Low I

-0.6 -0.6 Figure 18. Waveforms for change in temperature during exploratory fatigue tests. I A333-Gr. 6.Carbon Steel Tube specimen; 3-mm wall

= 0.6%;

-ý c = 0.002 %/s 104~

. .I.. . . I . . .I . . . . i . . . . i . . . i . . .

A333-Gr. 6 Carbon Steel Tube specimen; 1-mm walt i0 I Temperature]

A Constant i M phasieI I

E. = 0.6%.: = 0.002 %Is C) S = 0.012 wt.%; DO 1 ppm S = 0.012 wt.%, DO t ppmat 10,

.............

. ....................

[................* '"

75 103 I

LL a) 102 Temperature.

0 Constant I

I

• 2) 76

.................

. .. ...................

.

. . .... . . . ......... , * , *

.I *.......

I A In phase --.. . . 0)

Oofut p"rase 0 50 100 150 200 Temperature (°C) 250 300 35 03 50 100 150 Temperature (°C) 200 250 300 350 I Figure 19. Fatigue life of A333-Gr 6 carbon steel tube specimens under varying temperature, indicated by horizontal bars. I However, the nearly identical fatigue lives of the in-phase and out-of-phase tests are somewhat surprising. If we consider that the tensile-load cycle is primarily responsible for environmentally assisted reduction in fatigue life, and that the applied strain and temperature must be above a minimum threshold I

value for environmental effects to occur, then fatigue life for the out-of-phase tests should be longer than for the in-phase tests, because applied strains above the threshold strain occur at temperatures above 150"C for in-phase tests, whereas they occur at temperatures below 150'C for the out-of-phase tests. If I environmental effects on fatigue life are considered to be minimal below the threshold values of 150'C for temperature and <0.25 % for strain range, the average temperatures for the out-of-phase tests at I

28 I

50-290°C, 50-200'C, and 200-290'C should be 195, 160, and 236°C, respectively, instead of 220, 175, and 245°C, as plotted in Fig. 19. Thus, the fatigue lives of out-of-phase tests should be at least 50%

higher than those of the in-phase tests. Most likely, difference in the cyclic hardening behavior of the material is affecting fatigue life of the out-of-phase tests.

• In L WR environments, the effect of temperature on the fatigue life of carbon and low-alloy steels is explicitly considered in Fen defined in Eqs. 27 and 28 (Section 4.2.13). Also, an average temperature may be used to calculate Fen for a specific stress cycle or load set pair.

4.2.5 Dissolved Oxygen The dependence of fatigue life of carbon steel on DO content in water 1 2, 13 , 16 is shown in Fig. 20.

The test temperature, applied strain amplitude, and S content in steel were above, and strain rate was below, the critical threshold value. The results indicate a minimum DO level of 0.04 ppm above which environment decreases the fatigue life of the steel. The effect of DO content on fatigue life saturates at 0.5 ppm, i.e., increases in DO levels above 0.5 ppm.do not cause further decreases in life. In Fig. 20, for DO levels between 0.04 and 0.5 ppm, fatigue life appears to decrease logarithmically with DO. Estimates of fatigue life from a trained ANN also show a similar effect of DO on the fatigue life of carbon steels and low-alloy steels.

1 04 . . . . .J . . . . . . . .I . . . .1 04 . . . . . . . . .. t . . . .

A333-6 Steel 288'C A333-6 Steel 250'C Strain Amplitude: 0.6% Strain Amplitude: 0.6%

'AA O103-U W 13C 0U 00

--J O 0) . O Strain Rate (%Nsi i o' Strain Rate (%/s)

, C0 0.004 (0012% S) ' 0 0.004 (0.012% S) 10 001 S ý 0A 1005 i ,0;51, 1102 C0 0.0 (0.012% S) 0.002 102 0 0o* (0.015%

iO,

",,r.,.,,:S'* 00102 0.002 012% S)

.

F - I I .l I I I, I 'd I - - f llll . I I I 1 1 - I -

1 1 .' .I I - I I . I . ...  ! . ' , , ;

10-3 10-2 10-1 100 101 10-3 10-2 10-1 100 10' Dissolved Oxygen (ppm) Dissolved Oxygen (ppm)

Figure 20. Dependence on DO of fatigue life of carbon steel in high-purity water.

Environmental effects on the fatigue life of carbon and low-alloy steels are minimal at DO levels below 0.04 ppm, i.e., in low-DO PWR or hydrogen-chemistry BWR environments. In contrast, 104 environmental enhancement of CGRs has been observed in low-alloy steels even in low-DO water.

This apparent inconsistency of fatigue c-N data with the CGR data may be attributed to differences in the environment at the crack tip. The initiation of environmentally assisted enhancement of CGRs in low-alloy steels requires a critical level of sulfides at the crack tip. 104 The development of this critical sulfide concentration requires a minimum crack extension of 0.33 mm and CGRs in the range of 1.3 x 10-4 to 4.2 x 10-7 mrm/s. These conditions are not achieved under typical s-N tests. Thus, environmental effects on fatigue life are expected to be insignificant in low-DO environments.

e In L WR environments, effect of DO level on the fatigue life of carbon and low-alloy steels is explicitly consideredin Fen, defined in Eqs. 27 and 28 (Section 4.2.13).

29

I 4.2.6 Water Conductivity In most studies the DO level in water has generally been considered the key environmental I parameter that affects the fatigue life of materials in LWR environments. Studies on the effect of the concentration of anionic impurities in water (expressed as the overall conductivity of water), are somewhat limited. The limited data indicate that the fatigue life of WB36 low-alloy steel-,at 177°C in water with :8 ppm DO decreased by a factor of z6 when the conductivity of water was increased from 0.06 to 0.5 gS/cm. 48' 1 06 A similar behavior has also been observed in another study of the effect of conductivity on the initiation of short cracks.

107 *

• Normally, plants are unlikely to accumulate many fatigue cycles under off-normal conditions. Thus, effects of water conductivity on fatigue life have not been consideredin the determination of Fen.

4.2.7 Sulfur Content in Steel It is well known that S content and morphology are the most important material-related parameters 10 8 1 I

that determine susceptibility of low-alloy steels to environmentally enhanced fatigue CGRs. - 11 A critical concentration of S2- or HS- ions is required at the crack tip for environmental effects to occur.

Both the corrosion fatigue CGRs and threshold stress intensity factor AKth are a function of the S content l in the range 0.003-0.019 wt.%. 1 10 The probability of environmental enhancement of fatigue CGRs in precracked specimens of low-alloy steels appears to diminish markedly for S contents <0.005 wt.%.

The fatigue c-N data for low-alloy steels also indicate a dependence of fatigue life on S content.

• When all the threshold conditions are satisfied, environmental effects on the fatigue life increase with increased S content. The fatigue lives of A508-C1 3 steel with 0.003 wt.% S'and A533-Gr B steel with 0.010 wt.% S are plotted as a function of strain rate in Fig. 21. However, the available data sets are too sparse to establish a functional form for dependence of fatigue life on S content and to define either a threshold for S content below which environmental effects are unimportant or an upper limit above which the effect of S on fatigue life may saturate. A linear dependence of fatigue life on S content has been assumed in correlations for estimating fatigue life of carbon steels and low-alloy steels in LWR environments. 4 , 79 The limited data suggest that environmental effects on fatigue life saturate at S 4

contents above 0.015 wt.%.

The existing fatigue c-N data also indicate significant reductions in fatigue life of some heats of carbon steel with S levels as low as 0.002 wt.%. The fatigue lives of several heats of A333-Gr 6 carbon Low-Alloy Steel 288TC Strain Amplitude: 0.6% I . I 8 ppm Dissolved Oxygen , . Figure 21.

1 3 . . ..... . . *.... .* :igure 21

" . . I Effect of strain rate on fatigue life of low-alloy V I steels with different S contents (JNUFAD U-102 -- database and Ref. 4).

I50-Cl 3 (0o003 Mio% S) 1 0i V A533-Gr B (0,010 M.1%S) 101 5

10- 10-4 10-3 10-2 10-1 100 Strain Rate (%is) 30 lI

steel with S contents of 0.002-0.015 wt.% in high-DO water at 288°C and 0.6% strain amplitude are plotted as a function of strain rate in Fig. 22.4 Environmental effects on the fatigue life of these steels seem to be independent of S content in the range of 0.002-0.015 wt.%. However, these tests were conducted in air-saturated water (z8 ppm DO). The fatigue life of carbon steels seems to be relatively insensitive to S content in very high DO water, e.g., greater than I ppm DO; under these conditions, the effect of DO dominates fatigue life. In other words, the saturation DO level of 0.5 ppm most likely is for medium- and high-S steels (i.e., steels with >0.005 wt.% S); it may be higher for low-S steels.

4 10 A333-Gr 6 Carbon Steel 288°C Strain Amplitude: 0.6%

8 ppm Dissolved Oxygen

. .. . .... ... .

L) 103 . .. ............... . Figure 22.

V Effect of strain rate on the fatigue life of

.A333-Gr6 - carbon steels with different S contents.

0 102 . .... ................

Le Sulfur (wt.%)

0 0.006 V7 0.012 A 0.002 101'1 -J I 1 10-5 10-4 10-3 10-2 10-1 100 Strain Rate (%Is)

In L WR environments, the effect of S content on the fatigue life of carbon and low-alloy steels is explicitly considered in Fen, defined in Eqs. 27 and 28 (Section 4.2.13). However, evaluation of experimental data on low-S steels (<O.005 wt. % S) in water with >1 ppm DO should be done with caution; the effect of S may be larger than that predicted by Eqs. 2 7 and 28.

4.2.8 Tensile Hold Period Fatigue tests conducted using trapezoidal waveforms indicate that a hold period at peak tensile strain decreases the fatigue life of carbon steels in high-DO water at 289'C. 4 , 18 However, a detailed examination of the data indicated that these results are either due to limitations of the test procedure or caused by a frequency effect. Loading waveforms, hysteresis loops, and fatigue lives for the tests on A106-Gr B carbon steel in air and water environments are shown in Fig. 23.4 A 300-s hold period is sufficient to reduce fatigue life by '50% (z2000 cycles without and z 1000 cycles with a hold period); a longer hold period of 1800 s results in only slightly lower fatigue life than that with a 300-s hold period.

For example, two 300-s hold tests at 288°C and z0.78% strain range in oxygenated water with 0.7 ppm DO gave fatigue lives of 1,007 and 1,092 cycles; life in a 1800-s hold test was 840 cycles. These tests were conducted in stroke-control mode and are somewhat different from the conventional hold-time test in strain-control mode, where the total strain in the sample is held constant during the hold period.

However, a portion of the elastic strain is converted to plastic strain because of stress relaxation. In a stroke-control test, there is an additional plastic strain in the sample due to relaxation of elastic strain from the load train (Fig. 23). Consequently, significant strain changes occur during the hold period; the

,measured plastic strains during the hold period were z0.028% from relaxation of the gauge and 0.05-0.06% from relaxation of the load train. These conditions resulted in strain rates of 0.005-0.02%/s during the hold period. The reduction in life may be attributed to the slow strain rates during the hold period.

Also, frequency effects may decrease the fatigue life of hold time tests, e.g., in air, the fatigue life of stroke-control test with hold period is z50% lower than that without the hold period.

31

I Hold-time tests have also been conducted on STS410 carbon steel at 289°C in water with 1 ppm I

DO. The results are given in Table 5.M8 The most significant observation is that a reduction in fatigue life occurs only for those hold-time tests that were conducted at fast strain rates, e.g., at 0.4%/s. At lower strain rates, fatigue life is essentially the same for the tests with or without hold periods. Based on these I

results, Higuchi et al. 18 conclude that the procedures for calculating Fen need not be revised. Also, as discussed in Section 4.2.11, the differences in fatigue life of these tests are within the data scatter for the fatigue E-N data in LWR environments. I The existing data do not demonstrate that hold periods at peak tensile strain affect the fatigue life of carbon and low-alloy steels in L WR environments. Thus, any revision/modazcation of the method to determine Fen is not warranted I

600 600 600 K I , I , I, T1 I

400 400 i i -! / ..... ; ... /i..

400

.....--

i .i. '.

....- -i.-- -. . ----- - i 02 200

-ijI I 'I I a

200 R

200

- 7/l

....

/ii I

/, V I

0 0 0

_,j

-200

-200 -200 _ / i /. i _

-400 -400 -400

-600

-/.. ./,"

~...! ] i I

-600 -600

-0.4 -0.2 0 0.2 0.4 -0.4 -0.2 0 0.2 04 Strain (o Fraction of strain at slow rate: 0 Strain (%) Strain (%)

I I

Strain or Stroke Control Strain Control Stroke Control (5-min hold)

Air: 4740+/-1250 cycles Air: 6,275 cycles Air: 2592 cycles PWR: 1965+/-385 cycles High DO: 1007, 1092 cycles Stroke Control (30-min hold)

High DO: 2077 cycles High DO: 840 cycles I Figure 23. Fatigue life of Al 06-Gr B steel in air and water environments at 288°C, 0.78% strain range, and hold period at peak tensile strain (Ref. 4). Hysteresis loops are for tests in air.

I I

I 32

Table 5. Fatigue data for STS410 steel at 289°C in water with 1 ppm DO and trapezoidal waveform.

Strain Ampl. Hold Period at Peak Tensile / Compressive Strain Rate (%/s)

(%) Tensile Strain (s) 0.4 /0.4 0.04 /0.4 0.01 /0.4 0.004 /0.4 0.6 0 489 240 118 0.6 60 328,405 238 138 0.6 600 173,217 - -

0.3 0 3270 1290 737 508 0.3 60 1840, 1760 1495 875 587 0.3 600 436, 625 - -

4.2.9 Flow Rate Nearly all of the fatigue E-N data for LWR environments have been obtained at very low water flow rates. Recent data indicate that, under the environmental conditions typical of operating BWRs, environmental effects on the fatigue life of carbon steels are at least a factor of 2 lower at high flow rates (7 m/s) than at 0.3 m/s or lower. 19 ,20 ,44 The beneficial effects of increased flow rate are greater for high-S steels and at low strain rates. 19 ,2 0 The effect of water flow rate on the fatigue life of high-S (0.0 16 wt.%) A333-Gr 6 carbon steel in high-purity water at 289°C is shown in Fig. 24. At 0.3% strain amplitude, 0.01%/s strain rate, and all DO levels, fatigue life is increased by a factor ofz2 when the flow rate is increased from z10-5 to 7 m/s. At 0.6% strain amplitude and 0.001%/s strain rate, fatigue life is increased by a factor ofz6 in water with 0.2 ppm DO and by a factor of=3 in water with 1.0 or 0.05 ppm DO. Under similar loading conditions, i.e., 0.6% strain amplitude and 0.001%/s strain rate, a low-S (0.008 wt.%) heat of A333-Gr 6 carbon steel showed only a factor of *2 increase in fatigue life with increased flow rates. Note that the beneficial effects of flow rate are determined from a single test on each material at very low flow rates; data scatter in LWR environments is typically a factor ofZ2.

A factor of 2 increase in fatigue life was observed (Fig. 25) at KWU during component tests with 180' bends of carbon steel tubing (0.025 wt.% S) when internal flow rates of up to 0.6 m/s were established. 4 4 The tests were conducted at 240'C in water that contained 0.2 ppm DO.

e Because of the uncertainties in the flow conditions at or near the locations of crack initiation, the beneficial effect offlow rate on the fatigue life is presently not included in fatigue evaluations.

104 ....

A333-Gr 6 Carbon Steel (High-S) A333-Gr 6 Carbon Steel (High-S).

104 z

,_..---------------------

.. .. 0---- _o. . . - - - --

-q103 ...... O 0-

2) g 102 _8" ELLu DO (ppm) DO (ppm) 289'C e 1.0 289°C G 1.0 Strain Amplitude 0.3% 0.2 Strain Amplitude 0.6% 0.2 102 Strain Rate 0.01%/s A-- 0.05 Strain Rate 0.001%/s A 0.05 L

,I , , , . , .I , _ _"* , , ,l, 0,5H...I * ,10 1 ,,, -I , , ,, ,,,,, ,, , , , I .,, ",I . ,1 5 4 3 4 10- 10- 10- 10-2 10-1 100 101 10-5 10- 10-3 10-2 10-1 100 101 Flow Rate (m/s) Flow Rate (m/s)

(a) (b)

Figure 24. Effect of water flow rate on fatigue life of A333-Gr 6 carbon steel at 289°C and strain amplitude and strain rates of (a) 0.3% and 0.01 %/s and (b) 0.6% and 0.001 %/s.

33

I

. . . . . . ..ii I . . i. . . . . ;. ;. i. . i. i. I '

I Carbon Steel (0.025% S)

W 1.0 240'C, Strain Rate: 0.001%/s

' "

Best-Fit Curv RRT Air ve I

6 ASME Code I

"

Figure 25.

Design Curve 00 " Effect of flow rate on low-cycle fatigue of E carbon steel tube bends in high-purity water at 2400C (Ref. 44). RT = room temperature.

I Dissolved Oxygen 0; 0.2 ppm O 0.01 ppm Open Symbols; Low flow Closed Symbols: 0.6 m/s flow rate o111 I

I I 0.1 10I1 102 103 104 Fatigue Life (Cycles) 4.2.10 Surface Finish I Fatigue testing has been conducted on specimens of carbon and low-alloy steels that were intentionally roughened in a lathe, under controlled conditions, with 50-grit sandpaper to produce circumferential scratches with an average roughness of 1.2 ltrm and Rq of 1.6 ltm (z62 micro in.). 3 9 The I

results for A106-Gr B carbon steel and A533-Gr B low-alloy steel are shown in Fig. 26. In air, the fatigue life of rough A106-Gr B specimens is a factor of 3 lower than that of smooth specimens, and, in high-DO water, it is the same as that of smooth specimens. In low-DO water, the fatigue life of the roughened A I06-Gr B specimen is slightly lower than that of smooth specimens. The effect of surface I

roughness on the fatigue life of A533-Gr B low-alloy steel is similar to that for A106-Gr B carbon steel; in high-DO water, the fatigue lives of both rough and smooth specimens are the same. The results in water are consistent with a mechanism of growth by a slip oxidation/dissolution process, which seems I

unlikely to be affected by surface finish. Because environmental effects are moderate in low-DO water, surface roughness would be expected to influence fatigue life.

I I I I I II I I I'll r 11ITjII _ ý TTI 11,1 . . 1I ý I I  :. . . l - H. 1 ý - - I . i ý~~

I A533 Gr B Low-Alloy Steel . Ai 0.004%/s A106 Gr B Carbon Steel . Air. 0.04D',/s 289°C A Air, 0.0 %/s 289°C A Air. 0.4%/s 1.0 =700 ppb DO Water, 1.0 0 =700 ppb DO Water, 0.004%/s

• 0.004%/s V 5 ppb DO Water, a) n 0.

x... I F ."

0 ASME Code C

CV U) 0.1 Design Curve Open Symbols: Smooth Specimens Closed Symbols: Rough Surface, 50 grit paper 0 ".

0.1 ASME Code Design Curve Open Symbols: Smooth Specimens I

Closed Symbols: Rough Surface, 50 grit paper 102 10 3

104 Fatigue Life (Cycles) 10 5

101 102 1(0 3

10 4

105 Fatigue Life (Cycles) 106 I Figure 26. Effect of surface roughness on fatigue life of (a) A106-Gr B carbon steel and (b) A533 low-alloy steel in air and high-purity water at 289°C. I 34 i

i

• The effect of surface finish is not considered in the environmental fatigue correction factor; it is included in the subfactorfor "surfacefinish and environment" that is applied to the mean data curve to develop the Code fatigue design curve in air.

4.2.11 Heat-to-Heat Variability The effect of material variability and data scatter on the fatigue life of carbon and low-alloy steels has also been evaluated for LWR environments. The fatigue behavior of each of the heats or loading conditions is characterized by the value of the constant A in the ANL models (e.g., Eq. 6). The values of A for the various data sets are ordered, and median ranks are used to estimate the cumulative distribution of A for the population. Results for carbon and low-alloy steels in water environments are shown in Fig. 27. The median value of A in water is 5.951 for carbon steels and 5.747 for low-alloy steels. The results indicate that environmental effects are approximately the same for the various heats of these steels.

For example, the cumulative distribution of data sets for specific heats is approximately the same in air and water environments. The ANL model seem to overestimate the effect of environment for a few heats, e.g., the ranking for A533-Gr B heat 5 is --42 percentile in air and =95 percentile in water, and for A106-Gr B heat A, it is z17 percentile in air and varies from 2 to 60 percentile in water. Monte Carlo analyses were also performed for the fatigue data in LWR environments.

1.0 .. .I I . .1 1 1. . 1.. . I 11 0 W-carbon Steel t .... Low-Alloy Steel

-Water Environmet Water Environment-,* .

0.8 - 75th P_ -- _.. 0.8 75th Percentile 7 r e . . . / . ...... .

01 P-0.6

---

Median5951 "- ] ...-. ___.

_T__

_

0.6 0

-

--rr 1

- .--- 13Heat Median 5. .--

L

-

• I-------- > --Median

- --..-5.747 i , 327 Data S ,e~ Points a 8 Data Points 0.4 12 Heats 7 A302-8 U 5ti 0jjIF'IU j--

I ( A,0-

-es: (T---25th - T---*

Prcetil~~~0 ---- *

• "I A533-t3 (5)*

A533-B

.A533-B (1.)

(1)

Percenttile 0 A__-B (A) 02__ X A533-B (M) 0.2 0.2i......} 4 .. i - A A333-6 A306-B

..........-

(2)

(2) 0.2 -

-- - . _*. ._

__.._...._{*..

-

V 0l A508-2 A-o--

A508-3 (1)

(1

( 1) 0 A333-6 (3) 8 A0- 1

__ ",A A333-6 (5) 0 A508-3 (7)

A333-6 7) A508(M) 0.0 ,, 0.0 11.. . .. ............ .

4 4.5 5 5.5 6 6.5 7 7.5 8 4 4.5 5 5.5 6 6.5 7 7.5 8 Constant A Constant A Figure 27. Estimated cumulative distribution of parameter A in the ANL models for fatigue life for heats of carbon and low-alloy steels in LWR environments.

The results for carbon and low-alloy steels in LWR environments are summarized in Tables 6 and 7, respectively, in terms of values for A that provide bounds for the portion of the population and the confidence that is desired in the estimates of the bounds. In LWR environments, the 5th percentile value of parameter A at 95% confidence level is 5.191 for carbon steels and 4.748 for low-alloy steels. From Fig. 27, the median value of A for the sample is 5.951 for carbon steels and 5.747 for low-alloy steels.

Thus, the 95/95 value of the margin to account for material variability and data scatter is 2.1 and 2.7 on life for carbon steels and low-alloy steels, respectively. These margins are needed to provide 95%

confidence that the resultant life will be greater than that observed for 95% of the materials of interest.

35

I Table 6. Values of parameter A in the ANL fatigue life model for carbon steels in water and the I

margins on life as a function of confidence level and percentage of population bounded.

Confidence Percentage of Population Bounded (Percentile Distribution of A)

I Level 95(5) 90(10) 75(25) 67 (33i 50(50) 50 75 5.333 5.275 5.469 5.417 Values of Parameter A 5.697 5.652 5.786 5.742 5.951 5.906 I

I 95 5.191 5.342 5.587 5.678 5.840 Margins on Life 50 1.9 1.6 1.3 1.2 1.0 75 2.0 1.7 1.3 1.2 1.0 Table 7.

95 2.1 1.8 1.4 1.3 Values of parameter A in the ANL fatigue life model for low-alloy steels in water and the 1.1 I

margins on life as a function of confidence level and percentage of population bounded.

Confidence Level 95(5)

Percentage of Population Bounded (Percentile Distribution of A) 90(10) 75(25) 67(33) 50 (50)

I Values of Parameter A 50 75 95 4.950 4.867 4.748 5.126 5.052 4.944 5.420 5.355 5.261 5.534 5.470 5.378 5.747 5.680 5.585 I

I Margins on Life 50 2.2 1.9 1.4 1.2 1.0 75 2.4 2.0 1.5 1.3 1.1 95 2.7 2.2 1.6 1.4 1.2 The effect of heat-to-heat variabilityis not considered in the environmentalfatigue correctionfactor; it is included in the subfactorfor "data scatter and material variability" that is applied to the mean data I

curve to develop the Code fatigue design curve in air.

4.2.12 Fatigue Life Model I

Fatigue-life models for estimating the fatigue lives of carbon and low-alloy steels in LWR environments based on the existing fatigue F-N data have been developed at ANL. 4 ,39 The effects of key parameters, such as temperature, strain rate, DO content in water, and S content in the steel, are included I

in the correlations; the effects of these and other parameters on the fatigue life are discussed below in detail. The functional forms for the effects of strain rate, temperature, DO level in water, and S content in the steel were based on the data trends. For both carbon and low-alloy steels, the model assumes I

threshold and saturation values of 1.0 and 0.001%/s, respectively, for strain rate; 0.001 and 0.015 wt.%,

respectively, for S; and 0.04 and 0.5 ppm, respectively, for DO. It also considers a threshold value of 150'C for temperature.

I In the present report these models have been updated based on the analysis presented in Section 4.2.11, e.g., constant A in the models differs from the value reported earlier in NUREG/CR-6583 and -

6815. Relative to the earlier model, the fatigue lives predicted by the updated model are z6% lower for I

carbon steels and z2% higher for low-alloy steels. In LWR environments, the fatigue life, N, of carbon steels is represented by I ln(N) = 5.951 - 1.975 ln(Ea-0.113)+0.10l S* T* O* *, (20)

I 36 I

I

106

• 105-

"3 104 .. . V' *

~~0 a)1021 3 lo , ,4*

Ae 101,/ .

101 102 103 104 105 3 4 106 102 10 10 105 Observed Life (Cycles) Observed Life (Cycles)

(a) (b)

Figure 28. Experimental and predicted fatigue lives of (a) carbon steels and (b) low-alloy steels in LWR environments.

and that of low-alloy steels, by ln(N) =5.747 -l.808 ln(aa,, -0.15 1) +0.101 S* T*0* t * (21) where S*, T*, 0*, and *

• are transformed S content, temperature, DO level, and strain rate, respectively, defined as:

S* = 0.015 (DO > 1.0 ppm)

S* = 0.001 (DO <1.0 ppm and S < 0.001 wt.%)

S*= S (DO _<1.0 ppm and 0.001 < S < 0.015 wt.%)

S*= 0.015 (DO <1.0 ppm and S > 0.015 wt.%) (22)

T* =0 (T

• 150-C)

T* = T - 150 (150 < T < 350-C) (23) 0*6=0 (DO < 0.04 ppm) 0*= ln(DO/0.04) (0.04 ppm < DO < 0.5 ppm) 0* = ln(12.5) (DO > 0.5 ppm) (24)

(t > 1%/s)

(0.001 < a _<1%/s) t ln(0.00 1) (t; < 0.001%/s). (25)

These models are recommended for predicted fatigue lives <100 cycles. Also, as discussed in Section 4.2.7, because the effect of S on the fatigue life of carbon and low-alloy steels appears to depend on the DO level in water, Eqs. 20-25 may yield nonconservative estimates of fatigue life for low-S

(<0.007 wt.%) steels in high-temperature water with >1 ppm DO. The experimental values of fatigue life and those predicted by Eqs. 20 and 21 are plotted in Fig. 28. The predicted fatigue lives show good agreement with the experimental values; the experimental and predicted values differ by a factor of 3.

37

I The ANL fatigue life models represent the mean values offatigue life as a function of applied strain amplitude, temperature, strain rate, DO level in water, and S content of the steel. The effects of parameters (such as mean stress, surface finish, size and geometry, and loading history) that are known to influencefatigue life are not included in the model.

I 4.2.13 Environmental Fatigue Correction Factor The effects of reactor coolant environments on fatigue life have also been expressed in terms of environmental fatigue correction factor, Fen, which is -defined as the ratio of life in air at room temperature, NRTair, to that in water at the service temperature, Nwater. Values of Fen can be obtained from the ANL fatigue life model, where ln(Fen) = ln(NRTair) - ln(Nwater). (26) I The environmental fatigue correction factor for carbon steels is given by Fen = exp(0.632 - 0.101 S* T* 0* t*) (27) and for low-alloy steels, by Fen = exp(O.702 - 0.101 S* T* 0* t*) (28) where the constants S*, T*, r *, and 0* are defined in Eqs. 22-25. Note that because the ANL fatigue m life models have been updated in the present report, the constants 0.632 and 0.702 in Eqs. 27 and 28 are different from the values reported earlier in NUREG/CR-6583 and -6815.

expressions, correction factors determined from Eq. 27 for carbon steels are Relative to the earlier

'8% higher, and those determined from Eq. 28 for low-alloy steels are z18% lower. A threshold strain amplitude (one-half of I

the applied strain range) is also defined, below which LWR coolant environments have no effect on fatigue life, i.e., Fen = 1. The threshold strain amplitude is 0.07% (145 MPa stress amplitude) for carbon and low-alloy steels. To incorporate environmental effects into a ASME Section I1 fatigue evaluation, I

the fatigue usage for a specific stress cycle of load set pair based on the current Code fatigue design curves is multiplied by the correction factor. Further details for incorporating environmental effects into fatigue evaluations are presented in Appendix A.

0 The Fen approach may be used to incorporateenvironmental effects into the Codefatigue evaluations. n 4.2.14 Modified Rate Approach Nearly all of the existing fatigue a-N data were obtained under loading histories with constant strain rate, temperature, and strain amplitude. The actual loading histories encountered during service of nuclear power plants are far more complex. Exploratory fatigue tests have been conducted with waveforms in which the test temperature and strain rate were changed. 4 , 15 , 18 The results of such tests provide guidance for developing procedures and rules for fatigue evaluation of components under complex loading histories.

The modified rate approach has been proposed to predict fatigue life under changing test m conditions. 3 1,32 It allows calculating Fen under conditions where temperature and strain rate are changing. The correction factor, Fen(t, T), is assumed to increase linearly from I with increments of m 38 I

strain from a minimum value Emin (%) to a maximum value Fmax (%). Increments of Fen, dFen, during increments of strain, de, are calculated from dFen = (Fen - 1) de /(Fmax - Emin)- (29)

Integration of Eq. 29 from Smin to max provides the environmental fatigue correction factor under changing temperature and strain rate. The application of the modified rate approach to a strain transient is illustrated in Fig. 29; at each strain increment, Fen( ,T) is determined from Eqs. 27 and 28. Thus, Fen for the total strain transient is given by n A en = *- Fk*k,Tk) - , (30) k=1 rmax min where n is the total number of strain increments, and k is the subscript for the k-th incremental segment.

As discussed in Section 4.2.3, a minimum threshold strain, Eth (one-half of the applied strain range),

is required for an environmentally assisted decrease in fatigue life. During a strain cycle, environmental effects are significant only after the applied strain levelexceeds the threshold value. In application of the modified rate approach when a threshold strain Eth is considered, F,, for the total strain transient is given by n k=Ik e

Fe I~ Fen~k('k,Tk) (31)n k1Fmax -(mn+ Eth)()

Emax - - - - - - - - - - - - - -

Figure 29.

Application of the modified rate approach to T Adetermine the environmental fatigue correction U) T3 AE b*y&

factor Fen during a transient.

E,--------- ---- E AEk Ek =-- Atk Time The modified rate approach has been used to evaluate fatigue life under cyclic loading conditions where both temperature and strain rate were varied during the test. 18 , 3 1,32 The studies demonstrate the applicability of the damage rate approach to variable loading conditions such as actual plant transient.

Also, the following conclusions may be drawn from these studies.

(a) The use of a strain threshold, Fth, for calculating Fen by the modified rate approach (i.e., Eq. 3 1) is not necessary because it does not improve the accuracy of estimation. 3 2 As discussed earlier in 39

I Section 4.2.3, application of the modified rate approach, without the consideration of a strain threshold, gives the best estimates of fatigue life.

(b) Under load cycles that involve variable strain rate, estimates of Fen based on an average strain rate

[i.e., in Fig. 29, total strain (Fmax - Emin) divided by the total time for the transient] are the most conservative.1 8 Thus, calculations of Fen based on an average strain rate for the transient will always yield a conservative estimate of fatigue life.

(c) An average temperature for the transient may be used to estimate Fen during a load cycle.

0 Where information is availableregarding the transients associatedwith a specific stress cycle or load set pair,the modified rate approach may be used to determine Fen.

I I

I I

I I

I I

I I

I 40i I

5 Austenitic Stainless Steels The relevant fatigue a-N data for austenitic SSs in air include the data compiled by Jaske and O'Donnell 72 for developing fatigue design criteria for pressure vessel alloys, the JNUFAD database from Japan, studies at EdF in France, 6 9 and the results of Conway et al. 73 and Keller. 74 In water, the existing fatigue a-N data include the tests performed by GE in a test loop at the Dresden I reactor; 8- 11 the JNUFAD data base; studies at MHI, IHI, and Hitachi in Japan; 18- 30 the work at ANL; 6 ,7,36 -40 and the studies sponsored by EdF. 7 0- 71 Nearly 60% of the tests in air were conducted at room temperature, 20%

at 250-325°C, and 20% at 350-450'C. Nearly 90% of the tests in water were conducted at temperatures between 260 and 325°C; the remainder were at lower temperatures. The data on Type 316NG in water have been obtained primarily at DO levels >0.2 ppm, and those on Type 316 SS, at <0.005 ppm DO; half of the tests on Type 304 SS are at low-DO and the remaining at high-DO levels.

5.1 Air Environment 5.1.1 Experimental Data The fatigue a-N data for Types 304, 316, and 316NG SS in air at temperatures between room temperature and 456'C are shown in Fig. 30. The best-fit curve based on the updated ANL fatigue life model (Eq. 32 in Section 5.1.7) and the ASME Section II1mean-data curves are included in the figures.

The results indicate that the fatigue life of Type 304 SS is comparable to that of Type 316 SS; the fatigue Type 304 SS Type 316 SS P.i (C 290-C

, 325^C V-*- > 400'C 1.0 >v 300'C - 0 V 456'C ASMECode -

ASME Code E I MeanIO ° Curve E Mean Curve -

Best PdAirt Best-Fit Air ,

0.1 -ANL Model 0.1 ANLModel H 11. I rI ...... r1.. . I rrrrrrrr I r ....... I 4 5 3 5 102 103 10 10 106 107 102 10 104 10 106 107 108 108 Fatigue Life (Cycles) Fatigue Life (Cycles)

-. 1.

. 1.. ...... 1 . . ... ..1 Type 316NG L- 320WC Figure 30.

.1 1.0 Fatigue E-N behavior for Types 304, 316, and ASME Code 316NG austenitic stainless steels in air at various Mean Curve temperatures E

(Refs. JNUFAD data, 7, 36-38, 72, 73, 74).

U)

Best-FitAir 0.1 ANL Model 3 4 102 10 10 105 106 107 108 Fatigue Life (Cycles) 41

I life of Type 316NG is slightly higher than that of Types 304 and 316 SS at high strain amplitudes. Some of the tests on Type 316 SS in room-temperature air have been conducted in load-control mode at stress levels in the range of 190-230 MPa. The data are shown as triangles in Fig. 30, with strain amplitudes of U 0.1-0.12% and fatigue lives of 7 x 104-3 x 101. For these tests, the strain amplitude was calculated only as elastic strain. Based on cyclic stress-vs.-strain correlations for Type 316 SS,38 actual strain amplitudes for these tests should be 0.23-0.32%. These results were excluded from the analysis of the fatigue s-N data to develop the model for estimating the fatigue life of these steels in air.

I The results also indicate that the current Code mean-data curve is not consistent with the existing fatigue s-N data. At strain amplitudes <0.3% (stress amplitudes <585 MPa), the Code mean curve predicts significantly longer fatigue lives than those observed experimentally for several heats of austenitic SSs with composition and tensile strength within the ASME specifications. The difference between the Code mean curve and the best-fit of the available experimental data is due most likely to differences in the tensile strength of the steels. The Code mean curve represents SSs with relatively high strength; the fatigue s-N data obtained during the last 30 years were obtained on SSs with lower tensile strengths.

Furthermore, for the current Code mean curve, the 106 cycle fatigue limit (i.e., the stress amplitude at a fatigue life of 106 cycles) is 389 MPa, which is greater than the monotonic yield strength of austenitic SSs in more common use (z303 MPa). Consequently, the current Code design curve for austenitic SSs I

does not include a mean stress correction for fatigue lives below 106 cycles. Recent studies by Wire et al. 1 12 and Solomon et al. 70 on the effectfof residual stress on fatigue life clearly demonstrate that mean stress can decrease the 106 cycle fatigue limit of the material; the extent of the effect depends on the cyclic hardening behavior of the material, and the resultant decrease in strain amplitude developed during load-controlled cycling. Strain hardening is more pronounced at high temperatures (e.g., 288-320'C) or at high mean stress (e.g., >70 MPa); therefore, as observed by Wire et al. and Solomon et al., fatigue life for load-controlled tests with mean stress is actually increased at high temperatures or large values of I

mean stress. In both studies, under load control, mean stress effects were observed at low temperatures (150'C) or at relatively low mean stress (<70 MPa). I Wire et al.1 12 performed fatigue tests on two heats of Types 304 SS to establish the effect of mean stress under both strain control and load control. The strain-controlled tests indicated "an apparent reduction of up to 26% in strain amplitude in the low- and intermediate-cycle regime (<106 cycle) for a mean stress of 138 MPa." However, the results were affected both by mean stress and cold work.

Although the composition and vendor-supplied tensile strength for the two heats of Type 304 SS were within the ASME specifications, the measured mechanical properties showed much larger variations than indicated by the vendor properties. Wire et al. state, "at 288°C, yield strength varied from 152-338 MPa.

I These wide variations are attributed to variations in (cold) working from the surface to the center of the thick cylindrical forgings." After separating the individual effect of mean stress and cold work, the Wire et al. results indicate a 12% decrease in strain amplitude for a mean stress of 138 MPa. These results are I

consistent with the predictions based on conventional mean stress models such as the Goodman correlation.

• The current Code mean data curve, and therefore the Code design curve, is nonconservative with respect to the existing fatigue E-N datafor austenitic SSs. A new Code fatigue design curve, which is consistent with the existingfatigue data, has been proposed (see Section 5.1.8for details).

I 42 I

5.1.2 Specimen Geometry The influence of specimen geometry (hourglass vs. gauge length specimens) on the fatigue life of Types 304 and 316 SS is shown in Fig. 31. At temperatures up to 300'C, specimen geometry has little or no effect on the fatigue life of austenitic SSs; the fatigue lives of hourglass specimens are comparable to those of gauge specimens.

Type 316 SS I RT Best-Fit Air ANL Model 0 2900C 1.0 1.0 0*

to 0U)

Hourglass Specimens

0. 0.1 C01pen Symbols: Type 316 SS

- 0 osed Symbols: Type 316L SS 3 4 3 5 6 7 102 10 10 105 106 101 108 102 10 104 10 10 10 108 Fatigue Life (Cycles) Fatigue Life (Cycles)

Figure 31. Influence of specimen geometry on fatigue life of Types 304 and 316 stainless steel (JNUFAD data).

Fatigue c-N data obtained either on hourglass or straight gauge specimens may be used to develop the Codefatigue design curves.

5.1.3 Temperature The fatigue life of Types 304 and 316 SS in air at temperatures between 100 and 325°C is plotted in Fig. 32; the best-fit curve based on the ANL model (Eq. 32 in Section 5.1.7) and the ASME Code mean curve are also shown in the figures. In air, the fatigue life of austenitic SSs is independent of temperature from room temperature to 400'C. Although the effect of strain rate on fatigue life seems to be significant

. ..I. .. . . ... . .. .. F II.

. . ..Ij . . ...... I. . r .i ..I Type 304 SS Type 316 SS Air~r Air V 100°C ** * . an2Curv n 290D'C L 260COC

- 00 2881C 2 1.0 > 300°0

- ASME Code 0 1.0 0

0 325 "C .ASME Code

_ " Mean Curve

.-. rMean Curve E - Best-Pit Air U*l E ANL Model m- .. Best-Fit Air - .

ANL Model -- - --

0) 0.1 0.1 I.. I. I. .... -... I

. . . .I . . . . I . . . . .I . . ......1 I. . II

,,I. . .. . . ... .

2 3 7 2 3 5 7 10 10 104 105 106 10 108 10 10 104 10 106 10 108 Fatigue Life (Cycles) Fatigue Life (Cycles)

Figure 32. Influence of temperature on fatigue life of Types 304 and 316 stainless steel in air (Ref. 38, JNUFAD database).

43

I at temperatures above 400 0 C, variations in strain rate in the range of 0.4-0.008%/s have no effect on the I

fatigue lives of SSs at temperatures up to 4000C. 69 In air, the fatigue s-N data can be represented by a single curve for temperatures from room temperature Up to 400 0 C. I Recent data indicate that temperature can influence the fatigue limit of austenitic SSs because of differences in the secondary hardening behavior of the material due to dynamic strain aging. 7 1 For a heat of Type 304L SS, the fatigue limit was higher at 300'C than at 150'C because of significant secondary I

hardening at 300'C.

• Temperature has no significant effect on the fatigue life of austenitic SSs at temperaturesfrom room temperature to 400'C. Variations in fatigue life due to the effects of secondary hardening behavior are I

accountedfor in the factor applied on stress to obtain the design curvefrom the mean data curve.

5.1.4 Cyclic Strain Hardening Behavior I

Under cyclic loading, austenitic SSs exhibit rapid hardening during the first 50-100 cycles; as shown in Fig. 33 the extent of hardening increases with increasing strain amplitude and decreasing I temperature and strain rate. 3 8 The initial hardening is followed by a softening and saturation stage at high temperatures, and by continuous softening at room temperature.

• The cyclic strain hardening behavior is likely to influence the fatigue limit of the material; variations I

I infatigue life due to such effects are accountedfor in thefactor of 2 on stress.

Type 316NG SS Strain Amplitude (%) Type 304 SS Strain Amplitude (%)

t*

300 250 288°C Air Strain Rate (%Is)

Open Symbols: 0.5 Closed Symbols: 0.005 a' < 0.50 0.75

)

-

300 -

288°C Air a 007 0.50 I

7a 250 o a E 200 S.. ... &. E 200 -

AA a.'

a A

-

I U) t 150 100*

I- 150 10 ý Strain Rate (%/s)

Open Symbols: 0.4 Closed Symbols: 0.004

.. .. I . . ."..

'.. ý. . .. . I . . - ..'

A

. . ....-J . . . ... I I

2 3 4 3 10 10 105 10 6 I

100 101 10 100 lot 102 10 104 5 10 106 Number of Cycles Number of Cycles 35 0 . . .. . . . . I . . . ..... . . . ..I. 1..

..1 1. . ...... I 3C : a v Type 316NG SS I

v,. Room Temperature Air 3001 -:,

Figure 33.

250 ouu. Effect of strain amplitude, temperature, and strain rate on cyclic strain-hardening behavior of ad Strain Amplitude (%)

A CL 200 n 1."

Types 304 and 316NG SS in air.

u)) - 0.50 150 - 0.35 I

0.27 Strain Rate: 0.17 - 1.0 %Is IG L . . ..I . I. . ..... . . . I.... .I ...... . . II.. . . . .. .

0 2 3 4 10 101 10 10 10 105 106 Number of Cycles I

44

5.1.5 Surface Finish Fatigue tests have been conducted on Types 304 and 3 16NG SS specimens that were intentionally roughened in a lathe, under controlled conditions, with 50-grit sandpaper to produce circumferential cracks with an average surface roughness of 1.2 gin. The results are shown in Figs. 34a and b, respectively, for Types 316NG and 304 SS. For both steels, the fatigue life of roughened specimens is a factor ofz3 lower than that of the smooth specimens.

- The effect of surfacefinish was not investigated in the mean data curve used to develop the Code fatigue design curves; it is included as part of the subfactor that is applied to the mean data curve to accountfor "surfacefinish and environment."

Type 316NG SS Heat D432804 Heat P91576 Type 304 SS Sawtooth Waveform 289°C r A Air 289°C Strain Rate = 0.004/0.4%/s 1.0 7 1.0-.

Z7 1.0Strain Rate. 0.4 - 0.0040/as a ~ .A -a

• . ASME Code 12AStAE Code m Design Curve Design Curve 0.1 - .

Open Symbols: Smooth Specimens Open Symbols: Smooth Specimens Closed Symbols: Rough Surface, 50 grit paper Closed Symbols: Rough Surface, 50 grit paper

_____________________ I________1_____I_____________________

11__ 1_1_i_______ iil__i____ i I , I , I II 103 104 105 106 103 104 105 106 Fatigue Life (Cycles) Fatigue Life (Cycles)

(a) (b)

Figure 34. Effect of surface roughness on fatigue life of (a) Type 316NG and (b) Type 304 SSs in air.

5.1.6 Heat-to-Heat Variability The effects of material variability and data scatter must be included to ensure that the design curves not only describe the available test data well, but also adequately describe the fatigue lives of the much larger number of heats of material that are found in the field. As mentioned earlier for carbon and low-alloy steels, material variability and data scatter in the fatigue s-N data for austenitic SSs are also evaluated by considering the best-fit curves determined from tests on individual heats of materials or loading conditions as samples of the much larger population of heats of materials and service conditions of interest. The fatigue behavior of each of the heats or loading conditions is characterized by the value of the constant A in Eq. 6. The values of A for the various data sets were ordered, and median ranks were used to estimate the cumulative distribution of A for the population. The distributions were fit to lognormal curves. Results for various austenitic SSs in air are shown in Fig. 35. The median value of the constant A is 6.891 for the fatigue life of austenitic SSs in air at temperatures not exceeding 400'C. The values of A that describe the 5th percentile of these distributions give a fatigue s-N curve that is expected to bound the lives of 95% of the heats of austenitic SSs. A Monte Carlo analysis was performed to address the uncertainties in the median value and standard deviation of the sample used for the analysis.

For austenitic SSs, the values for A that provide bounds for the portion of the population and the confidence that is desired in the estimates of the bounds are summarized in Table 8. From Fig. 35, the median value of A for the sample is 6.891. From Table 8, the 95/95 value of the margin to account for material variability and data scatter is 2.3 on life. This margin is needed to provide reasonable confidence that the resultant life will be greater than that observed for 95% of the materials of interest.

45

I 1.0 ---Austenitic. .SSs . I .. .... . .. ./ ' ! .."... .... .

I 0.8

--Air

-.................

...........

75th Percentile-

.......... . ........

...

---

....

_ ............

--- .............. I t1L C:

0 .............--

I  : ;I4;::*

- ------ ----- * - -

I .  !

I 0.6 "*a*-*:8*!:::::* ::::--I~~...... ..........

• • O aD~*s ..

Median 6.891

. 1} 357 Data Points 38 Heats.

Figure 35.

Estimated cumulative distribution of constant A I in the ANL model for fatigue life for heats of 0.4 > 316N-1

.- austenitic SS in air.

I

, ... .. 316N-A

___0

~' 304-3 304-10 Pece i .. I X 304-21 Percen 'il-- e-_[--,-- V 304-A 0.2 - *i

• 304-G I

.... ......-.

. 316-1 S "< 316-3

.. ,i,..... ........

. Z

_ _............._ 316-12

_ __- _

n nL 5.5 6 6.5 7 Constant A 7.5 8 8.5 9 I

Table 8. Values of parameter A in the ANL fatigue life model and the margins on life for austenitic Confidence SSs in air as a function of confidence level and percentage of population bounded, Percentage of Population Bounded (Percentile Distribution of A)

I Level 95(5) 90(10) 75(25) 67(33) 50(50) 50 75 6.205 6.152 6.356 6.309 Values of Parameter A 6.609 6.569 6.707 6.668 6.891 6.851 I

95 6.075 6.241 6.510 6.611 6.793 50 75 2.0 2.1 1.7 1.8 Margins on Life 1.3 1.4 1.2 1.2 1.0 1.0 I

95 2.3 1.9 1.5 1.3 1.1

• The Codefatigue design curves are based on the mean data curves; heat-to-heat variability is included I

in the subfactor that is applied to the mean data curve to account for "data scatter and material variability."

I 5.1.7 Fatigue Life Model The database used to develop the new air mean data curve is much larger and developed for more representative materials than were used as the basis for the existing ASME fatigue design curves. It is an I

updated version of the PVRC database; the sources are listed in Table 1 of the present report. The data were obtained on smooth specimens tested under strain control with a fully reversed loading (i.e., R = -1) in compliance with consensus standard approaches for the development of such data. The database for austenitic SSs consists of some 520 tests on Types 304, 316, 304L, 316L and 316NG SS; Z220 for Type 304 SS; 150 for Type 316 SS; and 150 for Types 316NG, 304L, and 316L SS. The austenitic SSs used in these studies are all in compliance with the compositional and strength requirements of the ASME I

Code specifications. I 46 I I

Several different best-fit mean &-N curves for austenitic SSs have been proposed in the literature.

Examples include Jaske and O'Donnell, 72 Diercks,1 13 Chopra, 38 Tsutsumi et al., 2 8 and Solomon and Amzallag. 1 14 These curves differ by up to 50%, particularly in the 10 4 to 107 cycle regime; the differences primarily occur because different database were used in developing the models for the mean

&-N curves. The analyses by Jaske and O'Donnell and by Diercks are based on the Jaske and O'Donnell database. The details regarding the database used by Tsutsumi et al. are not available. The database used in NUREG/CR-5704 included the Jaske and O'Donnell data, data obtained in Japan (including the JNUFAD database), and some additional data obtained in the U.S. In the earlier ANL reports, separate models were presented for Type 304 or 316 SS and Type 316NG SS. In the present report, the existing data were reanalyzed to develop a single model for the fatigue c-N behavior of austenitic SSs. The model assumes that the fatigue life in air is independent of temperature and strain rate. Also, to be consistent with the models proposed by Tsutsumi et al. 28 and Jaske and O'Donnell, 72 the value of the constant C in the modified Langer equation (Eq. 6) was lower than that in earlier reports (i.e., 0.112 instead of 0.126).

The proposed curve yields an R 2 value of 0.851 when compared with the available data; the R2 values for the mean curves derived by Tsutsumi et al., Jaske and O'Donnell, and the ASME Code are 0.839, 0.826, and 0.568, respectively.

In air, at temperatures up to 400'C, the fatigue data for Types 304, 304L, 316, 316L, and 316NG SS are best represented by the equation:

ln(N) = 6.891 - 1.920 ln(ca - 0. 112) (32) where F, is applied strain amplitude (%). The experimental values of fatigue life and those predicted by Eq. 32 for austenitic SSs in air are plotted in Fig. 36. The predicted lives show good agreement with the experimental values; for most tests the difference between the experimental and predicted values is within a factor of 3, and for some, the observed fatigue lives are significantly longer than the predicted values.

The ANL fatigue life models represent mean values offatigue life. The effects ofparameters such as mean stress, surfacefinish, size and geometry, and loading history, which are known to influence fatigue life, are not explicitly considered in the model; such effects are accountedfor in the factors of 20 on life and 2 on stress that are applied to the mean data curve to obtain the Codefatigue design curve.

Austenitic SSs Ausienitic S5s 21-325°C Air 21-325°Air Z -

106 106 2-2' i

- - . - 14°- - - -- - - - - - .,

., co,  ! 'o:2a .

104 ~0 -0 101 0~0 103 --- _

..---- .- __ ._ ................_

.................... ....... + 1032 - ---

102 --------- 102A '04LSS 304SS 0 316L S 0 316SS Ss-E~

101 "101 2 4

101 102 103 10 10i 106 101 10 103 104 105 106 Observed Life (Cycles) Observed Life (Cycles)

Figure 36. Experimental and predicted fatigue lives of austenitic SSs in air.

47

I 5.1.8 New Fatigue Design Curve I

As discussed in Section 5.1.1, the current Code mean-data curve that was used to develop the Code fatigue design curve, is not consistent with the existing fatigue c-N data. A fatigue design curve that is I

consistent with the existing database may be obtained from the ANL model (Eq. 32) by following the same procedure that was used to develop the current ASME Code fatigue design curve. However, the discussions presented later in Section 7.5 indicate that the current Code requirement of a factor of 20 on cycles, to account for the effects of material variability and data scatter, specimen size, surface finish, and I

loading history, is conservative by at least a factor of 1.7. Thus, to reduce this conservatism, fatigue design curve based on the ANL model for austenitic SSs (Eq. 32) may be developed by first correcting for mean stress effects using the modified Goodman relationship and then lowering the mean-stress-adjusted I

curve by a factor of 2 on stress or 12 on cycles, whichever is more conservative. This curve and the current Code design curve are shown in Fig. 37; values of stress amplitude vs. cycles for the current and the proposed design curves are given in Table 9. A fatigue design curve that is consistent with the I

existing fatigue c-N data but is not based on the ANL model (Eq. 32) has also been proposed by the 89 ASME Subgroup on Fatigue Strength.

Table 9. The new and current Code fatigue design curves for austenitic stainless steels in air.

I Cycles Stress Amplitude (MPa/ksi)

New Design Curve Current Design Curve Cycles Stress Amplitude (MPa/ksi)

New Design Curve Current Desin Curve I

I E+01 6000 (870) 4881 (708) 2 E+05 168 (24.4) 248 (35.9) 2 E+01 5 E+01 1 E+02 2 E+02 4300(624) 2748 (399) 1978 (287) 1440(209) 3530(512) 2379 (345) 1800 (261) 1386(201) 5 E+05 I E+06 2 E+06 5 E+06 142 (20.6) 126(18.3) 113 (16.4) 102 (14.8) 214 (31.0) 195 (28.3) 157 (22.8) 127(18.4)

I 5 E+02 974(141) 1020(148) I E+07 99(14.4) 113(16.4) 1 E+03 2 E+03 5 E+03 745 (108) 590 (85.6) 450 (65.3) 820 (119) 669 (97.0) 524 (76.0) 2 E+07 5 E+07 I E+08 97.1 (14.1) 105 (15.2) 98.6(14.3) 97.1 (14.1)

I 1 E+04 368 (53.4) 441 (64.0) I E+09 95.8 (13.9) 95.8 (13.9) 2 E+04 5 E+04 1 E+05 300 (43.5) 235 (34.1) 196 (28.4) 383 (55.5) 319(46.3) 281 (40.8)

I E+10 I E+ll 2 E+10 94.4 (13.7) 93.7 (13.6) 94.4 (13.7) 93.7 (13.6) I I

103 I E

I 65 10~2 101 .102 103 104 105 106 107 108 10 9

1010 1011 I

Figure 37.

Number of Cycles N Fatigue design curve for austenitic stainless steels in air.

I 48 I I

The proposed curve extends up to 1011 cycles; the two curves are the same beyond 108 cycles.

Although the curve is based primarily on data for Types 304 and 316 SS, it may be used for wrought Types 304, 310, 316, 347, and 348 SS, and cast CF-3, CF-8, and CF-8M SS for temperatures not exceeding 371°C (700'F).

9 The current Code fatigue design curve for austeniticstainless steels is nonconservative with respect to the existingfatigue E-N datafor fatigue lives in the range of 103 to 5 x 106 cycles. A new design curve, that is consistent with the existing data, has been developed. To reduce the conservatism in the current Code requirement of 20 on life, the new curve was obtained by usingfactors ofe2 on life and 2 on stress.

5.2 LWR Environment 5.2.1 Experimental Data The fatigue lives of austenitic SSs are decreased in LWR environments; the fatigue s-N data for Types 304 and 316NG SS in water at 288°C are shown in Fig. 38. The c-N curves based on the ANL model (Eq. 32 in Section 5.1.7 and Eq. 34 in Section 5.2.13) are also included in the figures. The fatigue life is decreased significantly when three threshold conditions are satisfied simultaneously, viz., applied strain range and service temperature are above a minimum threshold level, and the loading strain rate is below a threshold value. The DO level in the water and, possibly, the composition and heat treatment of the steel are also important parameters for environmental effects on fatigue life. For some steels, fatigue life is longer in high-DO water than in low-DO PWR environments. Although, in air, the fatigue life of Type 316NG SS is slightly longer than that of Types 304 and 316 SS, the effects of LWR environments are comparable for wrought Types 304, 316, and 316NG. Also, limited data indicate that the fatigue life of cast austenitic SSs in both low-DO and high-DO environments is comparable to that of wrought SSs in low-DO environment.

. . ...'.......I I .. . 1. ...- , . ., ' '. . I I . 1 . . -1

. ....

MI . . II I Type 304 SS Tensile/Compressiwe Type 316NG 288'C Water Strain Rate (%/s) 288°C Water 0.005 ppm DO 8ppm DO

" 004

. - Best-Fit Air C0004 7N S1.0 - - ANLModel [I 1.0 Best-Fit Air F ..

ANI Modl a.

) ANL Model S

"i "* Tensile/Compreissive >- - .

U) Strain Rate (%/s)ve 288°C Low 00 Water  :. . 4 288°C High DO Water 0.1 0.004%/s Strain Rate 0.1 'D 0,04.0.04 0.04%Is Strain Rate 3 5 3 4 5 102 10 104 10 106 102 10 10 10 106 Fatigue Life (Cycles) Fatigue Life (Cycles)

(a) (b)

Figure 38. Strain amplitude vs. fatigue life data for (a) Type 304 and (b) Type 316NG SS in water at 288°C (JNUFAD and Refs. 7,38).

The existing fatigue data indicate that a slow strain rate applied during the tensile-loading cycle (i.e., up-ramp with increasing strain) is primarily responsible for the environmentally assisted reduction in fatigue life. Slow rates applied during both tensile- and compressive-loading cycles (i.e., up- and down-ramps) do not further decrease fatigue life compared with that observed for tests with only a slow 49

Figure 39.

(a) (b)

Higher-magnification photomicrographs of oxide films that formed on Type 316NG stainless 1

steel in (a) simulated PWR water and (b) high-DO water.

tensile-loading cycle (Fig. 38b). Consequently, loading and environmental conditions during the tensile-loading cycle (strain rate, temperature, and DO level) are important for environmentally assisted reduction of the fatigue lives of these steels. 3 For austenitic SSs, lower fatigue lives in low-DO water than in high-DO water are difficult to reconcile in terms of the slip oxidation/dissolution mechanism, which assumes that crack growth rates increase with increasing DO in the water. The characteristics of the surface oxide films that form on austenitic SSs in LWR coolant environments can influence the mechanism and kinetics of corrosion U

processes and thereby influence the initiation stage, i.e., the growth of MSCs. Also, the reduction of fatigue life in high-temperature water has often been attributed to the presence of surface micropits that I may act as stress raisers and provide preferred sites for the fonnation of fatigue cracks.

Photomicrographs of the gauge surfaces of Type 3 16NG specimens tested in simulated PWR water and high-DO water are shown in Fig. 39. Austenitic SSs exposed to LWR environments develop an oxide film that consists of two layers: a fine-grained, tightly-adherent, chromium-rich inner layer, and a crystalline, nickel-rich outer layer composed of large and intermediate-size particles. The inner layer forms by solid-state growth, whereas the crystalline outer layer forms by precipitation or deposition from the solution. A schematic representation of the surface oxide film is shown in Fig. 40. The structure and composition of the inner and outer layers and their variation with the water chemistry have been identified.] 15,116 Large-size Particles Intermediate-size Particles Outer Layer Outer Layer I

Figure 40. Schematic of the corrosion oxide film formed on austenitic stainless steels in LWR environments.

5 50 I

Experimental data indicate that surface micropits or minor differences in the composition or structure of the surface oxide film have no effect on the formation of fatigue cracks. Fatigue tests were conducted on Type 316NG (Heat P91576) specimens that were preexposed to either low-DO or high-DO water and then tested in air or water environments. The results of these tests, as well as data obtained earlier on this heat and Heat D432804 of Type 316NG SS in air and low-DO water at 288'C, are plotted in Fig. 41. The fatigue life of a specimen preoxidized in high-DO water and then tested in low-DO water is identical to that of specimens tested without preoxidation. Also, fatigue lives of specimens preoxidized at 288°C in low-DO water and then tested in air are identical to those of unoxidized specimens (Fig. 41).

If micropits were responsible for the reduction in life, the preexposed specimens should show a decrease in life. Also, the fatigue limit of these steels should be lower in water than in air, but the data indicate this limit is the same in water and air environments. Metallographic examination of the test specimens indicated that environmentally assisted reduction in fatigue lives of austenitic SSs most likely is not 7 36 37 caused by slip oxidation/dissolution but some other process, such as hydrogen-induced cracking. , ,

• An LWR environment has a significant effect on the fatigue life of austenitic SSs; such effects are not considered in the current Code design curve. Environmental effects may be incorporated into the Code fatigue evaluation using the Fen approachdescribed in Section 5.2.14.

5.2.2 Strain Amplitude As in the case of the carbon and low-alloy steels, a minimum threshold strain range is required for the environmentally induced decrease in fatigue lives of SS to occur. Exploratory fatigue tests have also been conducted on austenitic SSs to determine the threshold strain range beyond which environmental effects are significant during a fatigue cycle. 24 ,29 The tests were performed with waveforms in which the slow strain rate is applied during only a fraction of the tensile loading cycle. The results indicate that a minimum threshold strain is required for an environmentally assisted decrease in the fatigue lives of SSs (Fig. 42). The threshold strain range Asth appears to be independent of material type (weld or base metal) 24 29 and temperature in the range of 250-3251C, but it tends to decrease as the strain range is decreased. ,

The threshold strain range may be expressed in terms of the applied strain range Ae by the equation AF-th/AE = - 0.22 AE + 0.65. (33)

The results suggest that A-th is related to the elastic strain range of the test and does not correspond to the strain at which the crack closes.

Type 316NG SS Open Symbols: Air

_Figure 289oC Closed Symbols: Low-DO water 41.

Heat D432804 Effects of environment on formation of 1.0 A 0- 0.4%/s 0.004%/s Best-Fit Air fatigue cracks in Type 316NG SS in air

" Heat 91576 ANL Model and low-DO water at 2880C.

Preoxidized Preoxidized specimens were exposed E Low-DO

> 0.4%/s-. for 10 days at 288'C in water that Preoxidized contained either <5 ppb DO and o,9 High-DO =23 cm 3 /kg dissolved H 2 or =500 ppb DO and no dissolved H2 (Ref. 7).

103 104 105 Fatigue Life (Cycles) 51

• I Type 316 SS, 325°C Strain Range Ac = 1.2%

DO =0.005 ppmn 103 G

0I Figure 42.

Results of strain rate change tests on 0

o Type 316 SS in low-DO water at 325°C. Low strain rate was applied during only a fraction of 0o - tensile loading cycle. Fatigue life is plotted as

".0 a function of fraction of strain at high strain rate Threshold Strain (Refs. 24,29).

102 I .

0.0 0.2 0.4 0.6 0.8 1.0 Aefast Ac In LWR environments, the procedureforcalculatingFen, defined in Eq. 38 (Section 5.2.14), includes a threshold strain range below which LWR coolant environments have no effect on fatigue life, i.e., Fe, = 1.

However, a threshold strain should not be considered when the damage rate approach is used to determine Fen for a stress cycle or load set pair.

5.2.3 Hold-Time Effects Environmental effects on fatigue life occur primarily during the tensile-loading cycle and at strain levels greater than the threshold value. Information on the effect of hold periods on the fatigue life of austenitic SSs in water is very limited. In high-DO water, the fatigue lives of Type 304 SS tested with a trapezoidal waveform (i.e., hold periods at peak tensile and compressive strain) 8 are comparable to those tested with a triangular waveform, 2 5 as shown in Fig. 43. As discussed in Section 4.2.8, a similar behavior has been observed for carbon and low-alloy steels: the data show little or no effect of hold periods on fatigue lives of the steels in high-DO water.

I Type 304 SS High-DO Water 1.0 - d O -\ Best-FitAir ANL Model Figure 43.

0 oeFatigue life of Type 304 stainless steel

-a ---. 0 tested in high-DO water at 260-288°C E

< ASME Design Curve 0.

9069

-

with trapezoidal or triangular waveform (Refs. 8,25).

I 0 260'C, 0.2 ppm, =0.03%/s . es ,

Trapezoidal Waveform 0.1 0 288°C, 8 ppm DO, =0.04%/s Triangular Waveform 1,,,1 I . . . ..... I , i , 1,, 1,,r.. . .. I . . . .

3 106 102 10 104 105 Fatigue Life (Cycles)

The existing data do not demonstrate that hold periods at peak tensile strain affect the fatigue life of austenitic SSs in LWR environments. Thus, any revision/modification of the method to determine Fei is not warranted.

52

5.2.4 Strain Rate The fatigue life of Types 304L and 316 SSs in low-DO water is plotted as a function of tensile strain rate in Fig. 44. In low-DO PWR environment, the fatigue life of austenitic SSs decreases with decreasing strain rate below Z0.4%/s; the effect of environment on fatigue life saturates at -0.0004%/s (Fig. 4 4 ).7,18,21-25,28,29,38-40 Only a moderate decrease in life is observed at strain rates greater than 0.4%/s. A decrease in strain rate from 0.4 to 0.0004%/s decreases the fatigue life by a factor of l 0.

105 1 1 ý' " . , m ir"'1 - " q .

300°C; DO 0 - 0.1 ppm ............... 325°C; DO 0.005 ppm Type 304L SS Air 0.2 Open Symbols: Type 304 Air 0.25%

Closed Symbols: Type 316 -. Air 0.3% ......

1047,--* J..J*1" j

'a a) Air 0.50% a,

~0 /, f',

104 Estimated Air 0.6%.,*

0 a,

103 C a) 1037

3 05

• LLS o oSt rain Amplitude (%) ('3 - Strain Amplitude (%

U-0 0.50%

0 0.35%

102 Ls 0.30 102 A 0.25% 0 0.25 II HI ! , ~ ,I , , , ,,l,I , - - .,,n- -TJl II 'l , , , , , l , ,Ir i , , , Il ,

10-5 10-4 10-3 10-2 10-1 10S 10r5 10-4 t 10-2 Ra- 10-1 10o Strain Rate (%/s)

Strain Rate (%Is)

Figure 44. Dependence of fatigue lives of austenitic stainless steels on strain rate in low-DO water (Refs. 7,38,40,71).

In high-DO water, the effect of strain rate may be less pronounced than in low-DO water (Fig. 45).

For example, for Heat 30956 of Type 304 SS, strain rate has no effect on fatigue life in high-DO water, whereas life decreases linearly with strain rate in low-DO water (Fig. 45a). For Heat D432804 of Type 316NG, some effect of strain rate is observed in high-DO water, although it is smaller than that in low-DO water (Fig. 45b).

105- =: - ,

• Type 304 SS (Heat 30956) K Type 316NG SS (Heat D432804) 288°C ... ..............

.............

.. F 288°C A-D -

104[ a) 104ý -

a, L a) L 0) 15 a) .

Strain Amplitude L 103 Strain Amplitude 103 U-O =0.38% 1_

-

E) =0.4%

O =0.25% '0 =0.25%

Open Symbols: <0.005 ppm DO Open Symbols: <0.005 ppm DO Closed Symbols: =0.7 ppm DO Closed Symbols: >0.2 ppm DO 2

10 . L ,

5 5 3 10- 10-4 10-3 10-2 10-1 100 10- 10-4 10- 10-2 10-1 100 Strain Rate (%/s) Strain Rate (%/s)

(a) (b)

Figure 45. Dependence of fatigue life of Types (a) 304 and (b) 316NG stainless steel on strain rate in high- and low-DO water at 288°C (Ref. 7,38,40).

53

I In LWR environments, the effect of strain rate on the fatigue life of austenitic SSs is explicitly consideredin Fen defined in Eq. 38 (Section 5.2.14). Also, guidance is provided to define the strain rate to be used to calculateFen for a specific stress cycle or loadset pair.

5.2.5 Dissolved Oxygen 3 In contrast to the behavior of carbon and low-alloy steels, the fatigue lives of austenitic SSs decrease significantly in low-DO (i.e., <0.05 ppm DO) water. In low-DO water, the fatigue life is not influenced by the composition or heat treatment condition of the steel. The fatigue life, however, 7 18 23 25 2 8 29 3 8 0 continues to decrease with decreasing strain rate and increasing temperature. , , - , , , -4 In high-DO water, the fatigue lives of austenitic SSs are either comparable to 23 ,2 8 or, in some cases, higher 7,38,4 0 than those in low-DO water, i.e., for some SSs, environmental effects may be lower in high-DO than in low-DO water. The results presented in Figs. 45a and 45b indicate that, in high-DO water, environmental effects on the fatigue lives of austenitic SSs are influenced by the composition and heat treatment of the steel. For example, for high-carbon Type 304 SS, environmental effects in high-DO water are insignificant for the mill-annealed (MA) material (Fig. 45a), whereas as discussed in Section 5.2.8, for sensitized material the effect of environment is the same in high- and low-DO water. For the low-C Type 316NG SS, some effect of strain rate is apparent in high-DO water, although it is smaller than that in low-DO water (Fig. 45b). The effect of material heat treatment on the fatigue life of Type I

304 SS is discussed in Section 5.2.8; in high-DO water, material heat treatment affects the fatigue life of SSs. I o In LWR environments, the effect of DO on the fatigue life of austenitic SSs is explicitly considered in Fen, defined in Eq. 38. Also, guidance is provided to define the DO content to be used to calculateFen for a specific stress cycle or load set pair.

5.2.6 Water Conductivity The studies at ANL indicate that, for fatigue tests in high-DO water, the conductivity of water and the ECP of steel are important parameters that must be held constant.7, 38,40 During laboratory tests, the time to reach stable environmental conditions depends on the autoclave volume, DO level, flow rate, etc.

In the ANL test facility, fatigue tests on austenitic SSs in high-DO water required a soaking period of 5-6 days for the ECP of the steel to stabilize. The steel ECP increased from zero or a negative value to above 150 mV during this period. The results shown in Fig. 45a for MA Heat 30956 of Type 304 SS in high-DO water (closed circles) were obtained for specimens that were soaked for 5-6 days before the test. The same material tested in high-DO water after soaking for only 24 h showed a significant reduction in fatigue life, as indicated by Fig. 46.

The effect of the conductivity of water and the ECP of the steel on the fatigue life of austenitic SSs is shown in Fig. 46. In high-DO water, fatigue life is decreased by a factor of'2 when the conductivity of water is increased from z0.07 to 0.4 11S/cm. Note that environmental effects appear more significant for the specimens that were soaked for only 24 h. For these tests, the ECP of steel was initially very low and increased during the test.

5 54I

Type 304 SS 288°C Strain range =0.77%

Strain rate tensile 0.004%/s Air ............ & compressive 0.4 %/s DO =0.8 ppm 4? 104 Figure 46.

Effects of conductivity of water and soaking

'1)

LL period on fatigue life of Type 304 SS in high-DO water (Ref. 7,38).

Simulated PWR 0 Open Symbols: ECP 155 mV (=120 h soak)

Closed Symbols: ECP 30-145 mV (=24 h soak)

I n3 . .I . , I ,I . . I 10-2 10-1 100 Conductivity of Water (pS/cm) e Effects of water chemistry on fatigue life have not been considered in the determination of Fe,.

Additionalguidance may be neededfor excursions of off-normal water chemistry conditions.

5.2.7 Temperature The change in fatigue lives of austenitic SSs with test temperature at two strain amplitudes and two strain rates is shown in Fig. 47. The results suggest a threshold temperature of 150°C, above which the environment decreases fatigue life in low-DO water if the strain rate is below the threshold of 0.4%/s. In the range of 150-325 0 C, the logarithm of fatigue life decreases linearly with temperature. Only a moderate decrease in life occurs in water at temperatures below the threshold value of 1501C.

[ 4 Austenitic SSs c5 =0.6%, DO*0.005 ppm A 0.4%/s 0.01%/s (ni - - , Closed Symbols: Type 316 & 316NG 0

a) A 'A-..A #o

. ,j A CU)

Austenitic SSs 0- 3 -

U- ea = 0.3%, DO 0.005 ppm L1 Open Symbols: Type 304 0 0.4%/s A 0 -3 A Closed Symbols: Type 316 & 316NG An 0.01%/s 50 100 150 200 250. 300 350 400 50 100 150 200 250 300 350 400 Temperature (TC) Temperature (°C)

Figure 47. Change in fatigue lives of austenitic stainless steels in low-DO water with temperature (Refs.

7,23-25,28,38-40).

Fatigue tests have been conducted at MHI in Japan on Type 316 SS under combined mechanical and thermal cycling. 23 Triangular waveforms were used for both strain and temperature cycling. Two sequences were selected for temperature cycling: (i) an in-phase sequence, in which temperature cycling was synchronized with mechanical strain cycling, and (ii) a sequence in which temperature and strain were out of phase, i.e., maximum temperature occurred at minimum strain level and vice versa. Two temperature ranges, 100-325°C and 200-325°C, were selected for the tests. The results are shown in Fig. 48, along with data obtained from tests at constant temperature. An average temperature is used in 55

Fig. 48 for the thermal cycling tests. Because environmental effects are considered to be moderate below threshold values of 150'C for temperature and -0.25% for strain range, the average temperature for the thermal cycling tests was determined from higher value between 150'C and temperature at threshold strain for in-phase tests, and the lower value between maximum temperature and temperature at threshold strain for out-of-phase tests.

The results in Fig. 48 indicate that for load cycles involving variable temperature, average temperature gives the best estimate of fatigue life. Also, as expected, the fatigue lives of the in-phase tests are shorter than those for the out-of-phase tests. For the thermal cycling tests, fatigue life is longer for out-of-phase tests than for in-phase tests, because applied strains above the threshold strain occur at high temperatures for in-phase tests, whereas they occur at low temperatures for out-of-phase tests. The results from the thermal cycling tests (triangles) agree well with those from the constant-temperature tests (open circles).

104 Type 316 SS 325°C Ca=0.6%

DO = <0.005 ppm

• " Strain Rate 0.002%/s o _, - Figure 48.

I Stan ae 02/............s Fatigue life of Type 316 stainless steel under constant and varying test I,, temperature (Ref. 23).

'_ Temperature (Strain Rate, %/s)

• Constant (0.01)

A In phase (0.002)

• Out of phase (0.002) 10 2 . . . i.Ii l . ii l .l.

i ~I .l .ii i l 0 50 100 150 200 250 300 350 Temperature (°C)

Another study conducted by the' Japan Nuclear Safety Organization on Type 316 SS under combined mechanical and thermal cycling in PWR water showed similar results, e.g., the in-phase tests had lower fatigue lives than the out-of-phase tests. 30,32 These results indicate that load cycles involving variable temperature conditions may be represented by an average temperature.

9 In LWR environments, the effect of temperature on the fatigue life of austenitic SSs is explicitly considered in Fen, defined in Eq. 38 (Section 5.2.14). Also, guidance is provided to define the temperature to be used to calculate Fen for a specific stress cycle or load set pair.

5.2.8 Material Heat Treatment Limited data indicate that, although heat treatment has little or no effect on the fatigue life of austenitic SSs in low-DO and air environments, in a high-DO environment, fatigue life may be longer for nonsensitized or slightly sensitized SS.40 The effect of heat treatment on the fatigue life of Type 304 SS in air, BWR, and PWR environments is shown in Fig. 49. Fatigue life is plotted as a function of the EPR (electrochemical potentiodynamic reactivation) value for the various material conditions. The results indicate that heat treatment has little or no effect on the fatigue life of Type 304 SS in air and PWR environments. In a BWR environment, fatigue life is lower for the sensitized SSs; fatigue life decreases with increasing EPR value.

56

Type 304 Stainless Steel

- 289oC 2-

,1104, _ Figure 49.

The effect of material heat treatment on fatigue life of Type 304 stainless steel in air, BWR and PWR environments at 2890C, =0.38% strain

-amplitude, sawtooth waveform, and 0.004%/s 0 tensile strain rate (Ref. 40).

Strain amplitude =0.38% -- --- Air 103 Saw-tooth waveform 0 BWR Strain Rate 0.004%/s tensile -----A---- PWR 0.4%/s compressive 0 5 10 15 20 25 30 35 2

EPR (C/cm )

These results are consistent with the data obtained at MHI on solution-annealed and sensitized Types 304 and 316 SS. 2 1 ,25 In low-DO (<0.005 ppm) water at 325°C, a sensitization annealing had no effect on the fatigue lives of these steels. In high-DO (8 ppm) water at 300'C, the fatigue life of sensitized Type 304 SS was a factor of -2 lower than that of the solution-annealed steel. However, a sensitization anneal had little or no effect on the fatigue life of low-C Type 316NG SS in high-DO water at 288'C, and the lives of solution-annealed and sensitized Type 316NG SS were comparable.

• The effect of heat treatment is not considered in the environmentalfatigue correctionfactor; estimates of Fen based on Eq. 38 (Section 5.2.14) may be conservativefor some SSs in high-DO water.

5.2.9 Flow Rate It is generally recognized that flow rate most likely affects the fatigue life of LWR materials because it may cause differences in local environmental conditions in the enclaves of the microcracks formed during early stages in the fatigue c-N test. As discussed in Section 4.2.9, data obtained under typical operating conditions for BWRs indicate that environmental effects on the fatigue life of carbon 9

steels are a factor of z2 lower at high flow rates (7 mrns) than at low flow rates (0.3 m/s or lower).1 ,20 However, similar tests in both low-DO and high-DO environments indicate that increasing flow rate has no effect or may have a detrimental effect on the fatigue life of austenitic SSs. Figure 50 shows the effect of water flow rate on the fatigue life of Types 316NG and 304 SSs in high-purity water at 289°C. Under Austenitic 0 SSs 0.6% Rate: 0.001%/s 289 C Water

.... 104... 316NG SS

-A---- 304SS e.: 0.3% Rate: 0.01%/s

...... N.

.sGt' .. - iur.0

.... 'S- F igure 50 .

, A-. Effect of water flow rate on the fatigue life of 10: . ... - austenitic SSs in high-purity water at 289°C

. -- -- - - --. . (Ref. 20).

Dissolved Oxygen (ppm)

Open Symbols: 0.2 Closed Symbols: 0.05 L, I,,,,,,, ,,,,.l I,

, ,,d1 ,,,,,,,I , I I l 10-5 10-4 10-3 10-2 10-1 100 101 Flow Rate (m/s) 57

I all test conditions, the fatigue lives of these steels are slightly lower at high flow rates than those at lower rates or semi-stagnant conditions.

Fatigue tests conducted on SS pipe bend specimens in simulated PWR primary water at 240'C also indicate that water flow rate has no effect on the fatigue life of austenitic SSs. Increasing the flow rate from 0.005 m/s to 2.2 m/s had no effect on fatigue crack initiation in z26.5-mm diameter tube specimens.

These results appear to be consistent with the notion that, in LWR environments, the mechanism of I

fatigue crack initiation in austenitic SSs may differ from that in carbon and low-alloy steels.

9 Because of the uncertaintiesin theflow conditions at or near the locations of crack initiation and the insignificant effect offlow rate,flow rate effects on the fatigue life of austenitic SSs in L WR environments arepresently not consideredin the fatigue evaluations.

5.2;10 Surface Finish Fatigue tests have been conducted on Types 304 and 316NG SS specimens that were intentionally I roughened in a lathe, under controlled conditions, with 5-grit sandpaper to produce circumferential cracks with an average surface roughness of 1.2 .tm. The results are shown in Figs. 51 a and b, respectively, for Types 316NG and 304 SS. For both steels, the fatigue life of roughened specimens is lower than that of the smooth specimens in air and low-DOwater environments. In high-DO water, the fatigue life of Heat I

P91576 of Type 316NG is the same for rough and smooth specimens.

I I Type 316NG SS IIIIII I I I IIII III Heat D432804 II I III 1

Heat P91576 1 1 Type 304 SS I i i iiiii Sawtooth Waveform I

289°C N Air A Air 289°C Strain Rate = 0.004/0.4%/s 1.0 .. > PWR 0 PWR 1.0 C. Air V BWR 0 BWR - A Simulated PWR Water Strain Rate: 0.004%/s in Water -

-d a)V [>D "- "*., N 0.4 - 0.004%/s in Air - z.. "" e Best-Fit i Air Best-Fit Air .

E E7-

<- N- <

.9 ASME Code -. . . ASME Code -

0 . Design Curve - " 0.1 Design Curve Open Symbols: Smooth Specimens Open Symbols: Smooth Specimens Closed Symbols: Rough Surface, 50 grit paper Closed Symbols: Rough Surface, 50 grit paper 3 4 4 5 10 10 105 106 103 10 10 106 Fatigue Life (Cycles) Fatigue Life (Cycles)

(a)

(b)

Figure 51. Effect of surface roughness on fatigue life of (a) Type 316NG and (b) Type 304 stainless steels in air and high-purity water at 289°C.

The effect of surface finish is not considered in the environmental fatigue correction factor, it is included in the subfactorfor "surfacefinish and environment, " which is applied to the mean data curve to develop the Code fatigue design curve in air.

5.2.11 Heat-to-Heat Variability The effect of material variability and data scatter on the fatigue life of austenitic SSs has been evaluated for the data in LWR environments. The fatigue behavior of each of the heats or loading conditions is characterized by the value of the constant A in the ANL model (e.g., Eq. 6). The values of A for the various data sets are ordered, and median ranks are used to estimate the cumulative distribution of A for the population. The results in water environments are shown in Fig. 52. The median value of A I

58I

in water is 6.157. The results indicate that environmental effects are approximately the same for the various heats of these steels. For example, the cumulative distribution of data sets for specific heats is approximately the same in air and water environments. The ANL model seems to over-estimate the effect of environment for a few heats, e.g., the ranking for Type 304 SS heat 3 is z25 percentile in air (Fig. 35) and 485 percentile in water (Fig. 52).

The values for constant A that provide bounds for the portion of the population and the confidence that is desired in the estimates of the bounds for austenitic SSs in LWR environments are summarized in Table 10. In LWR environments, the 5th percentile value of Parameter A at a 95% confidence level is 5.401. Thus, for the median value of 6.157 for the sample (Table 10), the 95/95 value of the margin to account for material variability and data scatter is 2.3 on life. This margin is needed to provide 95%

confidence that the resultant life will be greater than that observed for 95% of the materials of interest.

1.0

-Austenitic SSs ... -

-- Water - - - __"

j- _

0.8 75th Percentile I OI __

U-0 i i F "

L L

.. 0.6 Median 6.157 Figure 52.

a)

___

__~___ * -* 276 Data Points-Estimated cumulative distribution of constant

-I 0.4 14 Heats A in the ANL model for fatigue life for heats of E . ..... ..... * - +, . ... . austenitic SSs in water.

0.3 2____ 1 ,.....:

Z4, h V 0 31N-A -

316N-1

-25th I- 0 304-3 -

Percen file - T @ 304-9 0.2 X 304-21 V 304-A

.... . .---

- o304L-E C) 316-14

/1 316-12 -

I 4.5 5 5.5 6 6.5 7 7.5 8 Constant A Table 10. Values of parameter A in the ANL fatigue life model and the margins on life for austenitic SSs in water as a function of confidence level and percentage of population bounded.

Confidence Percentage of Population Bounded (Percentile Distribution of A)

Level 95 (5) 90(10) 75 (25) 67(33) 50(50)

Values of Parameter A 50 5.481 5.630 5.880 5.976. 6.157 75 5.414 5.570 5.828 5.925 6.104 95 5.317 5.483 5.752 5.851 6.028 Margins on Life 50 2.0 1.7 1.3 1.2 1.0 75 2.1 1.8 1.4 1.3 1.1 95 2.3 2.0 1.5 1.4 1.1 The heat-to-heat variability is included in the Code fatigue design curves as part of the subfactor that is applied to the room-temperature mean data curve to account for "data scatter and material variability."

59

I 5.2.12 Cast Stainless Steels Available fatigue a-N data 2 3 ,28 ,37 ,38 indicate that, in air, the fatigue lives of cast CF-8 and CF-8M I SSs are similar to that of wrought austenitic SSs. The fatigue lives of cast austenitic SSs also decrease in LWR coolant environments. Limited data suggest that the fatigue lives of cast SSs in high-DO water are approximately the same as those in low-DO water. In LWR environments the fatigue lives of cast SSs are comparable to those of wrought SSs in low-DO water. Also, the fatigue lives of these steels are relatively insensitive to changes in ferrite content in the range of 12-28%.23,28 Also, existing data are inadequate to establish the dependence of fatigue life on temperature in LWR environments. I The effect of thermal aging at 250-400'C on the fracture toughness properties of cast SSs are well established, fracture toughness is decreased significantly after thermal aging because of the spinodal decomposition of the ferrite phase to form Cr-rich cc' phase. 117,118 The cyclic-hardening behavior of cast austenitic SSs is also influenced by thermal aging. 38 At 288°C, cyclic stresses of cast SSs aged for 10,000 h at 400'C are higher than those for unaged material or wrought SSs. Also, strain rate effects on cyclic stress are greater for aged than for unaged steel, i.e., cyclic stresses increase significantly with decreasing strain rate. The existing data are too sparse to establish the effects of thermal aging on strain-I rate effects on the fatigue life of cast SSs in air. Limited data in low-DO water at 288°C indicate that thermal aging for 10,000 h at 400'C decreases the fatigue life of CF-8M steels, Fig. 53b. 38 Note that thermal aging of another heat of CF-8M steel for 25,200 h at 465°C, Fig. 53a, had little or no effect on I

fatigue life. The different behavior for the two steels may be attributed to differences in the microstructure produced after thermal aging at 400'C as apposed to 465°C. Thermal aging at 400'C results in spinodal decomposition of the ferrite phase which strengthens the ferrite phase and increases cyclic hardening. Thermal aging at 465°C results in the nucleation and growth of large &x' particles and other phases such as sigma phase, which do not change the tensile or cyclic hardening properties of the material.

CF-8M Cast SS (FN 19.7) CF-8M Cast SS I 325°C; DO 0.005 ppm Heat 74 Ferrite =18%

Strain Amplitude (%) Heat 75 Ferrite =28%

0.6

• A 0.3 ..

0 0.25 a) Open Symbols: Aged 25.200 h at 465'C 0 '. -

103 Closed Symbols: Unaged n

IL I_

288°C: Ea =0.38%

DO <0.005 ppm

• 74 Unaged I

2102 74 Aged 10,000 h at 400°C 10 75 Aged 10,000 h at 400'C 3 4 1 0 -6 10-5 10-4 10- 10-2 10-1 100 10 -6 10-5 10- 10-3 10-2 10-1 100 Strain Rate (%Is) Strain Rate (%Is)

(a) (b)

Figure 53. Dependence of fatigue lives of CF-8M cast SSs on strain rate in low-DO water.at various strain amplitudes (Refs. 23,28,37,38).

The decrease in fatigue life with decreasing strain rate for three heats of CF-8M cast SS in low-DO water at 325 and 288°C is shown in Fig. 53; the effects of strain rate on the fatigue life of cast SSs are similar to those for wrought SSs. However, for an unaged heat of CF-8M steel with -20% ferrite, environmental effects on life do not appear to saturate even at strain rates as low as 0.00001%/s.23,28 Similar results have also been reported for unaged CF-8M steels in low-DO water at 325°C.1 19 Based 60

on these results, the saturation strain rate of 0.0004%/s, recommended for wrought SSs (Eq. 36 in Section 5.2.13), has been decreased to 0.00004%/s for cast SS. However, thermal aging may have influenced the results at very low strain rates. All of the tests at low strain rates were obtained on unaged material; as discussed above, available data indicate that thermal aging decreases the fatigue life of CF-8M steel (Fig. 53b). Limited data indicate that the effects of strain rate are the same in low- and high-DO water.

Also, such low strain rates (i.e., less than 0.0004%/s) are not likely to occur in the field. In the present report the effects of strain rate and temperature on the fatigue life of cast austenitic SSs are assumed to be similar to those for wrought SSs.

The estimated cumulative distribution of constant A in the ANL model for fatigue life for austenitic SSs, including several heats of cast SSs, in air and water environments are shown in Fig. 54. The results for cast SSs are evenly distributed and have insignificant effect on the median value of the constant A, e.g., the values with and without the cast SS data are 6.878 and 6.891, respectively, in air, and 6.147 and 6.157, respectively, in water. Thus, the ANL model for austenitic SSs adequately represent both wrought and cast SSs.

1 ._ A .i. ..- .. . . ..... ... .......

.......... . -..-W.. at.e r . .. ..

--Austenitic S5s ---- -AusteniticSSss-

. . ...............

0.8

• *e~~~~rc Z117- _eln

-5*t---h_ ,..

... _ ----:ii-----I

---

-i~L7*-i ----. 75--P-re~/~~ *.... --* *°-

U 75th Percentile 1 . hPretl .

-_0.6o...-. -~. . . .~

. ... - - --. . . . .-- -----

. .. . ..1.... . ..---------

-.

0.6. . . . .t .. . *

-Median 687-8 -. . . ". ..Median 6,147 -.

--- ----..---.--..---.--- -.-.----....

.50.4 -------

E E ----

--

.....

....... ... . ...... ... . ..... . .... ...

,....... .... ......

ercen, ie Percen tle - - .

0.2 - Cf_8M 0.2 . .... a... CFSM --

S - V 316NG .-.---.---- V 316NG -

.. . ... .....

S 014SS -.... .. A 304SS

. 316*s _ 316SS 00-~" "" . .... ..... 316 --

5.5 6 6.5 7 7.5 8 8.5 9 4.5 5 5.5 6 6.5 7 7.5 8 Constant A Constant A (a) (b)

Figure 54. Estimated cumulative distribution of constant A in the ANL model for fatigue life of wrought and cast austenitic stainless steels in (a) air and (b) water environments.

5.2.13 Fatigue Life Model In LWR environments, the fatigue life of austenitic SSs depends on strain rate, DO level, and temperature; the effects of these and other parameters on the fatigue life of austenitic SSs are discussed in detail below. The functional forms for the effects of strain rate and temperature are based on the data trends. For both wrought and cast austenitic SSs, the model assumes threshold and saturation values of 0.4 and 0.0004%/s, respectively, for strain rate and a threshold value of 150'C for temperature.

The influence of DO level on the fatigue life of austenitic SSs is not well understood. As discussed in Section 5.2.5, the fatigue lives of austenitic SSs are decreased significantly in low-DO water, whereas in high-DO water they are either comparable or, for some steels, higher than those in low-DO water. In 61

~I high-DO water, the composition and heat treatment of the steel influence the magnitude of environmental effects on austenitic SSs. Until more data are available to clearly establish the effects of DO level on fatigue life, the effect of DO level on fatigue life is assumed to be the same in low- and high-DO water and for wrought and cast austenitic SSs.

The least-squares fit of the experimental data in water yields a steeper slope for the F-N curve than i the slope of the curve obtained in air. 38,8 2 These results indicate that environmental effect may be more pronounced at low than at high strain amplitudes. Differing slopes for the E-N curves in air and water environments would add complexity to the determination of the environmental fatigue correction factor Fen, discussed in the next section. In the ANL model, the slope of the *-N curve is assumed to be the I

same in LWR and air environments. In LWR environments, fatigue data for austenitic SSs are best represented by the equation:

In(N) = 6.157 - 1.920 ln(F-a - 0. 112) + T' C' O', (34) where T', E, and 0' are transformed temperature, strain rate, and DO, respectively, defined as follows: i T'= 0 T'= (T-T' 1 150)/175 (T < 150-C)

(150 < T < 325-C)

(T _>325-C) (35)

I S=0 (t > 0.4%/s) ln( 2/0.4) (0.0004 < t < 0.4%/s)

'= ln(0.0004/0.4) (t < 0.0004%/s) (36) 0'= 0.281 (all DO levels). (37) i These models are recommended for predicted fatigue lives :S106 cycles. Note that Eq. 34 is based on the updated ANL model for austenitic SSs in air (Eq. 32) and the analysis presented in Section 5.2.11.

A single expression is used for Types 304, 304L, 316, 316L, and 316NG SSs, and constant A and slope B i

in the equation are different from the values reported earlier in NUREG/CR-5704, -6815, and -6878.

Equations 34-37 can also be used for cast austenitic SSs such as CF-3, CF-8, and CF-8M. Also, because the influence of DO level on the fatigue life of austenitic SSs may be influenced by the material composition and heat treatment, the ANL fatigue life model may be somewhat conservative for some SSs in high-DO water.

The experimental values of fatigue life and those predicted by Eq. 34 for austenitic SSs in LWR environments are plotted in Fig. 55. The predicted fatigue lives show good agreement with the experimental values. The difference between the experimental and predicted values is within a factor of 3 for most tests; the experimental fatigue lives of a few tests on Type 304 SS are up to a factor of z4 lower I

than the predicted values, all of these tests were on tube specimens with 1- or 3-mm wall thickness.

• The ANL model represent the mean values offatigue life as a function of applied strain amplitude, temperature, strain rate, and DO level in water. The effects ofparameters such as mean stress, surface I

finish, size and geometry, and loading history, which are known to influencefatigue life, are not included in the model.

62

106

(*10510

.....

~0 '

4? -----

101 . 101 a 10 A V0, 0 0 1 102 --------- -- 345 103 0 5.2.4 Enionetl 3 oretoFco 4 5 3 f igue 102 10 1re whic102 10 106 a 101 10 10efe 106 Observed Life (Cycles) Observed Life (Cycles)

Figure 55. Experimental and predicted values of fatigue lives of austenitic SSs in LWR environments.

5.2.14 Environmental Correction Factor The effects of reactor coolant environments on fatigue life have also been expressed in terms of a fatigue life correction factor F,,,, which is defined as the ratio of life in air at room temperature to that in water at the service temperature. The fatigue life correction factor for austenitic SSs, based on the ANL model, is given by F,, = exp(0.734 - T' ' O'), (38) where the constants T', E', and 0' are defined in Eqs. 35-37. Note that because the ANL model for austenitic SSs has been updated in the present report, the constant 0.734 in Eq. 38 is different from the values reported earlier in NUREG/CR-5704, 6815, and 6878. Relative to the earlier expressions, correction factors determined from Eq. 38 are 45-60% lower. A threshold strain amplitude (one-half of the applied strain range) is also defined, below which LWR coolant environments have no effect on fatigue life, i.e., Fen = 1. The threshold strain amplitude is 0.10% (195 MPa stress amplitude) for austenitic SSs. To incorporate environmental effects into a Section III fatigue evaluation, the fatigue usage for a specific stress cycle, based on the proposed new fatigue design curve (Fig. 37 and Table 9 in Section 5.1.8), is multiplied by the correction factor. Further details for incorporating environmental effects into fatigue evaluations are presented in Appendix A.

• The Fen approach may be used to incorporateenvironmental effects into the Codefatigue evaluations.

63

This page is intentionally left blank.

64

6 Ni-Cr-Fe Alloys and Welds The relevant fatigue E-N data for Ni-Cr-Fe alloys and their welds in air and water environments include the data compiled by Jaske and O'Donnell 72 for developing fatigue design criteria for pressure vessel alloys; the JNUFAD database from Japan; studies at MHI, IHI, and Hitachi in Japan; 33 studies at 75 Knolls Atomic Power Laboratory; 76,7 7 work sponsored by EPRI at Westinghouse Electric Corporation; 8

the tests performed by GE in a test loop at the Dresden I reactor; and the results of Van Der Sluys et al. 7 8 For Alloys 600 and 690, nearly 70% of the tests in air were conducted at room temperature and the remainder at 83-325°C. For Ni-Cr-Fe alloy welds (e.g., Alloys 82, 182, 132, and 152) nearly 85% of the tests in air were conducted at room temperature. In water, nearly 60% of the tests were conducted in simulated BWR environment (z0.2 ppm DO) and 40% in PWR environment (<0.01 ppm DO); tests in BWR water were performed at 288°C and in PWR water at 315 or 325'C. The existing fatigue data also include some tests in water with all volatile treatment (AVT) and at very high frequencies, e.g., 20 Hz to 40 kHz. 75 As expected, environmental effects on fatigue life were not observed for these tests; the results in AVT water are not included in the present analysis.

6.1 Air Environment 6.1.1 Experimental Data The fatigue e-N data for Alloys 600 and 690 in air at temperatures between room temperature and 316'C are shown in Fig. 56, and those for Ni-Cr-Fe alloy welds (e.g., Alloys 82, 182, 132, and 152) in air at temperatures between room temperature and 315'C are shown in Fig. 57. The best-fit curve for austenitic SSs based on the updated ANL model (Eq. 32 in Section 5.1.7) and the ASME Section III mean-data curve are included in the figures. The results indicate that although the data for Alloy 690 are very limited, the fatigue lives of Alloy 690 are comparable to those of Alloy 600 (Fig. 56). Similarly, the fatigue lives of Alloy 152 weld are comparable to those of Alloys 82, 182, and 132 welds (Fig. 57). Also, the fatigue lives of the Ni-Cr-Fe alloy welds are comparable to those of the wrought Alloys 600 and 690 in the low-cycle regime (i.e., <105 cycles) and are slightly superior to the lives of wrought materials in the high-cycle regime.

Alloy 600 Alloy 690, Air Air . , Ream remp 204'C c' 315'C 0 260-316°C 1.0 -_ J 1.0 ASME Code er ASME Code

<> Mean Curve Z3 Mean Curve

-- SSs AAustenitic Austenitic SSs CL Best-Fit Air C. . Best-Fit Air

< ANL Model <A Austenitic SSs . . Austenitic SSs 0.1 3 4 4 7 102 10 10 105 106 107 108 102 103 10 105 106 10 108 Fatigue Life (Cycles) Fatigue Life (Cycles)

Figure 56. Fatigue c-N behavior for Alloys 600 and 690 in air at temperatures between room temperature and 315'C (Refs. JNUFAD data, 72, 75-78).

65

I Alloys 82, 182, 132 Welds, Air Alloy 152 Weld, Air

• .. I*

> 21Tc (A* 82) 11 315Tc A 315-C (A 82)

?1.0 1.0 ASME Code ASME Code A. .- , ,*. *,Mean Curve "c- . MeanCure E Austenitic SSs .. Austenitic SSs S~ Best-Fit 0.1 Best-Fit Air ANL Model Austenitic SSs 0.1 Best-Fit Air ANLModel Austenitic SSs I

102 103 104 105 106 107 108 102 103 104 .105 106 107 108 Fatigue Life (Cycles) Fatigue Life (Cycles)

Figure 57. Fatigue c-N behavior for Alloys 82, 182, 132, and 152 welds in air at various temperatures (Refs. JNUFAD data, 72-78).

The fatigue lives of Alloy 600 are generally longer at high temperatures than at room temperature (Fig. 56a). 75 - 7 7 A similar behavior is observed for its weld metal, e.g., Alloy 82 (Fig. 57a). However, limited data for Alloy 690 (Fig. 56b) and its weld metal, Alloy 152 (Fig. 57b), indicate little or no effect of temperature on their fatigue lives. The existing data are inadequate to determine the effect of strain rate on the fatigue life of Ni-Cr-Fe alloys.

The results also indicate that the fatigue data for Ni-Cr-Fe alloys, including welds, are not consistent with the current ASME Code mean curve for austenitic SSs. The data for Alloys 600 and 690 show very good agreement with the updated ANL fatigue life model for austenitic SSs (Fig. 56a). Also, the fatigue data for Alloys 82, 182, and 132 are consistent with the updated ANL model in the low-cycle I

regime and somewhat conservative with respect to the model in the high-cycle regime (Fig. 57a).

• For Alloys 600 and 690 and their welds, the updated ANL fatigue life model proposed in the present reportfor austeniticSSs (Eq. 32) is either consistent or conservative with respect to the fatigue e-N data.

6.1.2 Fatigue Life Model I For Ni-Cr-Fe alloys, fatigue evaluations are based on the fatigue design curve for austenitic SSs.

However, the existing fatigue c-N data for Ni-Cr-Fe alloy and their welds are not consistent with the current ASME Code fatigue design curve for austenitic SSs. As discussed above, the data are either I

comparable or slightly conservative with respect to the updated ANL model for austenitic SSs, e.g., Eq. 32. Thus, the new fatigue design curve proposed in the present report for austenitic SSs and presented in Fig. 37 and Table 9 adequately represents the fatigue c-N behavior of Ni-Cr-Fe alloys and their welds.

• The new design curve for austeniticSSs may also be usedfor Ni-Cr-Fe alloys and their welds. I I

I I

6.2 LWR Environment 6.2.1 Experimental Data The fatigue lives of Ni-Cr-Fe alloys and their welds are also decreased in LWR environments; the fatigue s-N data for various Ni-Cr-Fe alloys in simulated BWR water at z289°C and PWR water at 315-325°C are shown in Figs. 58 and 59, respectively. The s-N curves based on the ANL model for austenitic SSs (Eq. 32 in Section 5.1.7) and the ASME Section III mean-data curve for austenitic SSs are also included in the figures. The results indicate that environmental effects on the fatigue life of Ni-Cr-Fe alloys are strongly dependent on key parameters such as strain rate, temperature, and DO level in water.

Similar to SSs, the effect of coolant environment on the fatigue life of Ni-Cr-Fe alloys is greater in the low-DO PWR environment than in the high-DO BWR environment. However, under similar loading and environmental conditions, the extent of the effects of environment is considerably less for the Ni-Cr-Fe alloys than for austenitic SSs. In general, environmental effects on fatigue life are the same for wrought and weld alloys.

rm- ......................................................

I'"1. . I. .. .'I ...... . . .. . . . . ... . . . ... I . . .. .I Alloy 600 Strain Rate (%/s) Ni-Alloy Welds 289°C BWR Water * , j* Sri1' Re% 0,0 289°C BWR Water Strain Rate (%Is)

V 0.04 0 0.04 20 A 0.004 C* 0.001 1.0 1.0 10 0

ASME Code ASMECode -

a Mean Curve -

2  ;.*

<.*Austenitic

• >"

Mean CurveSSs E "AustniticSSs j 10 U,

Best Fit Air Best-Fit Air r 0.1 ANL Model 0.1 ANL Model Austenitic SSs Austenitic SSs

,,,. ... .. . . . ,,I . ... .. I . . . ..... I . . . ..... I . . . . ,.. . .. I , . . . . I . . . .. I . . ...... I . .. . I . . ...... I -

3 7 4 5 102 10 104 105 106 10 108 102 103 10 10 106 107 108 Fatigue Life (Cycles) Fatigue Life (Cycles)

Figure 58. Fatigue s-N behavior for Alloy 600 and its weld alloys in simulated BWR water at =289°C (Refs. JNUFAD data, 33).

Alloy 600 & Alloys 132 & 82 Welds Strain Rate (%Is) 215-325°C PWR Water Alloy 600/132 V 0.4 Alloy 82 :F 1.0 "0 0.1 1.0 ASME Code > 0.01 Mean Curve V \ Austenitic SSs E

0.1 Best-Fit Air ANL Model 0.1 Austenitic SSs 4 7 102 103 10 105 106 10 108 102 103 104 105 106 107 108 Fatigue Life (Cycles) Fatigue Life (Cycles)

Figure 59. Fatigue s-N behavior for Alloys 600 and 690 and their weld alloys in simulated PWR water at 315 or 325°C (Refs. 33, 78).

67

I 105 1 .I.T... 10l 5 Strain Amplitude

. Ni-Alloys Alloy 600 .

Alloy 690/152 325°C PWR Water 289TC BWR Water I 04%

o).6% lI I Strain C Amplitude 0.3%

Q"*'104 -- Alloy 600/132 '5%

>In.) 104 Z n- 075% .. i....

0 04/s .,

LL103.... I . k -- r ..... ......*'$' - -

.I ........

4 4 10-5 10- 10-3 10-2 10-1 100. 10-5 10- 10-3 10-2 10-1 100 Strain Rate (%Is) Strain Rate (%Is) i Figure 60. Dependence of fatigue lives of Alloys 690 and 600 and their weld alloys in PWR water at 325°C and Alloy 600 in BWR water at 289°C (Refs. JNUFAD data, 33, 78). 3 6.2.2 Effects of Key Parameters The existing fatigue &-N data for Ni-Cr-Fe alloys in LWR environments are very limited; the effects of the key loading and environmental parameters (e.g., strain rate,temperature, and DO level) on I

fatigue life of these alloys have been evaluated by Higuchi et al. 3 3 The fatigue lives of Alloys 600 and 690 and their weld metals (e.g., Alloys 132 and 152) in simulated PWR and BWR water at different strain amplitudes are plotted as a function of strain rate in Fig. 60. The fatigue life of these alloys decreases I

logarithmically with decreasing strain rate. Although fatigue data at strain rates below 0.001%/s are not available, for Ni-Cr-Fe alloys, the effect of strain rate is assumed to be similar to that for austenitic SSs; the effect saturates at 0.0004%/s strain rate. Also, the threshold strain rate below which environmental effects are significant cannot be determined from the present data. Higuchi et al. 33 have defined a threshold strain rate of 1.8%/s in high-DO BWR water and 26.1%/s in low-DO PWR water. As discussed in Section 6.2.3, an average threshold value of 5%/s provides good estimates of fatigue lives of Ni-Cr-Fe alloys in LWR environments.

I The results also indicate that the effects of environment are greater in the low-DO PWR water than in high-DO BWR water. For example, a three orders of magnitude decrease in strain rate decreases the fatigue life of these alloys by a factor ofz3 in PWR water and by ;2 in BWR water.

The existing data are inadequate to determine accurately the functional form for the effect of temperature on fatigue life or to define the, threshold strain amplitude below which environmental effects I

on fatigue life do not occur. Such effects are assumed to be similar to those observed in austenitic SSs. It is also assumed that a slow strain rate applied during the tensile-loading cycle (i.e., up-ramp with increasing strain) is primarily responsible for the environmentally assisted reduction in fatigue life. Slow rates applied during both tensile- and compressive-loading cycles (i.e., up- and down-ramps) do not further decrease fatigue life compared with that observed for tests with only a slow tensile-loading cycle.

Thus, loading and environmental conditions during the tensile-loading cycle are important for I

environmentally assisted reduction of the fatigue lives of Ni-Cr-Fe alloys.

6.2.3 Environmental Correction Factor The effects of reactor coolant environments on fatigue life of Ni-Cr-Fe alloys can also be expressed in terms of a fatigue life correction factor Fen, which is defined as the ratio of life in air at room I 68

temperature to that in water at the service temperature. The existing fatigue data are very limited to develop a fatigue life model for estimating the fatigue life of Ni-Cr-Fe alloys in LWR environments.

However, as discussed above in Section 6.2.2, environmental effects for these alloys show the same trends as those observed for austenitic SSs. Thus, Fen for Ni-Cr-Fe alloys can be expressed as Fen = exp(T' t ' 0'), (39) where T', C', and 0' are transformed temperature, strain rate, and DO, respectively. The functional forms for these transformed parameters were obtained from the best fit of the experimental data and are defined as follows:

T' = T/325 (T < 325-C)

T'=1 (T > 325°C) (40) t=0 (t > 5.0%/s)

S Iln(E/5.0) (0.0004 < t < 5.0%/s)

C= ln(0.0004/5.0) (t < 0.0004%/s) (41)

O' =0.09 (NWC BWR water)

O' =0.16 (PWR or HWC BWR water). (42)

The fatigue life of Ni-Cr-Fe alloys in LWR environments can be estimated from Eqs. 32 and 39-42. The experimental and estimated fatigue lives of various Ni-Cr-Fe alloys in BWR and PWR water are plotted in Fig. 61; the estimated values are either comparable or longer than those observed experimentally.

106 106 105~

105 103 103 4 3 103 10 105 106 10 104 105 101 Observed Life (Cycles) Observed Life (Cycles)

Figure 61. The experimental and estimated fatigue lives of various Ni alloys in BWR and PWR environments (Refs. JNUFAD data, 33, 78).

A threshold strain amplitude (one-half of the applied strain range) is also defined, below which LWR coolant environments have no effect on fatigue life, i.e., Fen = 1. The value is assumed to be the same as that for austenitic SSs. The threshold strain amplitude is 0.10% (195 MPa stress amplitude) for Ni-Cr-Fe alloys. To incorporate environmental effects into a Section I11 fatigue evaluation, the fatigue 69

usage for a specific stress cycle, based on the proposed new fatigue design curve for austenitic SSs (Fig. 37 and Table 9 in Section 5.1.8), is multiplied by the correction factor. Further details for incorporating environmental effects into fatigue evaluations are presented in Appendix A.

• The Fen approach may be used to incorporateenvironmental effects into the Codefatigue evaluations.

70

7 Margins in ASME Code Fatigue Design Curves Conservatism in the ASME Code fatigue evaluations may arise from (a) the fatigue evaluation procedures and/or (b) the fatigue design curves. The overall conservatism in ASME Code fatigue evaluations has been demonstrated in fatigue tests on components. 120,121 Mayfield et al. 120 have shown that, in air, the margins on the number of cycles to failure for elbows and tees were 40-310 and 104-510, respectively, for austenitic SS and 118-2500 and 123-1700, respectively, for carbon steel. The margins for girth butt welds were significantly lower, 6-77 for SS and 14-128 for carbon steel. Data obtained by Heald and Kiss 12 1 on 26 piping components at room temperature and 288'C showed that the design margin for cracking exceeds 20, and for most of the components, it is >100. In these tests, fatigue life was expressed as the number of cycles for the crack to penetrate through the wall, which ranged in thickness from 6 to 18 mm. Consequently, depending on wall thickness, the actual margins to form a 3-mm crack may be lower by a factor of more than 2.

Deardorff and Smith 122 discussed the types and extent of conservatism present in the ASME Section III fatigue evaluation procedures and the effects of LWR environments on fatigue margins. The sources of conservatism in the procedures include the use of design transients that are significantly more severe than those experienced in service, conservative grouping of transients, and use of simplified elastic-plastic analyses that lead to higher stresses. The authors estimated that the ratio of the CUFs computed with the mean experimental curve for test specimen data in air and more accurate values of the stress to the CUFs computed with the Code fatigue design curve were -60 and 90, respectively, for PWR and BWR nozzles. The reductions in these margins due to environmental effects were estimated to be factors of 5.2 and 4.6 for PWR and BWR nozzles, respectively. Thus,

Deardorff and Smith 122 argue that,

after accounting for environmental effects, factors of 12 and 20 on life for PWR and BWR nozzles, respectively, account for uncertainties due to material variability, surface finish, size, mean stress, and loading sequence.

However, other studies on piping and components indicate that the Code fatigue design procedures do not always ensure large margins of safety.t 23 ,12 4 Southwest Research Institute performed fatigue tests in room-temperature water on 0.9 1-m-diameter carbon and low-alloy steel vessels. 123 In the low-cycle regime, z5-mm-deep cracks were initiated slightly above (a factor of <2) the number of cycles predicted by the ASME Code design curve (Fig. 62a). Battelle-Columbus conducted tests on 203-mm or 914-mm carbon steel pipe welds at room temperature in an inert environment, and Oak Ridge National Laboratory (ORNL) performed four-point bend tests on 406-mm-diameter Type 304 SS pipe removed from the C-reactor at the Savannah River site. 124 The results showed that the number of cycles to produce a leak was lower, and in some cases significantly lower, than that expected from the ASME Code fatigue design curves (Fig. 62a and b). The most striking results are for the ORNL "tie-in" and flawed "test" weld; these specimens cracked completely through the 12.7-mm-thick wall in a life 6 or 7 times shorter than expected from the Code curve. Note that the Battelle and ORNL results represent a through-wall crack; the number of cycles to initiate a 3-mm crack may be a factor of 2 lower.

Much of the margin in the current evaluations arises from design procedures (e.g., stress analysis rules and cycle counting) that, as discussed by

Deardorff and Smith,

122 are quite conservative. However, the ASME Code permits new and improved approaches to fatigue evaluations (e.g., finite-element analyses, fatigue monitoring, and improved Ke factors) that can significantly decrease the conservatism in the current fatigue evaluation procedures.

71

I I

(5.

'i 103 Type 304 Stainless Steel Room Temp.

Room emp ORNL

• ,, Tie-in we'd Tes weid lack of penetrati ail I

6

'0 ASME Code I

SDesignCurve )

'a 102 I-I I

,,

3 5 103 106 10 104 10 106 104 105 Number of Cycles, N Number of Cycles, N (a) (b)

Figure 62. Fatigue data for (a) carbon and low-alloy steel and (b) Type 304 stainless steel components (Refs. 123,124). I The factors of 2 on stress and 20 on cycles used in the Code were intended to cover the effects of variables that can influence fatigue life but were not investigated in the tests that provided the data for the curves. It is not clear whether the particular values of 2 and 20 include possible conservatism. A study I

sponsored by the PVRC to assess the margins of 2 and 20 in fatigue design curves concluded that these margins should not be changed.

12 5 I

The variables that can affect fatigue life in air and LWR environments can be broadly classified into three groups:

(a) Material I

(i)

(ii)

Composition Metallurgy: grain size, inclusions, orientation within a forging or plate (iii) Processing: cold work, heat treatment I

(iv) Size and geometry (v) Surface finish: fabrication surface condition (vi) Surface preparation: surface work hardening I (b) Loading (i)

(ii)

Strain rate: rise time Sequence: linear damage summation or Miner's rule I (iii) Mean stress (c)

(iv) Biaxial effects: constraints Environment I (i) Water chemistry: DO, lithium hydroxide, boric acid concentrations (ii) Temperature (iii) Flow rate I The existing fatigue c-N database covers an adequate range of material parameters (i-iii), a loading parameter (i), and the environment parameters (i-ii); therefore, the variability and uncertainty in fatigue life due to these parameters have been incorporated into the model. The existing data are most likely I

conservative with respect to the effects of surface preparation because the fatigue c-N data are obtained for specimens that are free of surface cold work. Fabrication procedures for fatigue test specimens I 72 I

I

generally follow American Society for Testing and Materials (ASTM) guidelines, which require that the final polishing of the specimens avoid surface work-hardening. Biaxial effects are covered by design procedures and need not be considered in the fatigue design curves.

As discussed earlier, under the conditions typical of operating BWRs, environmental effects on the fatigue life are a factor ofz2 lower at high flow rates (7 m/s) than those at very low flow rates (0.3 m/s or lower) for carbon and low-alloy steels and are independent of flow rate for austenitic SSs.19,20 However, because of the uncertainties in the flow conditions at or near the locations of crack initiation, the beneficial effect, of flow rate on the fatigue life of carbon and low-alloy steels is presently not included in fatigue evaluations.

Thus, the contributions of four groups of variables, namely, material variability and data scatter, specimen size and geometry, surface finish, and loading sequence (Miner's rule), must be considered in developing fatigue design curves that are applicable to components.

7.1 Material Variability and Data Scatter The effects of material variability and data scatter must be included to ensure that the design curves not only describe the available test data well, but also adequately describe the fatigue lives of the much larger number of heats of material that are found in the field. The effects of material variability and data scatter have been evaluated for the various materials by considering the best-fit curves determined from tests on individual heats of materials or loading conditions as samples of the much larger population of heats of materials and service conditions of interest. The fatigue behavior of each of the heats or loading conditions is characterized by the value of the constant A in Eq. 6. The values of A for the various data sets are ordered, and median ranks are used to estimate the cumulative distribution of A for the population. The distributions were fit to lognormal curves. The median value of A and standard deviation for each sample, as well as the number of data sets in the sample, are listed in Table 11. The 95/95 value of the margin on the median value to account for material variability and data scatter vary from 2.1 to 2.8 for the various samples. These margins applied to the mean value of life determined from the ANL fatigue life models provide 95% confidence that the fatigue life of 95 percentile of the materials and loading conditions of interest will be greater than the resultant value.

Table 11. The median value of A and standard deviation for the various fatigue c-N data sets used to evaluate material variability and data scatter.

Air Environment Water Environment Median Value Standard Number of Median Value Standard Number'of ofA Deviation Data Sets ofA Deviation Data Sets Carbon Steel 6.583 0.477 17 5.951 0.376 33 Low-Alloy Steel 6.449 0.375 32 5.747 0.484 26 Stainless Steel 6.891 0.417 51 6.328 0.462 36 7.2 Size and Geometry The effect of specimen size on the fatigue life was reviewed in earlier reports. 6 ,39 Various studies conclude that "size effect" is not a significant parameter in the design curve margins when the fatigue curve is based on data from axial strain control rather than bending tests. No intrinsic size effect has been observed for smooth specimens tested in axial loading or plain bending. However, a size effect does occur in specimens tested in rotating bending; the fatigue endurance limit decreases by z25% if the specimen size is increased from 2 to 16 mm but does not decrease further with larger sizes. Also, some effect of size and geometry has been observed on small-scale-vessel tests conducted at the Ecole 73

I Polytechnique in conjunction with the large-size-pressure-vessel tests carried out by the Southwest Research Institute. 123 The tests at the Ecole Polytechnique were conducted in room-temperature water on 19-mm-thick shells with =305-mm inner diameter nozzles and made of machined bar stock. The results indicate that the fatigue lives determined from tests on the small-scale-vessel are 30-50% lower than those obtained from tests on small, smooth fatigue specimen. However, the difference in fatigue lives in these tests cannot be attributed to specimen size alone, it is due to the effects of both size and surface finish.

I During cyclic loading, cracks generally form at surface irregularities either already in existence or produced by slip bands, grain boundaries, second phase particles, etc. In smooth specimens, formation of surface cracks is affected by the specimen size; crack initiation is easier in larger specimens because of the increased surface area and, therefore, increased number of sites for crack initiation. Specimen size is not likely to influence crack initiation in specimens with rough surfaces because cracks initiate at existing irregularities on the rough surface. As discussed in the next section, surface roughness has a large effect on fatigue life. Consequently, for rough surfaces, the effect of specimen size may not be considered in the margin of 20 on life. However, conservatively, a factor of 1.2-1.4 on life may be used to incorporate size effects on fatigue life in the low-cycle regime.

7.3 Surface Finish 3 The effect of surface finish must be considered to account for the difference in fatigue life expected in actual components with industrial-grade surface finish compared to the smooth polished surface of a test specimen. Fatigue life is sensitive to surface finish; cracks can initiate at surface irregularities that are 3

normal to the stress axis. The height, spacing, shape, and distribution of surface irregularities are important for crack initiation. The effect of surface finish on crack initiation is expressed by Eq. 12 in terms of the RMS value of surface roughness (Rq).

The roughness of machined surfaces or natural finishes can range from Z-0.8 to 6.0 gm. Typical surface finish for various machining processes is in the range of 0.2-1.6 gim for cylindrical grinding, 0.4-3.0 jim for surface grinding, 0.8-3.0 jim for finish turning, and drilling and 1.6-4.0 jtm for milling.

I For fabrication processes, it is in the range of 0.8-3.0 [im for extrusion and 1.6-4.0 gm for cold rolling.

Thus, from Eq. 12, the fatigue life of components with such rough surfaces may be a factor of 2-3.5 lower than that of a smooth specimen.

Limited data in LWR environments on specimens that were intentionally roughened indicate that the effects of surface roughness on fatigue life is the same in air and water environments for austenitic SSs, but are insignificant in water for carbon and low-alloy steels. Thus, in LWR environments, a factor of 2.0-3.5 on life may also be used to account for the effects of surface finish on the fatigue life of austenitic SSs, but the factor may be lower for carbon and low-alloy steels, e.g., a factor of 2 may be used for carbon and low-alloy steels.

I 7.4 Loading Sequence The effects of variable amplitude loading of smooth specimens were also reviewed in an earlier report. 39 In a variable loading sequence, the presence of a few cycles at high strain amplitude causes the fatigue life at smaller strain amplitude to be significantly lower than that at constant-amplitude loading, i.e., the fatigue limit of the material is lower under variable loading histories.

I 74.

As discussed in Section 2, fatigue life has conventionally been divided into two stages: initiation, expressed as the cycles required to form microstructurally small cracks (MSCs) on the surface, and propagation, expressed as cycles required to propagate these MSCs to engineering size. The transition from initiation to propagation stage strongly depends on applied stress amplitude; at stress levels above the fatigue limit, the transition from initiation to propagation stage occurs at crack depths in the range of 150 to 250 tim. However, under constant loading at stress levels below the fatigue limit of the material (e.g., AG, in Fig. 1), although microcracks z10 ýltm can form quite early in life, they do not grow to an engineering size. Under the variable loading conditions encountered during service of power plants, cracks created by growth of MSCs at high stresses (AG 3 in Fig. 1) to depths larger than the transition crack depth can then grow to an engineering size even at stress levels below the fatigue limit.

Studies on fatigue damage in Type 304 SS under complex loading histories 1 2 6 indicate that the loading sequence of decreasing strain levels (i.e., high strain level followed by low strain level) is more damaging than that of increasing strain levels. The fatigue life of the steel at low strain levels decreased by a factor of 2-4 under a decreasing-strain sequence. In another study, the fatigue limit of medium carbon steels was lowered even after low-stress high-cycle fatigue; the higher the stress, the greater the decrease in fatigue threshold.1 27 A recent study on Type 316NG and Ti-stabilized Type 316 SS on strain-controlled tests in air and PWR environment with constant or variable strain amplitude reported a factor 128 of 3 or more decrease in fatigue life under variable amplitude compared with constant amplitude.

Although the strain spectrum used in the study was not intended to be representative of real transients, it represents a generic case and demonstrates the effect of loading sequence on fatigue life.

Because variable loading histories primarily influence fatigue life at low strain levels, the mean fatigue F-N curves are lowered to account for damaging cycles that occur below the constant-amplitude fatigue limit of the material. However, conservatively, a factor of 1.2-2.0 on life may be used to incorporate the possible effects of load histories on fatigue life in the low-cycle regime.

7.5 Fatigue Design Curve Margins Summarized The ASME Code fatigue design curves are currently obtained from the mean data curves by first adjusting for the effects of mean stress, and then reducing the life at each point of the adjusted curve by a factor of 2 on strain and 20 on life, whichever is more conservative. The factors on strain are needed primarily to account for the variation in the fatigue limit of the material caused by material variability, component size, surface finish, and load history. Because these variables affect life through their influence on the growth of short cracks (<I 00 [tim), the adjustment on strain to account for such variations is typically not cumulative, i.e., the portion of the life can only be reduced by a finite amount. Thus, it is controlled by the variable that has the largest effect on life. In relating the fatigue lives of laboratory test specimens to those of actual reactor components, the factor of 2 on strain that is currently being used to develop the Code design curves is adequate to account for the uncertainties associated with material variability, component size, surface finish, and load history.

The factors on life are needed to account for variations in fatigue life in the low-cycle regime.

Based on the discussions presented above the effects of various material, loading, and environmental parameters on fatigue life may be summarized as follows:

(a) The results presented in Table 11 may be used to determine the margins that need to be applied to the mean value of life to ensure that the resultant value of life would bound a specific percentile (e.g., 95 percentile) of the materials and loading conditions of interest.

75

(b) For rough surfaces, specimen size is not likely to influence fatigue life, and therefore, the effect of specimen size need not be considered in the margin of 20 on life. However, conservatively, a factor of 1.2-1.4 on life may be used to incorporate size effects on fatigue life.

(c) Limited data indicate that, for carbon and low-alloy steels, the effects of surface roughness on fatigue life are insignificant in LWR environments. A factor of 2 on life may be used for carbon and low-alloy steels in water environments instead of the 2.0-3.5 used for carbon and low-alloy steels in air and for austenitic SSs in both air and water environments.

(d) Variable loading histories primarily influence fatigue life at low strain levels, i.e., in the high-cycle regime, and the mean fatigue a-N curves are lowered by a ýfactor of 2 on strain to account for damaging cycles that occur below the constant-strain fatigue limit of the material. Conservatively, a factor of 1.2-2.0 on life may be used to incorporate the possible effects of load histories on fatigue life in the low-cycle regime.

The subfactors that are needed to account for the effects of the various material, loading, and environmental parameters on fatigue life are summarized in Table 12. The total adjustment on life may vary from 6 to 27. Because the maximum value represents a relatively poor heat of material and assumes the maximum effects of size, surface finish, and loading history, the maximum value of 27 is likely to be quite conservative. A value of 20 is currently being used to develop the Code design curves from the mean-data curves.

Table 12. Factors on life applied to mean fatigue e-N curve to account for the effects of various material, loading, and environmental parameters.

Parameter Material Variability and Data Scatter Section III Criterion Document Present Report I (minimum to mean)

Size Effect Surface Finish, etc.

2.0 2.5 4.0 2.1-2.8 1.2-1.4 2.0-3.5*

I I

Loading History - 1.2-2.0 Total Adjustment 20 6.0-27.4

• Afactor of 2 on life may be used for carbon and low-alloy steels in LWR environments.

To determine the most appropriate value for the design margin on life, Monte Carlo simulations were performed using the material variability and data scatter results given in Table 11, and the margins needed to account for the effects of size, surface finish, and loading history listed in Table 12.

I A lognonnal distribution was also assumed for the effects of size, surface finish, and loading history, and the minimum and maximum values of the adjustment factors, e.g., 1.2-1.4 for size, 2.0-3.5 for surface finish, and 1.2-2.0 for loading history, were assumed to represent the 5th and 95th percentile, I

respectively. The cumulative distribution of the values of A in the fatigue E-N curve for test specimens and the adjusted curve that represents the behavior of actual components is shown in Fig. 63 for carbon and low-alloy steels and austenitic SSs.

I The results indicate that, relative to the specimen curve, the median value of constant A for the component curve decreased by a factor of 5.6 to account for the effects of size, surface finish, and loading I history, and the standard deviation of heat-to-heat variation of the component curve increased by 6-10%.

The margin that has to be applied to the mean data curve for test specimens to obtain a component curve that would bound 95% of the population, is 11.0-12.7 for the various materials; the values are given in I 76 I

I

Table 13. An average value of 12 on life may be used for developing fatigue design curves from the mean data curve. The choice of bounding the 95th percentile of the population for a design curve is somewhat arbitrary. It is done with the understanding that the design curve controls fatigue initiation, not failure. The choice also recognizes that there are conservatisms implied in the choice of log normal distributions, which have an infinite tail, and in the identification of what in many cases are bounding values of the effects as 95th percentile values.

1.0 0.8 L- U-C:

0 0.6 23 0

. 0.4 E

0.2 0.0 5

Constant A Constant A 1.0 0.8 L.

0 0.6 Figure 63.

Estimated cumulative distribution of parameter A in the ANL models that represent the fatigue life

-- 0.4 of test specimens and actual components in air.

0 0.2 0.0 4 5 6 7 8 Constant A Table 13. Margin applied to the mean values of fatigue life to bound 95% of the population.

Material Air Environment Carbon Steels 12.6 Low-Alloy Steels 11.0 Austenitic Stainless Steels 11.6 77

I I

These results suggest that for all materials, the current ASME Code requirements of a factor of 20 on cycles to account for the effects of material variability and data scatter, as well as specimen size, surface finish, and loading history, contain at least a factor of 1.7 conservatism (i.e., 20/12 ' 1.7). Thus, to reduce this conservatism, fatigue design curves may be obtained from the mean data curve by first correcting for mean stress effects using the modified Goodman relationship, and then reducing the mean-stress adjusted curve by a factor of 2 on stress or 12 on cycles, whichever is more conservative. Fatigue design curves have been developed from the ANL fatigue life models using this procedure; the curves for carbon and low-alloy steels are presented in Section 4.1.10 and for wrought and cast austenitic SSs in Section 5.1.8.I I

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8 Summary The existing fatigue F-N data for carbon and low-alloy steels, wrought and cast austenitic SSs, and Ni-Cr-Fe alloys have been evaluated to define the effects of key material, loading, and environmental parameters on the fatigue lives of these steels. The fatigue lives of these materials are decreased in LWR environments; the magnitude of the reduction depends on temperature, strain rate, DO level in water, and, for carbon and low-alloy steels, the S content of the steel. For all steels, environmental effects on fatigue life are significant only when critical parameters (temperature, strain rate, DO level, and strain amplitude) meet certain threshold values. Environmental effects are moderate, e.g., less than a factor of 2 decrease in life, when any one of the threshold conditions is not satisfied. The threshold values of the critical parameters and the effects of other parameters (such as water conductivity, water flow rate, and material heat treatment) on the fatigue life of the steels are summarized.

In air, the fatigue life of carbon and low-alloy steels depends on steel type, temperature, orientation, and strain rate. The fatigue life of carbon steels is a factor of -1.5 lower than that of low-alloy steels. For both steels, fatigue life decreases with increase in temperature. Some heats of carbon and low-alloy steels exhibit effects of strain rate and orientation. For these heats, fatigue life decreases with decreasing strain rate. Also, based on the distribution and morphology of sulfides, the fatigue properties in the transverse orientation may be inferior to those in the rolling orientation. The data indicate significant heat-to-heat variation; at 288°C, the fatigue life of carbon and low-alloy steels may vary by up to a factor of 3 above or below the mean value. Fatigue life is very sensitive to surface finish; the fatigue life of specimens with rough surfaces may be up to a factor of 3 lower than that of smooth specimens. The results also indicate that in room-temperature air, the ASME mean curve for low-alloy steels is in good agreement with the available experimental data, and the curve for carbon steels is somewhat conservative.

The fatigue lives of both carbon and low-alloy steels are decreased in LWR environments; the reduction depends on temperature, strain rate, DO level in water, and S content of the steel. The fatigue life is decreased significantly when four conditions are satisfied simultaneously, viz., the strain amplitude, temperature, and DO in water are above certain minimum levels, and the strain rate is below a threshold value. The S content in the steel is also important; its effect on life depends on the DO level in water.

Although the microstructures and cyclic-hardening behavior of carbon and low-alloy steels differ significantly, environmental degradation of the fatigue life of these steels is very similar. For both steels, only a moderate decrease in life (by a factor of <2) is observed when any one of the threshold conditions is not satisfied, e.g., low-DO PWR environment, temperatures <1501C, or vibratory fatigue. The existing fatigue S-N data have been reviewed to establish the critical parameters that influence fatigue life and define their threshold and limiting values within which environmental effects are significant.

In air, the fatigue lives of Types 304 and 316 SS are comparable; those of Type 316NG are superior to those of Types 304 and 316 SS at high strain amplitudes. The fatigue lives of austenitic SSs in air are independent of temperature in the range from room temperature to 427°C. Also, variation in strain rate in the range of 0.4-0.008%/s has no effect on the fatigue lives of SSs at temperatures up to 400'C. The fatigue E-N behavior of cast SSs is similar to that of wrought austenitic SSs. The results indicate that the ASME mean-data curve for SSs is not consistent with the experimental data at strain amplitudes <0.5% or stress amplitudes <975 MPa (<141 ksi); the ASME mean curve predicts significantly longer lives than those observed experimentally.

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The fatigue lives of cast and wrought austenitic SSs decrease in LWR environments compared to those in air. The decrease depends on strain rate, DO level in water, and temperature. A minimum threshold strain is required for an environmentally assisted decrease in the fatigue life of SSs, and this strain appears to be independent of material type (weld or base metal) and temperature in the range of 250-325°C. Environmental effects on fatigue life occur primarily during the tensile-loading cycle and at strain levels greater than the threshold value. Strain rate and temperature have a strong effect on fatigue life in LWR environments. Fatigue life decreases with decreasing strain* rate below 0.4%/s; the effect saturates at 0.0004%/s. Similarly, the fatigue s-N data suggest a threshold temperature of 150'C; in the range of 150-325°C, the logarithm of life decreases linearly with temperature.

The effect of DO level may be different for different steels. In low-DO water (i.e., <0.01 ppm DO) the fatigue lives of all wrought and cast austenitic SSs are decreased significantly; composition or heat treatment of the steel has little or no effect on fatigue life. However, in high-DO water, the environmental effects on fatigue life appear to be influenced by the composition and heat treatment of the steel; the effect of high-DO water on the fatigue lives of different compositions and heat treatment of SSs is not well established. Limited data indicate that for a high-C Type 304 SS, environmental effects are significant only for sensitized steel. For a low-C Type 3 16NG SS, some effect of environment was observed even for mill-annealed steel (nonsensitized steel) in high-DO water, although the effect was smaller than that observed in low-DO water. Limited fatigue s-N data indicate that the fatigue lives of cast SSs are approximately the same in low- and high-DO water and are comparable to those observed for wrought SSs in low-DO water. In the present report, environmental effects on the fatigue lives of wrought and cast austenitic SSs are considered to be the same in high-DO and low-DO environments.

The fatigue s-N data for Ni-Cr-Fe alloys indicate that although the data for Alloy 690 are very limited, the fatigue lives of Alloy 690 are comparable to those of Alloy 600. Also, the fatigue lives of the Ni-Cr-Fe alloy welds are comparable to those of the wrought Alloys 600 and 690 in the low-cycle regime, i.e., <105 cycles, and are slightly superior to the lives of wrought materials in the high-cycle regime. The fatigue data for Ni-Cr-Fe alloys in LWR environments are very limited; the effects of key loading and environmental parameters on fatigue life are similar to those for austenitic SSs. For example, the fatigue life of these steels decreases logarithmically with decreasing strain rate. Also, the effects of environment are greater in the low-DO PWR water than the high-DO BWR water. The existing data are inadequate to determine accurately the functional form for the effect of temperature on fatigue life.

Fatigue life models developed earlier to predict fatigue lives of small smooth specimens of carbon and low-alloy steels and wrought and cast austenitic SSs as a function of material, loading, and environmental parameters have been updated/revised using a larger fatigue s-N database. The functional form and bounding values of these parameters were based on experimental observations and data trends.

The models are applicable for predicted fatigue lives <106 cycles. The ANL fatigue life model proposed in the present report for austenitic SSs in air is also recommended for predicting the fatigue lives of small smooth specimens of Ni-Cr-Fe alloys.

An approach, based on the environmental fatigue correction factor, is discussed to incorporate the effects of LWR coolant environments into the ASME Code fatigue evaluations. To incorporate environmental effects into a Section III fatigue evaluation, the fatigue usage for a specific stress cycle of load set pair based on the current Code fatigue design curves is multiplied by the correction factor.

The report also presents a critical review of the ASME Code fatigue design margins of 2 on stress and 20 on life and assesses the possible conservatism in the current choice of design margins. These factors cover the effects of variables that can influence fatigue life but were not investigated in the tests 80

that provided the data for the design curves. Although these factors were intended to be somewhat conservative, they should not be considered safety margins because they were intended to account for variables that are known to affect fatigue life. Data available in the literature have been reviewed to evaluate the margins on cycles and stress that are needed to account for the differences and uncertainties.

Monte Carlo simulations were performed to determine the margin on cycles needed to obtain a fatigue design curve that would provide a somewhat conservative estimate of the number of cycles to initiate a fatigue crack in reactor components. The results suggest that for both carbon and low-alloy steels and austenitic SSs, the current ASME Code requirements of a factor of 20 on cycles to account for the effects of material variability and data scatter, as well as size, surface finish, and loading history, contain at least a factor of 1.7 conservatism. Thus, to reduce this conservatism, fatigue design curves have been developed from the ANL model by first correcting for mean stress effects, and then reducing the mean-stress adjusted curve by a factor' of 2 on stress and 12 on cycles, whichever is more conservative. A detailed procedure for incorporating environmental effects into fatigue evaluations is also presented in Appendix A.

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107. Solomon, H. D., R. E. DeLair, and E. Tolksdorf, "LCF Crack Initiation in WB36 in High-Temperature Water," Proc. 9th Intl. Symp. on Environmental Degradation of Materials in Nuclear Power Systems - Water Reactors, F. P. Ford, S. M. Bruemmer, and G. S. Was, eds., The Minerals, Metals, and Materials Society, Warrendale, PA, pp. 865-872, 1999.

108. Cullen, W. H., M. Kemppainen, H. H~nninen, and K. T6rr6nen, "The Effects of Sulfur Chemistry and Flow Rate on Fatigue Crack Growth Rates in LWR Environments," NUREG/CR-4121, 1985.

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i 109. Van Der Sluys, W. A., and R. H. Emanuelson, "Environmental Acceleration of Fatigue Crack Growth in Reactor Pressure Vessel Materials and Environments," Environmentally Assisted Cracking: Science and Engineering, ASTM STP 1049, W. B. Lisagor, T. W. Crooker, and B. N.

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755-765, 1996.

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pp. 213-228, 1999.

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115. Kim, Y. J., "Characterization of the Oxide Film Formed on Type 316 Stainless Steel in 288°C Water in Cyclic Normal and Hydrogen Water Chemistries," Corrosion 51 (11), 849-860, 1995.

116. Kim, Y. J., "Analysis of Oxide Film Formed on Type 304 Stainless Steel in 288°C Water Containing Oxygen, Hydrogen, and Hydrogen Peroxide," Corrosion 55 (1), 81-88, 1999.

I 117. Chopra, 0. K., "Estimation of Fracture Toughness of Cast Stainless Steels During Thermal Aging in LWR Systems," NUREG/CR-4513, ANL-93/22, Aug. 1994.

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ASM International, Materials Park, OH, pp. 479-485, 1995.

I 119. Higuchi, M., "Review and Consideration of Unsettled Problems on Evaluation of Fatigue Damage in LWR Water," Proc. of the 2005 ASME Pressure Vessels and Piping Conf., July 17-21, 2005, Denver, CO, paper # PVP2005-71306.

120. Mayfield, M. E., E. C. Rodabaugh, and R. J. Eiber, "A Comparison of Fatigue Test Data on Piping with the ASME Code Fatigue Evaluation Procedure," ASME Paper 79-PVP-92, American Society of Mechanical Engineers, New York, 1979.

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A. F., and J. K. Smith, "Evaluation of Conservatisms and Environmental Effects in ASME Code, Section I11, Class 1 Fatigue Analysis," SAND94-0187, prepared by Structural Integrity Associates, San Jose, CA, under contract to Sandia National Laboratories, Albuquerque, NM, 1994.

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124. Scott, P. M., and G. M. Wilkowski, "A Comparison of Recent Full-Scale Component Fatigue Data with the ASME Section III Fatigue Design Curves," in Fatigue and Crack Growth: Environmental Effects, Modeling Studies, and Design Considerations, PVP Vol. 306, S. Yukawa, ed., American Society of Mechanical Engineers, New York, pp. 129-138, 1995.

125. Hechmer, J., "Evaluation Methods for Fatigue - A PVRC Project," in Fatigue, Environmental Factors, and New Materials, PVP Vol. 374, H. S. Mehta, R. W. Swindeman, J. A. Todd, S. Yukawa, M. Zako, W. H. Bamford, M. Higuchi, E. Jones, H. Nickel, and S. Rahman, eds.,

American Society of Mechanical Engineers, New York, pp. 191-199, 1998.

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1. PVP2006-ICPVT-93833.

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APPENDIX A Incorporating Environmental Effects into Fatigue Evaluations Al Scope This Appendix provides the environmental fatigue correction factor (Fen) methodology that is considered acceptable for incorporating the effects of reactor coolant environments on fatigue usage factor evaluations of metal components for new reactor construction. The methodology for performing fatigue evaluations for the four major categories of structural materials, e.g., carbon steel, low-alloy steels, wrought and cast austenitic stainless steels, and Ni-Cr-Fe alloys, is described.

A2 Environmental Correction Factor (Fen)

The effects of reactor coolant environments on the fatigue life of structural materials are expressed in terms of a nominal environmental fatigue correction factor, Fen,nor, which is defined as the ratio of fatigue life in air at room temperature (Nair,RT) to that in water at the service temperature (Nwater):

Fen,nomn = Nair,RT/Nwater- (A.1)

The nominal environmental fatigue correction factor, Fennoir, for carbon steels is expressed as Fennom = exp(0.632 - 0.101 S* T* 0* t (A.2) and for low-alloy steels, it is expressed as Fen,norn = exp(0.702 - 0.101 S* T* 0* t ) (A.3) where S*, T*, 0*, and t *are transformed S content, temperature, DO level, and strain rate, respectively, defined as:

S*:

S: *0.001 (S < 0.00o1 wt.%)

.S (S < 0.015 wt.%)

S*: .0.015 (S > 0.015 wt.%) (A.4)

T* 0 (T < 150 0C)

T*= T - 150 (T = 150-350°C) (A.5)

O*

0* =0 (DO < 0.04 ppm) 0*

ln(DO/0.04) (0.04 ppm < DO < 0.5 ppm) ln(12.5) (DO > 0.5 ppm) (A.6)

=0 ( > 1%/s)

=ln(t) (0.001 < ý < l%/s) ln(0.001) (t < 0.001%/s). (A.7)

A. I

For both carbon and low-alloy steels, a threshold value of 0.07% for strain amplitude (one-half the strain I

range for the cycle) is defined, below which environmental effects on the fatigue life of these steels do not occur. Thus, I

Fen,nom = 1 (Fa < 0.07%). (A.8)

For wrought and cast austenitic stainless steels, I Fen,nom = exp(0.734 - T'0' t').

where T', t', and 0' are transformed temperature, strain rate, and DO level, respectively, defined as:

(A.9)

I T'=0 T'= (T- 150)/175 (T < 150°C)

(150 <T < 325-C)

I T'= I (T >_325-C) (A.10) t= 0 (t > 0.4%/s)

I I=n( t/0.4) (0.0004 < t < 0.4%/s)

= ln(0.0004/0.4) (a < 0.0004%/s) (A. 11) I O' = 0.281 (all DO levels). (A.12)

For wrought and cast austenitic stainless steels, a threshold value of 0.10% for strain amplitude (one-half I

the strain range for the cycle) is defined, below which environmental effects on the fatigue life of these steels do not occur. Thus, Fen,nom = 1 (Sa < 0.10%). (A.13)

I For Ni-Cr-Fe alloys, I Fen,nom = exp(- T' C0 '), (A.14) where T', E' ', and 0' are transformed temperature, strain rate, and DO, respectively, defined as:

I T' T/325 T'=I (T <325-C)

(T>325-C) (A.15)

I ln(E'/5.0)

In(0.0004/5.0)

( t > 5.0%/s)

(0.0004: t

• 5.0%/s)

(t < 0.0004%/s) (A. 16)

I O'

0' 0.09 0.16 (NWC BWR water)

(PWR or HWC BWR water). (A. 17)

I For Ni-Cr-Fe alloys, a threshold value of 0.10% for strain amplitude (one-half the strain range for the cycle) is defined, below which environmental effects on the fatigue life of these alloys do not occur.

Thus, A.2 I

I

Fen,nom = I (Ea< 0.10%). (A. 18)

A3 Fatigue Evaluation Procedure The evaluation method uses as its input the partial fatigue usage factors U1, U2 , U3, .. Un, determined in Class 1 fatigue evaluations. To incorporate environmental effects into the Section III fatigue evaluation, the partial fatigue usage factors for a specific stress cycle or load set pair, based on the current Code fatigue design curves, is multiplied by the environmental fatigue correction factor:

Uen.I UI *Fen,I. (A.19)

In the Class 1 design-by-analysis procedure, the partial fatigue usage factors are calculated for each type of stress cycle in paragraph NB-3222.4(e)(5). For Class I piping products designed using the NB-3600 procedure, Paragraph NB-3653 provides the procedure for the calculation of partial fatigue usage factors for each of the load set pairs. The partial usage factors are obtained from the Code fatigue design curves provided they are consistent, or conservative, with respect to the existing fatigue c-N data.

For example, the Code fatigue design curve for austenitic SSs developed in the 1960s is not consistent with the existing fatigue database and, therefore, will yield nonconservative estimates of usage factors for most heats of austenitic SSs that are used in the construction of nuclear reactor components. Examples of calculating partial usage factors are as follows:

(1) For carbon and low-alloy steels with ultimate tensile strength <552 MPa (<80 ksi), the partial fatigue usage factors are obtained from the ASME Code fatigue design curve, i.e., Fig. 1-9.1 of the mandatory Appendix I to Section II1 of the ASME Code. As an alternative, to reduce conservatism in the current Code requirement of a factor of 20 on life, partial usage factors may be determined from the fatigue design curves that were developed from the ANL fatigue life model, i.e., Figs. A.1 and A.2 and Table A.1.

Carbon Steels UTS -552 MPa (-580 ksi)

Air up to 371 -C(70F E = 206.8 GPa Figure A.1.

103 Fatigue design curve for Code Curve .ASME U) ANL Model & Eq. 18 carbon steels in air. The curve developed from the ANL model is based on E 4 -'

factors of 12 on life and 2 U) -.J on stress.

Carbon Steels o*u= 551.6 MPa oy = 275.8 MPa

... .. . . ... . ..., . . .. I . . ..

. . ..l 101 102 103 104 105 106 107 108 109 1010 1011 Number of Cycles N A.3

I I

Low-Alloy Steels 103 SUTS "E=

<552 MPa (<80 ksi)

Air up to 371TC (700TF) 206.8GPa - Figure A.2.

I

-"-..ASME Code Curve Fatigue design curve for I

I

" I- ANLModel&Eq. 18 low-alloy steels in air: The

- Ix curve developed from the 7a E

" .--- - - , I ANL model is based on 102

-Low-Alloy Sel factors of 12 on life and 2 on stress. I

• o IA I .1 .. ... ... - i-. __I ay=482.6MPa I

101 102 103 104 105 106 107 108 109 1010 1011 Table A.1.

Number of Cycles N Fatigue design curves for carbon and low-alloy steels and proposed extension to 1011 cycles.

I ASME Code Stress Amplitude (MPa/ksi)

Eqs. 15 & 18 Eqs. 16 & 18 Carbon Steel Low-Alloy Steel Cycles ASME Code Curve Stress Amplitude (MPa/ksi)

Eqs. 15 & 18 Eqs. 16 & 18 Carbon Steel Low-Alloy Steel I

Cycles Curve I E+0l 2 E+01 5 E+O1 1 E+02 3999 (580) 2827 (410) 1896 (275) 1413 (205) 5355 (777) 3830 (556) 2510 (364) 1820 (264) 5467 (793) 3880 (563) 2438 (354) 1760 (255) 2 E+05 5 E+05 1 E+06 2 E+06 114(16.5) 93 (13.5) 86(12.5) 176 (25.5) 154 (22.3) 142 (20.6) 130 (18.9) 141 (20.5) 116(16.8) 106 (15.4) 98 (14.2)

I 2 E+02 1069(155) 1355 (197) 1300(189) 5 E+06 120(17.4) 94(13.6) 5E+02 1 E+03 2 E+03 724(105) 572 (83) 441 (64) 935(136) 733(106) 584 (84.7) 900(131) 720(104) 576 (83.5) 1 E+07 2 E+07 5 E+07 76.5(11.1) 115(16.7) 110(16.0) 107 (15.5) 91 (13.2) 90(13.1) 88(12.8)

I 5 E+03 331 (48) 451 (65.4) 432 (62.7) 1 E+08 68.3 (9.9) 105 (15.2) 87(12.6) 1 E+04 2 E+04 5 E+04 262 (38) 214 (31) 159(23) 373 (54.1) 305 (44.2) 238(34.5) 342 (49.6) 276 (40.0) 210(30.5) 1 E+09 1 E+010 1 E+01 1 60.7 (8.8) 54.5 (7.9) 48.3(7.0) 102 (14.8) 97(14.1) 94(13.6) 83 (12.0) 80(11.6) 77(11.2)

I 1 E+05 138 (20.0) 201 (29.2) 172 (24.9)

(2) For wrought or cast austenitic SSs and Ni-Cr-Fe alloys, the partial fatigue usage factors are obtained from the new fatigue design curve proposed in the present report for austenitic SSs, i.e.,

I Fig. A.3 and Table A.2.

The cumulative fatigue usage factor, Uen, considering the effects of reactor coolant environments is I

then calculated as the following:

Uen = Ul'Fen,1 + U2"Fen,2 + U3"Fen,3 + Ui'Fen,i ... + Un'Fen,n, (A.20)

I where Fen,i is the nominal environmental fatigue correction factor for the "i"th stress cycle (NB-3200) or load set pair (NB-3600). Because environmental effects on fatigue life occur primarily during the tensile- I loading cycle (i.e., up-ramp with increasing strain or stress), this calculation is performed only for the tensile stress producing portion of the stress cycle constituting a load pair. Also, the values for key parameters such as strain rate, temperature, dissolved oxygen in water, and for carbon and low-alloy steels S content, are needed to calculate Fen for each stress cycle or load set pair. As discussed in I

Sections 4 and 5 of this report, the following guidance may be used to determine these parameters:

A.4 I

I

Austenitic Stainless Steel Air up to 371 'C (700TF)

..... ASME Code Curve i New Design Curve Based on the, ANI Mnoell Figure A.3.

Fatigue. design curve for 0-austenitic stainless steels E in air.

(A, E = 195.1 GPa 65

- ~ __________

102 _ = 648.1 MPa _________-L y= 303.4 MPa

. . .. . I . . . . . .. I -- ] .I.I 7 9 101 102 103 104 105 106 10 108 10 1010 1011 Number of Cycles N Table A.2. The new and current Code fatigue design curves for austenitic stainless steels in air.

Stress Amplitude (MPa/ksi) Stress Amplitude (MPa/ksi)

Cycles New Design Curve Current Design Curve Cycles New Design Curve Current Design Curve I E+01 6000 (870) 4881 (708) 2 E+05 168 (24.4) 248 (35.9) 2 E+01 4300 (624) 3530 (512) 5 E+05 142 (20.6) 214 (31.0) 5 E+01 2748 (399) 2379 (345) I E+06 126 (18.3) 195 (28.3) 1 E+02 1978 (287) 1800 (261) 2 E+06 113(16.4) 157 (22.8) 2 E+02 1440 (209) 1386 (201) 5 E+06 102 (14.8) 127(18.4) 5 E+02 974 (141) 1020 (148) I E+07 99(14.4) 113 (16.4) 1 E+03 745 (108) 820 (119) 2 E+07 105 (15.2) 2 E+03 590 (85.6) 669 (97.0) 5 E+07 98.6(14.3) 5 E+03 450 (65.3) 524 (76.0) I E+08 97.1 (14.1) 97.1 (14.1)

I E+04 368 (53.4) 441 (64.0) I E+09 95.8 (13.9) 95.8 (13.9) 2 E+04 300 (43.5) 383 (55.5) I E+10 94.4 (13.7) 94.4 (13.7) 5 E+04 235 (34.1) 319 (46.3) I E+1 1 93.7 (13.6) 93.7 (13.6)

I E+05 196 (28.4) 281 (40.8) 2 E+10 (1) An average strain rate for the transient always yields a conservative estimate of Fen. The lower bound or saturation strain rate of 0.001%/s for carbon and low-alloy steels or 0.0004%/s for austenitic SSs can be used to perform the most conservative evaluation.

(2) For the case of a constant strain rate and a linear temperature response, an average temperature (i.e.,

average of the maximum and minimum temperatures for the transients) may be used to calculate Fen. In general, the "average" temperature that should be used in the calculations should produce results that are consistent with the results that would be obtained using the modified rate approach described in Section 4.2.14 of this report. The maximum temperature can be used to perform the most conservative evaluation.

(3) The DO value is obtained from each transient constituting the stress cycle. For carbon and low-alloy steels, the dissolved oxygen content, DO, associated with a stress cycle is the highest oxygen level in the transient, and for austenitic stainless steels, it is the lowest oxygen level in the transient.

A value of 0.4 ppm for carbon and low-alloy steels and 0.05 ppm for austenitic stainless steels can be used for the DO content to perform a conservative evaluation.

A.5

I I

(4) The sulfur content, S, in terms of weight percent might be obtained from the certified material test report or an equivalent source. If the sulfur content is unknown, then its value shall be assumed as the maximum value specified in the procurement specification or the applicable construction Code.

The detailed procedures for incorporating environmental effects into the Code fatigue evaluations have been presented in several articles. The following two may be used for guidance:

(1) Mehta, H. S., "An Update on the Consideration of Reactor Water Effects in Code Fatigue Initiation Evaluations for Pressure Vessels and Piping," Assessment Methodologies for Preventing Failure: I Service Experience and Environmental Considerations, PVP Vol. 410-2, R. Mohan, ed., American Society of Mechanical Engineers, New York, pp. 45-51, 2000.

(2) Nakamura, T., M. Higuchi, T. Kusunoki, and Y. Sugie, "JSME Codes on Environmental Fatigue I Evaluation," Proc. of the 2006 ASME Pressure Vessels and Piping Conf., July 23-27, 2006, Vancouver, BC, Canada, paper # PVP2006-ICPVTI 1-93305. 3 I

I I

I I

I I

I I

I A.6 I

NRC FORM 335 U. S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER (2-89) (Assigned by NRC. Add Vol., Supp., Rev.,

NRCM 1102, and Addendum Numbers, if any.)

3201,3202 BIBLIOGRAPHIC DATA SHEET (See instructionson the reverse) NUREG/CR-6909

2. TITLE AND SUBTITLE ANL-06/08 Effect of LWR Coolant Environments on the Fatigue Life of Reactor Materials 3. DATE REPORT PUBLISHED Final Report MONTH YEAR February 2007

4. FIN OR GRANT NUMBER N6187

5. AUTHOR(S) 6. TYPE OF REPORT Technical

0. K. Chopra and W. J. Shack 7. PERIOD COVERED (Inclusive Dates)

8. PERFORMING ORGANIZATION - NAME AND ADDRESS (if NRC, provide Division, Office or Region, U.S. Nuclear Regulatory Commission, and mailing address; if contractor, provide name and mailing address.)

Argonne National Laboratory 9700 South Cass Avenue Argonne, IL 60439

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, and mailing address.)

Division of Fuel, Engineering, and Radiological Research Office of Nuclear Regulatory Research U.S. Nuclear Regulatory Commission Washington, DC 20555-0001

10. SUPPLEMENTARY NOTES H. J. Gonzalez, NRC Project Manager

11. ABSTRACT (200 words or less)

The existing fatigue strain-vs.-life (c-N) data illustrate potentially significant effects of LWR coolant environments on the fatigue resistance of pressure vessel and piping steels. Under certain environmental and loading conditions, fatigue lives in water relative to those in air can be a factor ofl z2 lower for austenitic stainless steels, z3 lower for Ni-Cr-Fe alloys, and z17 lower for carbon and low-alloy steels. This report summarizes the work performed at Argonne National Laboratory on the fatigue of piping and pressure vessel steels in LWR environments. The existing fatigue c-N data have been evaluated to identify the various material, environmental, and loading parameters that influence fatigue crack initiation, and to establish the effects of key parameters on the fatigue life of these steels. Statistical models are presented for estimating fatigue life as a function of material, loading, and environmental conditions. The environmental fatigue correction factor for incorporating the effects of LWR environments into ASME Section III fatigue evaluations is described. The report also presents a critical review of the ASME Code fatigue design margins of 2 on stress (or strain) and 20 on life and assesses the possible conservatism in the current choice of design margins.

12. KEY WORDS/DESCRIPTORS (List words orphrases that will assist researchers in locating this report.) 13. AVAILABILITY STATEMENT Unlimited Fatigue crack initiation 14. SECURITY CLASSIFICATION Fatigue life (This Page)

Environmental effects Unclassified Carbon and low-alloy steels (This Report)

Austenitic stainless steels Unclassified Ni-Cr-Fe alloys 15. NUMBER OF PAGES BWR environment PWR environment 16.ICE PR NRC FORM 335 (2-89)

Official Transcript of Proceedings NEC-JH-27 NUCLEAR REGULATORY COMMISSION

Title:

Advisory Committee on Reactor Safeguards Subcommittee on Materials, Metallurgy and Reactor Fuels Docket Number: (not applicable)

Location: Rockville, Maryland Date: Wednesday, December 6, 2006 Work Order No.: NRC-1 347 Pages 1-162 NEAL R. GROSS AND CO., INC.

Court Reporters and Transcribers 1323 Rhode Island Avenue, N.W.

Washington, D.C. 20005 (202) 234-4433

1 1 UNITED STATES OF AMERICA 2 NUCLEAR REGULATORY COMMISSION 3 +++++

4 ADVISORY COMMITTEE ON REACTOR SAFEGUARDS 5 SUBCOMMITTEE ON MATERIALS, METALLURGY, AND 6 REACTOR FUELS 7 .+.+++

8 WEDNESDAY, 9 December 6, 2006 10 11 The meeting was convened in Room T-2B3 of 12 Two White Flint North, 11545 Rockville Pike, 13 Rockville, Maryland, at 1:30 p.m., Dr. J. Sam Armijo, 14 Chairman of the subcommittee, presiding.

15 MEMBERS PRESENT:

16 J. SAM ARMIJO, CHAIRMAN 17 MARIO V. BONACA, ACRS MEMBER 18 SAID ABDET KHALIK, ACRS MEMBER 19 SANJOY BANERJEE, ACRS MEMBER 20 THOMAS S. KRESS, ACRS MEMBER 21 JOHN D. SIEBER, ACRS MEMBER 22 GRAHAM WALLIS, ACRS MEMBER 23 CHARLES G. HAMMER, DESIGNATED FEDERAL OFFICIAL 24 CAXETANO SANTOS, ACRS STAFF 25 NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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2 1 INDEX 2 Opening Remarks, S. Armijo, ACRS 3 3 Overview of Regulatory Guide 1.207 (DG-1144) 4 H. Gonzalez, RES 4 5 Discussion of Technical Basis for Regulatory 6 Guide 1.207 and NUREG/CF-6909, 0. Chopra, 7 Argonne National Laboratory 11 8 Discussion of Public Comments and Staff 9 Responses for Regulatory Guide 1.207 10 and NUREG/CR-6909, H. Gonzalez, RES and 11 0. Chopra, ANL 77 12 ASME Presentation, E. Ennis, ASME 85 13 AREVA Presentation, D. Cofflin, AREVA 142 14 Adjourn 161 15 16 17 18 19 20 21 22 23 24 25 NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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3 I

1 P-R-O-C-E-E-D-I-N-G-S 2 1:31 P.M.

3 CHAIRMAN ARMIJO: The meeting will now 4 come to order. This is a meeting of the Materials, 5 Metallurgy and Reactor Fuels Subcommittee. My name is 6 Sam Armijo, Chairman of the Committee. ACRS Members I 7 in attendance are Dr. Mario Bonaca, Mr. Jack Sieber, 8 Dr. Bill Shack is sitting as a member of the audience 9 or staff at this point, Dr. Thomas Kress and Dr.

10 Graham Wallis are also present.

11 Gary Hammer of the ACRS staff is the 12 Designated Federal Official for this meeting. 3 13 The purpose of this meeting is to discuss 14 Regulatory Guide 1.207, guidelines for evaluating I 15 fatigue analyses incorporating the life reduction of 3 16 metal components due to the effects of light-water 17 reactor environments for new reactors. We will hear I 18 presentations from the NRC's Office of Nuclear 19 Regulatory Research and their contractor, Argonne 20 National Laboratory. I 21 We will also hear presentations from 22 representatives of the American Society of Mechanical 23 Engineers and AREVA. 3 24 The Subcommittee will gather information, 25 analyze relevant issues and facts, and formulate I NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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4 1 proposed positions and actions, as appropriate for 2 deliberation by the Full Committee.

3 The rules for participation in today's 4 meeting have been announced as part of the notice of 5 this meeting previously published in the Federal 6 Register. We have received no written comments from 7 members of the public regarding today's meeting.

8 A transcript of the meeting is being kept 9 and will be made available as stated in the Federal 10 Register notice. Therefore, we request that 11 participants in this meeting use the microphones 12 located throughout the meeting when addressing the 13 Subcommittee.

14 Participants should first identify 15 themselves and speak with sufficient clarity and 16 volume so that they may be readily heard.

17 We will now proceed with the meeting and 18 I call on Mr. Hipolito Gonzales of the Office of 19 Nuclear Regulatory Research to begin.

20 MR. GONZALEZ: Thank you. I am Hipolito 21 Gonzalez. I'm the Project Manager for Regulatory 22 Guide 1.207. I'm from the Corrosion and Metallurgy 23 Branch and with me, Omesh Chopra. He's from Argonne 24 National Lab. He's going to be presenting part of the 25 regulatory basis, technical regulatory basis.

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5 1 I would like to acknowledge William Cullen 2 from the Office of Research and John Ferrer, NRR, for 3 their helpful reviews and comments on this project.

4 Next slide.

5 The agenda today, we're going to be 6 discussing Regulatory Guide 1.207. I'm going to give 7 a quick historical perspective and then we're going to 8 go over an overview the reg. guide. And then Omesh 9 will present the technical basis which is the NUREG 10 report CR, NUREG CR 6909, Revision 1.

11 I'm .going to give a summary of the 12 regulatory positions. And the last presentation is 13 going to be the resolution of public comments.

14 The ASME Section 3, fatigue design curves 15 were developed in the late 1960s and the early 1970s.

16 The tests conducted were in laboratory environments at 17 ambient temperatures. And the design curves included 18 adjusted factors of 2 constraint and 20 on cyclic life 19 to account for variations in materials, surface 20 finish, data scatter and size.

21 Results from the studies in Japan and 22 others in ANL, Argonne National Lab, as illustrated.

23 Potential significant effects of the light-water 24 reactor coolant environment on the fatigue life of the 25 steel, steel components.

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1 Next slide.

2 Since the late 1980s, the NRC staff has 3 been involved in the discussion with ASME co-4 committees, the PVRC and Technical Community to 5 address the issues related to the environmental 6 effects on fatigue.

7 In 1991, the ASME Board of Nuclear Code 8 and Standards requested the PVRC to examine worldwide 9 fatigue strain versus like data and develop 10 recommendations.

11 In 1995, it was resolution for GSI 166 12 which established that the risk to core damage from 13 fatigue failure of the reactor coolant system was 14 small. So no action was required for current plant 15 design life of 40 years. Also, the NRC staff 16 concluded that fatigue issues should be evaluated for 17 extended period of operation for license renewal and 18 this is under GSI-190.

19 In 1999, we had GSI-190 and the fatigue 20 evaluation of metal components for 60-year life plant, 21 plant life. Staff concluded that consistent with 22 requirements of 10 CFR 54.21, that aging management 23 programs for license renewal should address components 24 of fatigue including the effects of the environment.

25 On December 1, 1999, by letter to the NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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7 1 Chairman of the ASME Board of Nuclear Code and 2 Standards, the NRC requested ASME to revise the code 3 to include the environmental effects on the fatigue 4 design components.

5 Next slide.

6 ASME initiated the PVRC Steering Committee 7 on cyclic life and environmental effects and the PVRC 8 Committee recommended revising the code for design 9 fatigue curves. This was to WRC Bulletin 487.

10 After more than 25 years of deliberation, 11 there hasn't been any consensus regarding 12 environmental effects on fatigue life on the light-13 water reactor environments.

14 The NRR requested research under user need 15 requests to 504 to develop guidance for determining 16 the acceptable fatigue life of ASME pressure boundary 17 components with consideration of the light water 18 reactor environment and this guidance will be used for 19 supporting reviews of application that the Agency 20 expects to receive for new reactors. The industry was 21 immediately notified that the NRC staff initiated this 22 work, the development of the reg. guide. In addition, 23 this is one of the high priority reg. guides to be 24 completed by March 2007.

25 In February and August this year, NRC NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., NW.

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8 1 staff and ANL, we had presented at the ASME Code 2 Meetings the technical basis draft, NUREG CR6909. On 3 July 24, 2006, both the draft reg. guide and the NUREG 4 technical basis report were published for public 5 comments and the public comment period ended September 6 25.

7 In addition, on July 25, ANL presented a 8 paper on the technical basis again.

9 CHAIRMAN ARMIJO: Just to clarify 10 something, new reactors, does that include -- do these 11 rules apply to already certified design, such as the 12 ABWR and the API000? Are they grandfathered by virtue 13 of their certification?

14 MR. FERRER: This is John Ferrer from NRR 15 staff. They're grandfathered by virtue of their 16 certification that's already been addressed in the 17 reviews there, so we're not backfitting this reg.

18 guide to those certified designs.

19 DR. SIEBER: For 40 years though.

20 CHAIRMAN ARMIJO: Well, actually, if you 21 read the safety evaluation, the way it was written 22 said that. they were evaluated for 60 years.

23 DR. SIEBER: Okay.

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9 1 in the United States and if they were being certified 2 after this reg. guide is issued, that would be the 3 rule -- that would control the design, wouldn't it?

4 MR. FERRER: I wish I -- I agree with you.

5 Unfortunately, the way certified design works is once 6 we certify it, we'd have to go through a backfit 7 evaluation if we were going to apply this. And what 8 happened in the backfit evaluation, if you go back a 9 couple of slides on the GSI-166 and the GSI-190, we 10 did a backfit evaluation and showed the risk was not 11 high enough to justify a backfit, but the reason we 12 implemented it on license renewal was the fact that 13 the probability of leakage increased significantly 14 within 40 and 60 years.

15 But again, the risk which is the 16 probability of getting a pipe rupture that would lead 17 to core damage was still low.

18 CHAIRMAN ARMIJO: Thank you.

19 MR. GONZALEZ: Now I am going to go to an 20 overview of the reg. guide.

21 Next slide.

22 How the reg. guide 1.207 relates to the 23 regulatory requirements. GDC criterion, general 24 design criterion 1, quality standards and waivers.

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10 1 structures and components must be designed, 2 fabricated, erected and tested to the quality standard 3 commensurate with the importance of the safety 4 function performed.

5 GDC-30 states, in part, that components 6 included in a reactor pressure boundary must be 7 designed, fabricated, erected and tested to the 8 highest practical quality standards.

9 In 10 CFR 50.55A endorses the ASME boiler 10 pressure vessel code for design of safety-related 11 systems and components. These are Class 1 components.

12 ASME Code Section 3 includes the design 13 fatigue, includes the fatigue design curves. But 14 these fatigue design curves do not address the impact 15 of the reactor coolant system environment.

16 The objective of this regulatory guide is 17 to provide guidance for determining the acceptable 18 fatigue life of ASME pressure boundary components with 19 the consideration of the light water reactor 20 environment for major structural materials that will 21 be carbon steel, low-alloy steels, austenitic 22 stainless steel and nickel-based alloys. For example, 23 alloy-600, 690.

24 So in this guide, describes an approach 25 that the NRC staff considers acceptable to support NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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11 1 reviews about the applications that the Agency expects 2 to receive for new reactors.

3 Implementation, this will only apply to 4 new plants. And no backfitting is intended. And this 5 is due to the conservatism in the current fleet of 6 reactors because of the design practices for fatigue 7 work conservatisms all plants were designed.

8 Next slide, please.

9 Now I'm going to -- how the technical 10 basis was developed. Omesh is going to give the 11 presentation on the technical basis report.

12 MR. CHOPRA: Thanks, Hipo.

13 DR. BONACA: I have a question regarding 14 your last statement. No backfitting is intended, 15 conservatism on coolant reactors. If the approach was 16 conservative on coolant reactors, I mean could it be 17 used also for new reactors?

18 MR. FERRER: Let me try to answer that.

19 In reviewing GSI-166 which was backfit to current 20 operating plants, we evaluated the as-existing fatigue 21 analyses and there were a number of conservatisms in 22 the specification of transients and the methodology 23 and the analysis.

24 We don't know whether or not that same 25 conservatism will be applied in the new reactors. In NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., NW.

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12 1 addition, there have been some changes in the ASME 2 code criteria since those original analyses were done 3 that removed some of the conservatisms in the 4 analysis. So if somebody were to do code analysis to 5 the current code criteria may not have the same level 6 of conservatisms.

7 DR. BONACA: I understand. Thank you.

8 MR. CHOPRA: The issue we are discussing 9 here today is effect of light water reactor coolant 10 environments on the fatigue life of structural steels.

11 Over the last 20 to 30 years, there's been sufficient 12 data accumulated, both in the U.S. and worldwide, 13 especially in Japan, which shows that coolant 14 environments can have a significant effect on the 15 fatigue life of these steels.

16 And this data is very consistent. It 17 doesn't matter where it has been rated, all show 18 similar trends without any exception. And also, the 19 fatigue data is consistent with a much larger database 20 on fatigue crack growth rates affect on environment of 21 fatigue crack growth rates. There's no inconsistency.

22 The mechanisms are very similar and both show similar 23 trends, effects of radius parameters, material loading 24 and environmental parameters have similar inference on 25 fatigue crack initiation and fatigue crack growth.

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13 1 And this fatigue data has been evaluated 2 to clearly define which are the important parameters.

3 They're well defined and also the range of these 4 parameters for which environmental effects are 5 significant, it's clearly defined.

6 So we know the conditions under which 7 environment would have an effect on fatigue life. The 8 question. is do these conditions exist in the fleet?

9 If they exist, we will have an effect on the 10 environment and it should be considered. We know from 11 subsection 31.32.21 that the current fatigue design 12 curves do not include the effect of aggi-essive 13 environment which can accelerate fatigue failures and 14 has to be considered.

15 So the burden is on the designer to better 16 define these transients, to know what conditions 17 occurred during these transients and whether 18 environment would be involved.

19 Next, before getting into the 20 environmental. effects, I just want to cover a few 21 background information. We are talking about the 22 effect of environment on fatigue life. Let's 23 understand what do we mean by fatigue life? The 24 current code design curves were based on data which 25 was where the specimens were tested to failure. Quite NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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14 1 often, these design curves are termed as failure 2 codes, but I think the intent was to define fatigue 3 life as to prevent fatigue crack initiation, because 4 the data which has been obtained in the last 20 to 30 5 years in these results fatigue life is defined as the 6 number of sitings for the peak load to decrease by 25 7 percent.

8 And for the type of specimen, size of 9 specimens used in these tests, mostly quarter inch or 10 three-eighth round cylindrical specimens, this would

.11 correspond to creating a three millimeter crack. So 12 we can say the fatigue life is the number of cycles 13 for a given strain condition to initiate a three 14 millimeter crack and from several studies we know that 15 surface crack, about 10 micron deep form quite early 16 during fatigue cycling.

17 So we can say that fatigue life is nothing 18 but it's associated with growth of these cracks from 19 a 10 micron size to 3 millimeter size and typically 20 this is the behavior of the growth of these cracks is 21 in this shape where crack length is a fraction of 22 fatigue life varies like this and it's divided into 23 two stages, initiation stage and a propagation stage.

24 Initiation stage is characterized by decrease in crack 25 growth rates. It's very sensitive to micro structure.

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15 1 It involves sheer crack growth which is 45 degrees.to 2 the stress axis, whereas propagation stage is not very 3 sensitive to microstructure. It was tensile crack 4 growth which is perpendicular to the stress axis and 5 this is the stage where you see on the fracture 6 surface well defined striations..

7 Various studies have shown that this 8 transition from an initiation stage to a propagation 9 stage occurs around -- depending on the material, 150 10 micron or 300 micron, that range.

11 So initiation stage is growth of crack up 12 to 300 microns. Propagation stage is beyond that to 13 3000 or 3 millimeter size.

14 Next slide.

15 CHAIRMAN ARMIJO: Before you leave that 16 curve,, just for the benefit of people who don't 17 understand these curves, what is the time difference 18 between or the fatigue life difference from the three 19 millimeter crack initiated crack to through-wall 20 failure in the case of let's say a one-inch pipe, one-21 inch wall thickness?

22 MR. CHOPRA: We would use the crack growth 23 rate data.

24 CHAIRMAN ARMIJO: Would that typically 25 increase the number of cycles by a factor of 2 or a CNEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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16 1 factor of 10?

2 MR. CHOPRA: It depends on the conditions, 3 loading conditions and environment and so on. So we 4 know what the crack growth rates are for various 5 conditions. So we have to use that. But maybe I can 6 answer another way. In a test specimen, the 7 difference between 25 percent load drop and complete 8 failure of a specimen is very small. It's less than 9 one or two percent.

10 So whether we call it failure of a 11 specimen or defining it 25 percent drop, would be very 12 small difference. The idea of using 25 percent load 13 drop was to be consistent so that we define life as 14 some consistent -- all the labs do the same thing. So 15 that was the idea.

16 Otherwise, for a real component, if we 17 deal with three millimeter steel in a tube, it would 18 depend on crack growth rates.

19 CHAIRMAN ARMIJO: Okay.

20 MR. CHOPRA: Now the same curve I've 21 plotted a slightly different way where I plotted still 22 our cracked growth rates was the crack depths, 23 decreasing growth rates in the initiation stage and 24 increasing growth rates.

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17 1 on applied stress ranges. The higher the stress 2 range, the higher the crack growth. The delta sigma 3 one at very low stresses, the cracks which form during 4 cyclic loading may not growth to large enough size 5 that they can -- the propagation stage takes over.

6 DR. WALLIS: Crack velocity is really 7 growth rate and microns per cycle, not per unit of 8 time.

9 MR. CHOPRA: Right, but depending on the 10 time period one could convert it to --

11 DR. WALLIS: I know, but velocity is a 12 strange word.

13 MR. CHOPRA.: Yes, maybe this should be 14 crack growth rate.

15 DR. WALLIS: If there's no cycling, 16 there's no crack growth.

17 MR. CHOPRA: Yes, yes. Beta sigma one, 18 when the stresses are very low, cracks may grow to 19 large enough size for the propagation to take over and 20 this is known as the fatigue limit of the material.

21 This is true for constant loading.

22 MR. BANERJEE: What's the mechanism that 23 changes the velocity so much?

24 MR. CHOPRA: Initial sheer crack growth.

25 It will extent maximum couple of degrees. So it's a NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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18 1 sheer crack growth, 45 degrees, whereas, once you go 2 deep enough, large enough size, you get into a 3 different process where actually fracture mechanics 4 methodology can be used to express that. It's a 5 tensile crack growth.

6 MR. BANERJEE: It's a multi-grain sort of 7 size and then it starts -- a different mechanism.

8 MR. CHOPRA: Typically, a couple of 9 grains. Fatigue limit is applicable only under 10 constant stress conditions. If we have random 11 loading, as in the case of a real component, then we 12 can have situations where we have higher stresses, few 13 cycles of higher stresses, where cracks can grow 14 beyond this depth that you can grow even at stresses 15 which are much lower than fatigue limit.

16 So the history of cycling is also 17 important for evaluating fatigue damage.

18 DR. WALLIS: Delta sigma is the magnitude 19 of this?

20 MR. CHOPRA: Of the stress range, applied 21 extracted stress range. And environment also.

22 DR. WALLIS: Does it matter if it's 10 23 silo or compressible?

24 MR. CHOPRA: On the tests which are used 25 for obtaining fatigue data, the strain range ratio is NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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19 1 -1, completely reversed. So we go from tensile to 2 compressive.

3 Even in environment, corrosion processes 4 can cause the cracks to grow beyond this and then 5 . propagation can take over. So environment also could 6 accelerate. So the question is which part -- which of 7 these stages is affected by environment? Initiation 8 or propagation, or both?

9 DR. WALLIS: Your scales are linear, are 10 they?

11 MR. CHOPRA: This is a schematic.

12 DR. WALLIS: Schematic.

13 MR. CHOPRA: This portion is plotted here 14 where I have actual numbers. And I just wanted to 15 show you that we know from crack growth studies that 16 crack growth rates are affected by environment and 17 it's very well documented.

18 DR. WALLIS: These data look unreasonably 19 well behaved for materials data.

20 (Laughter.)

21 MR. CHOPRA:. If we plotted a few tests, we 22 will see this happen.

23 CHAIRMAN ARMIJO: Agreement is log, log.

24 DR. WALLIS: Even so, I mean.

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20 1 is also, has been studied in fatigue crack initiation.

2 DR. WALLIS: These are real data?

3 MR. CHOPRA: These are real data. But we 4 have calculated the crack growth rates in the fatigue 5 samples by benchmarking the fatigue crack front at 6 different stages during fatigue life. And so we can 7 see the three environments here: high oxygen -- high 8 dissolved oxygen water; low dissolved oxygen; PWR 9 water and air. And we see if you take 100 micron 10 crack length and air -- it took about 3,000 cycles to 11 reach that. In water, it took only 40 cycles, which 12 gives me an average growth rate of 2.5 micron per 13 cycle and this is this region here, average of this.

14 In this case, it's .0033 microns per 15 cycle. So *we see two orders of magnitude effect of 16 environment which suggests that even the initiation 17 stage may be affected even more than what crack growth 18 rate is affected.

19 I just wanted to show you that both stages 20 are affected by the environment, even the growth of 21 very small cracks.

22 Now next, the design curves, what do the 23 design curves --

24 DR. WALLIS: Presumably, this is not just 25 one batch of data like this.

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21 1 MR. CHOPRA: There's lots of data. I'm 2 just giving --

3 DR. WALLIS: There's a whole lot of data.

4 MR. CHOPRA: I'm just giving you one set, 5 yes. There's a lot of data.

6 DR. WALLIS: Because if there were 7 uncertainty in these, these curves might switch 8 positions.

9 MR. CHOPRA: sure, but I'm just presenting 10 that data to show that environment has a large effect.

11 It's the relative difference between air and water 12 which I was trying to show, not absolute crack growth 13 rates, just to show that it took only 40 cycles in 14 high oxygen water compared to 3,000 which suggests 15 that environment has a large effect on fatigue crack 16 initiation.

17 Now the design curves, we have -- the data 18 which we have obtained is on small specimens. They 19 are absolutely smooth and they were tested in room 20 temperature air. This is what was used to generate 21 the design curves in the current, code. And all of 22 them were tested under strain control, fully reversed, 23 strain ratio of -1.

24 Now this gives me the best behavior of a 25 specimen when a crack would be initiated in a NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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22 1 specimen. To apply those results to actual reactor 2 component we need to adjust these results to account 3 for parameters or variables which we know affect 4 fatigue life, but are not included in this data. And 5 these variables are mean stress, surface finish, size, 6 loading history.

7 DR. WALLIS: Does the humidity of the air 8 make a difference?

9 MR. CHOPRA: Actually, if you look at the 10 basis document of the current code, they use a 11 subfactor which included surface roughness and 12 environment and by that environment they meant a lab, 13 well-controlled lab environment.

14 DR. WALLIS: Does the humidity of the air 15 make a difference?

16 MR. CHOPRA: In some cases it would, but 17 again, that is not studied as a -- it's not addressed 18 as an explicit parameter in defining fatigue life.

19 All data which was used was room temperature air to 20 generate the design curves.

21 DR. WALLIS: Room temperature means 20 22 degrees Centigrade or something?

23 MR. CHOPRA: Yes, 25, yes. To account for 24 these other variables like mean stress, surface 25 roughness and so on, what the current code --

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23 I

1 DR. WALLIS: I'm sorry, when you -- maybe 2 you just said it. When you say PWR water, you mean at 3 room temperature or -- I 4 MR. CHOPRA: No, no. The design curves do 3 5 not address environment at all.

6 DR. WALLIS: But your data that you showed I 7 us, the well-behaved data.

8 MR. CHOPRA: Those are higher 9 temperatures.

10 DR. WALLIS: Those are higher 11 temperatures.

12 MR. CHOPRA: They would be at reactive 13 temperatures.

14 DR. WALLIS: Okay. Could be a temperature U 15 effect as well as an environment effect? 3 16 MR. CHOPRA: There is and I'll come to 17 that actually. In water, temperature is a very 18 important parameter. And to convert this data on 19 specimens to a real component, what the current code 20 does now is take the best -- 1 21 DR. WALLIS: Is the PWR water that is 22 borated at initial strength or something?

23 MR. CHOPRA: PWR is. It both has boron 24 and lithium.

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24 1 condition throughout the cycle?

2 MR. CHOPRA: Right, right. Typically, 3 people test around 1,000 ppm boron and 2ppm lithium.

4 TO adjust these curves to an actual 5 reactor component, what the code does is we take the 6 best of the specimen data and adjust it for mean 7 stress correction and then apply these adjustment 8 factors of two on stress. We decrease the specimen 9 curve by a factor of two on stress and 20 on life, 10 whichever is the lower gets the design curve. But as 11 1 mentioned, it does not include the effect of an 12 aggressive environment. In this case, what we are 13 talking about is light-water reactor environments.

14 Now to summarize some of the effects of 15 environment on carbon and low-alloy steels, there are 16 several parameters which are important. Steel type, 17 all of the data shows irrespective of steel type, it 18 doesn't matter which grade of carbon steel or low-19 alloy steel, effect of environment is about the same.

20 There is a strain threshold below which environments 21 do not -- environmental effects do not occur. And 22 this threshold is very close to slightly above the 23 fatigue life of the steel. Strain rate is an 24 important parameter. There is a threshold, 1 percent 25 per second above that. Environmental effects are more NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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25 1 great and lower the strain rate, higher the effect.

2 And it diffuses the saturation at around .001 percent 3 per second.

4 Similarly, temperature is very important.

5 Once again, there is a threshold; 150 degree C.

6 Higher temperatures, there's greater effect. Below 7 150 --

8 DR. WALLIS: Strain rate's lowest point is 9 .001 percent a second makes a difference?

10 MR. CHOPRA: Yes. I'll show you some of 11 the results.

12 DR. WALLIS: Really? That's awfully slow, 13 isn't it?

14 MR. CHOPRA: Some of the transients are.

15 DR. WALLIS: Abnormally slow.

16 MR. CHOPRA: Temperature also, there is 17 only a moderate effect below 150. Typically, when I 18 mean moderate effect, up to a factor of 2. Any water 19 touched surface may have up to a factor of --

20 DR. WALLIS: Linear decrease doesn't tell 21 me how fast it is. Linear decrease in life after 150 22 doesn't tell me how rapidly it decreases.

23 MR. CHOPRA: There are some slides, I'll 24 show you how much of a different it is.

25 MR. SANTOS: Do you have an equation?

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26 1 MR. CHOPRA: Yes.

2 DR. WALLIS: Which goes right through the 3 data?

4 MR. CHOPRA: Absolutely.

5 DR. WALLIS: Is this an Argonne equation 6 or a universal equation?

7 CHAIRMAN ARMIJO: You'll see.

8 DR. WALLIS: We'll see, okay.

9 MR. CHOPRA: Dissolved oxygen is also 10 similar. There's a threshold. In this case, low 11 oxygen environmental effects on carbon low-allow 12 steels-are less. There's a threshold .04 ppm. Higher 13 dissolved oxygen has an environmental effect, 14 saturates around .05 ppm.

15 DR. WALLIS: How much sulfur is there in 16 the reactor?

17 CHAIRMAN ARMIJO: That's in the steel.

18 DR. WALLIS: In the steel, I'm sorry. I 19 thought you were talking about the environment. Now 20 you're talking about the steel?

21 MR. CHOPRA: These are --

22 DR. WALLIS: Dissolved oxygen in the 23 steel.

24 MR. CHOPRA: These are loading parameters.

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27 1 parameters.

2 DR. WALLIS: Okay.

3 MR. CHOPRA: Sulfur also has a large 4 effect on fatigue crack initiation.

5 DR. WALLIS: There's no other effects, 6 copper and stuff like that? There's no other effects?

7 MR. CHOPRA: In the steel? No. At least 8 the ones which we have looked at. Sulfur is the one 9 because it deals with the mechanism. Actually, the 10 reason why these are higher for carbon and low-allow 11 steels which these are very well documented. It's the 12 sulfite iron density of the cracking. If we reach a 13 critical sulfite iron density crack enhancement 14 occurs. So these are very well documented in the 15 data. This is a mechanism. That's why sulfur is 16 important.

17 Roughness effects, we know if we have a 18 rough specimen surface it provides sites for 19 initiation. Life goes down. And in carbon low-alloy 20 steel, in air, there is an effect of surface 21 roughness, but some limited data suggests that in 22 water, rough and smooth specimens have about the same 23 life. So roughness effects may not be there for 24 carbon low-alloy steel.

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28 1 obtained on very low flow rates or semi-stagnant 2 conditions. If we do these tests in higher flow 3 rates, effect of the environment does go down. Means 4 fatigue life would increase in high flow rates by a 5 factor of about 2.

6 Similarly, the effects on austenitic 7 stainless steels, same parameters, steel type, again 8 different grades of austenitic stainless steel, 9 similar effects and even cast austenitic stainless 10 steel have similar effects on the environment.

11 Once again we see a strain threshold below 12 which there is no effect and it's very close to the 13 fatigue limit. The dependence of strain rate and 14 temperature are very similar to what we see in carbon 15 and low-alloy steels.

16 The next three, dissolved oxygen, surface 17 roughness and flow rate, the effects are very 18 different from carbon and low-alloy steels. In this 19 case, for austenitic stainless steel, it's the low 20 oxygen which gives you a larger effect. And 21 irrespective of what steel type we use or what heat 22 treatment, heat treatment that means sensitization.

23 Sensitized stainless steel or solution in the 24 stainless steel both show similar life in low oxygen.

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29 1 oxygen?

2 MR. CHOPRA: Pardon me?

3 DR. WALLIS: That extends down --

4 MR. CHOPRA: If we can achieve that, you 5 know, but typically in a PWR, we have around -- it's 6 a low -- less than 50 ppm.

7 Yes, low oxygen, irrespective of the steel 8 type or heat treatment, there's a large effect on 9 environment, but in high oxygen, non-water chemistry, 10 PWR conditions, some steels show less effect and these 11 are solution annealed high-carbon steels which are not 12 sensitized. All low carbon grades such as 316 nuclear 13 grade or 304 L may have less effect in high oxygen.

14 Surface roughness and this is both in air 15 and water environments, there's a reduction in life.

16 Even in water. In carbonate steel we did not see a 17 reduction in life for rough samples. In this case, 18 both in air and water there is an effect of roughness.

19 And flow rate, there is no effect of flow rate on 20 fatigue life for austenitic stainless steels in water.

21 The differences between these three 22 suggests that the mechanism may be different for 23 austenitic stainless steels compared to carbon and 24 low-alloy steel. I mention the mechanism for carbon 25 and low-allow steels, the sulfite iron density of the NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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30 1 crack depth. In this case, it's not well known --

2 there's no agreement on what is the mechanism. One 3 possible mechanism would be that as we expose stress 4 surface, hydrogen is created which changes the 5 definition of behavior and of the crack depth. But 6 this is one possible mechanism.

7 The next slides are details of what I 8 summarized. Unless there are specific questions, I'm 9 going to skip these next eight slides which basically 10 give the data which I summarized in the previous.

11 CHAIRMAN ARMIJO: I think it would be 12 better if you just highlight these things, just to 13 make the key points from these charts because I think 14 they're important.

15 MR. CHOPRA: This is the strain rate 16 effect. You were asking about the strain rate. I 17 plotted fatigue life for low-alloy steel, carbon steel 18 under certain conditions, strain amplitudes. In air, 19 PWR water and BWR.

20 DR. WALLIS: Are you claiming there's a 21 significant difference between air and PWR?

22 MR. CHOPRA: It's up to about a factor of 23 2 and this could be a factor of 15 or 20 lower 24 DR. WALLIS: We're not going to put in 25 that much oxygen, are we?

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31 1 MR. CHOPRA: BWR has 200 to 300 ppb oxygen 2 and in this case, there are correlations which will 3 tell you how much -- depending on the oxygen, what 4 would be the effect.

5 This is the maximum effect because this is 6 I think .7. Saturation is at .5. So this is the 7 maximum effect under these conditions.

8 This is strain threshold which I 9 mentioned, the threshold about which effect of 10 environment is there. This gives you dissolved oxygen 11 at .04, this is carbon steel, higher oxygen levels, 12 things go down. And again, in PWR there's only a 13 modern effect.

14 I mentioned that for stainless steel, the 15 effect of dissolved oxygen is different. Here, this 16 is now three or four stainless at two different 17 strainless amplitude. There are two different tests 18 at different conditions, .25 and .33 and high oxygen, 19 no effect upstream rate and low oxygen, it goes down.

20 Whereas, a 316 NG or low carbon grade shows some 21 reduction in life in high oxygen, but not at the same 22 extent as you see in low oxygen.

23 So these are just a few examples I'm 24 showing. There's a lot of data in Japan and Europe 25 which shows similar trends. This shows the effect of NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., NW.

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32 1 sensitization. Sensitization is defined as a number, 2 EPI number. Degree of sensitization is increasing and 3 same conditions. In air, low oxygen, high oxygen and 4 we see in high oxygen it decreases with degree of 5 sensitization.

6 Effect of -- this is temperature again at 7 150 and lower, depending on what are the strain rates 8 and what are the dissolved oxygen conditions. If it's 9 very low, no effect. These are low oxygen conditions, 10 no effect. High oxygen, depending on the strain rate 11 and dissolved oxygen levels to the extent of the 12 effect in pieces.

13 DR. WALLIS: You're just talking about a 14 hundred cycles there, failure.

15 MR. CHOPRA: No, a thousand. In some 16 cases in the environment, it is.

17 DR. WALLIS: Right.

18 MR. CHOPRA: There is up to a factor of 20 19 reduction in life.

20 Surface roughness again, stainless steel, 21 open circles, smooth specimens; closed circles are 22 symbols are rough samples. A factor of 3 in air, 23 factor about the same in water.

24 CHAIRMAN ARMIJO: I don't want to belabor 25 this, but I looked at these data and the one that NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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33 1 shows -- the curve on the left for the air data, the 2 right triangles. They don't go through the best fed 3 curve at all.

4 MR. CHOPRA: Actually, this is 316 NG.

5 316 NG has a steeper slope, but for convenience we are 6 using a curve for all steels.

7 CHAIRMAN ARMIJO: So that's the best fit 8 curve there is for all --

9 MR. CHOPRA: All stainless steels, all 10 grades, including high or low-carbon grades.

11 DR. WALLIS: The purpose of the ASME curve 12 is to be below all the data, is that the idea?

13 MR. CHOPRA: Once we take into account, 14 you know I mentioned those adjustment factors of 20 on 15 fatigue and 2 on stress. Once we take that into 16 account, once we do that adjustment, then we want to 17 make sure that we are above that.

18 But these are best fit curves. So they 19 give you the average behavior for all --

20 DR. WALLIS: The ASME code has a factor of 21 2 in it or something? I don't see that.

22 MR. CHOPRA: I'll come to that. Give me 23 a 24 25 DR. WALLIS: Okay. But the factor of 2 is NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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34 1 in this curve here?

2 MR. CHOPRA: No, these are --

3 CHAIRMAN ARMIJO: ASME codes.

4 MR. CHOPRA: The code curve has the factor 5 of 2.

6 DR. WALLIS: No safety factor.

7 MR. CHOPRA: This is the best fit. These 8 are showing that even --

9 DR. WALLIS: Oh, I see. So you've give u p 10 your margin of 2?

11 MR. CHOPRA: Right.

12 DR. WALLIS: Okay.

13 MR. CHOPRA: What we are saying is only 14 the margin or adjustment factors are gone for the --

15 CHAIRMAN ARMIJO: That's it.

16 MR. CHOPRA: Environment has taken care o:f 17 all that and still be within bound for a lot of other 18 factors like surface roughness and so on.

19 DR. WALLIS: You're going to tell us what 20 you're going to do about that?

21 MR. CHOPRA: Sure.

22 DR. WALLIS: Okay.

23 (Laughter.)

24 CHAIRMAN ARMIJO: Absolutely.

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35 1 flow rate. I mentioned that for carbon and low-alloy 2 steels, effect of environment is less.

3 Now a few slides for nickel alloy.

4 There's much less data on nickel alloys. Here, I've 5 plotted the data which is available --

6 DR. WALLIS: Much less data. So you're 7 showing us more than you showed us for steel?

8 MR. CHOPRA: What we do is rather than 9 coming with a new curve for nickel alloys, unless we 10 have enough data, what I'm trying to show is that we 11 can use the austenitic stainless steel to represent 12 the nickel alloys and even the few data we have for 13 alloy 690 suggests that we can use the austenitic 14 stainless steel code to determine usage factors, 15 fatigue usage factors for nickel alloys in air.

16 MR. BANERJEE: So temperature has almost 17 no effect here.

18 MR. CHOPRA: For carbon and low-alloy 19 steels there is some effect. Going from room 20 temperature to 300 may reduce life by about 50 21 percent, but stainless up to 400. There's not much 22 effect.

23 MR. BANERJEE: Including nickel alloys?

24 MR. CHOPRA: Nickel alloys, no. At 400, 25 in fact, they show longer life. But again, the data NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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36 1 is very limited. There's few data sets at 400 which 2 actually show longer life for alloy 600. But again, 3 at present, since all curves are based on room 4 temperature data, we are not taking any temperature 5 dependence for air. But for water effects, 6 temperature is important and explicitly defined in the 7 expressions to calculate fatigue life in water.

8 DR. WALLIS: That means it is through the 9 median of the data in some way?

10 MR. CHOPRA: I'll show you how we got the 11 best fit curves.

12 DR. WALLIS: It's supposed to be an 13 average right through the middle of the data.

14 MR. CHOPRA: Right.

15 DR. WALLIS: It's not best fit to a 95 16 percentile or something like that? You'll get to that 17 too, but what you're showing here is --

18 MR. CHOPRA: Average, right. These 19 results show nickel alloy data for alloy 600 and some 20 of the welds. In BWR, normal water chemistry, BWR 21 environment and PWR environment and again, what we see 22 is the effects are similar to what we get for 23 austenitic stainless steels. There's larger effect in 24 low oxygen than in high oxygen. PWR environment has 25 larger effect than BWR, but the focal effect is much NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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37 1 less than what you would see for austenitic stainless 2 steel.

3 Typically, under certain conditions in 4 austenitic stainless steel we see a reduction of a 5 factor of 14 or 15. In this, the maximum is a factor 6 of 3. So the effect is much less, but we can use this 7 limited data to define the important parameters and 8 how to estimate environmental effects.

9 Now we have all this data. How do we 10 generate the expressions? All -- in air, all data, 11 fatigue data I expressed by this modified Langer 12 equation where fatigue life is expressed in terms of 13 strain amplitude and these constants A, B, C --

14 DR. WALLIS: Is this an equation because 15 you plotted the data on log paper, is that why it is?

16 MR. CHOPRA: This is the expression used 17 and it presents the data best.

18 DR. WALLIS: It's because you plotted it 19 on log paper. It looks good on log paper and it's 20 linear.

21 MR. CHOPRA: Well, the trend is also -- it 22 does represent the trend.

23 DR. WALLIS: Okay.

24 MR. CHOPRA: And C is the fatigue limit or 25 related with the fatigue limit of the material. B is NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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38 1 the slope of that curve. A is a constant which would 2 vary with heat to heat. Depending on a more resistant 3 material would give a higher A or lower means it's 4 less resistant to fatigue damage.

5 We can do a best fit of the data and also 6 use this A to represent heat to heat variability and 7 come up with a median value, how median material would 8 behave. Best fit gives me the average behavior, 9 whereas a distribution would give me how various 10 materials behave and I get a median curve and then 11 come up with a number which would bound 95 percent of 12 the materials. And that's what I'm going to show.

13 One more thing, another term, D can be 14 added to impute in 1, which would include parameters 15 like temperature, strain rate and so on.

16 DR. WALLIS: Does the ASME curve have a 17 similar equation?

18 MR. CHOPRA: Yes. The Langer equation is 19 very -- yes.

20 This shows for low-alloy steels in air and 21 water various heats. Now each did define even if I 22 have 10 data points, it's 1 point. Another may have 23 500 data points. But if it's the same material, it's 24 just one point on this plot. This way, I can give 25 you, we can determine the median value for the NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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39 1 materials and if I select a fifth percentile number, 2 in this case, 5.56, if I select the A or 5.56, that 3 curve would bound 95 percent of the --

4 DR. WALLIS: It's the coefficient.

5 MR. CHOPRA: So this is how we obtain the 6 design curve by defining what subfactors I need to 7 adjust the best fit curve for average curve to come up 8 with a design curve which would bound 95 percent of 9 the materials.

10 I'll give the loca probability of track 11 initiation.

12 MR. BANERJEE: There's B and C as well, 13 right?

14 MR. CHOPRA: B and C, what I do is use it 15 for normalizing to get A for each heat which is the 16 average heat and I get a standard deviation. That's 17 what I've plotted here.. For the particular heat, I've 18 given the average value and the standard deviation for 19 the data set.

20 MR. BANERJEE: You lost me.

21 CHAIRMAN ARMIJO: B and C are relatively 22 constant.

23 MR. CHOPRA: A is the one that changes.

24 MR. BANERJEE: So you fix B and C to some 25 value?

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40 1 MR. CHOPRA: Right, right. And we know 2 even environment does not change. The strain 3 threshold was close to fatigue limit so I don't have 4 to change the fatigue limit. And there is no data 5 which suggests that C changes, means that the fatigue 6 limit changes for material.

7 DR. WALLIS: The range of that is not very 8 big, but if N is E to the A, so it's a factor of about 9 10 on the whole range.

10 MR. CHOPRA: Right.

11 MR. BANERJEE: Do B and C govern the shape 12 of the curve?

13 MR. CHOPRA: Yes. Right. The slope is B.

14 C is where at 106 or 107.

15 DR. WALLIS: I see where it's flat.

16 CHAIRMAN ARMIJO: So all the environmental 17 effects are just put into the A constant?

18 MR. CHOPRA: Right.

19 CHAIRMAN ARMIJO: Okay.

20 MR. CHOPRA: Now we come up with these 21 expressions which can be used for predicting fatigue 22 life under various conditions. Again, Langer equation 23 A, constant A; slope B and C. And this is the 24 environmental term B which would have these -- which 25 would depend on these three parameters for carbon low-NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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41 1 alloy steel, same for content, given by these 2 expressions, temperature, dissolved oxygen and strain 3 rate.

4 CHAIRMAN ARMIJO: Now the A is the five 5 percent number?

6 MR. CHOPRA: No. These are still the 7 average numbers.

8 CHAIRMAN. ARMIJO: These are average 9 numbers.

10 MR. CHOPRA: Next, I'll get to where we 11 apply those adjustment factors to get the design 12 growth.

13 DR. WALLIS: What does N mean here?

14 MR. CHOPRA: Cycles --

15 DR. WALLIS: Environment. N for 16 environment, is that PWR?

17 MR. CHOPRA: No, this is in error what the 18 expression is. This is in the light water reactor.

19 DR. WALLIS: Okay.

20 MR. CHOPRA: It doesn't matter whether 21 it's BWR or PWR because these are the parameters which 22 will change in various environments, reactor 23 environments.

24 MR. BANERJEE: Is there no effective 25 hydrogen on it at all?

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42 1 MR. CHOPRA: In BWR environment, there's 2 'about 2 ppm dissolved hydrogen, but I think it's the 3 hydrogen which is created by the austenitic reaction 4 which is more important than what is -- it does 5 control ECP, the electrical potential of the 6 environment. So hydrogen would change the ECP, but 7 below -250 electrical potential, effects are not that 8 much different. But you know, in crack growth rates 9 there is some effect, depending on -- well, in this 10 case all -- we use only 2 PPM hydrogen.

11 MR. BANERJEE: These are all done in 12 autoclaves or whatever?

13 MR. CHOPRA: And we do simulate these 14 conditions. BWR, it's high oxygen, high purity, very 15 high purity. And pressurized water reactor, again 16 high purity. Then we had boron or boric acid to get 17 boron, 1,000 PPM and 2 PPM lithium, by adding lithium 18 hydroxide. And measure the pH. We measure the 19 conductivity and maintain all these water chemistry 20 parameters constant during the test.

21 CHAIRMAN ARMIJO: These are flowing a loop 22 type --

23 MR. CHOPRA: Very small flow rates. I 24 think if you look at the -- my plot, they would amount 25 to 10-5 meter per second. Very low.

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.43 1 CHAIRMAN~ ARMIJO: They're not static 2 autoclaves?

3 MR. CHOPRA: They're not static and they 4 are continuously reconditioned. So if they are, it's 5 once through. They're not repeated.

6 DR. WALLIS: How long are the tests done 7 typically?

8 MR. CHOPRA: Depends on the conditions.

9 At low strain amplitudes and low strain rates, it may 10 take up to 5 to 8 months. and those results are very 11 limited. In the range which people have -- we have 12 tested .25 to .4 strain amplifies, it can take 13 anywhere from a few days to a month or two, depending 14 on the environmental effects. In air, they're much 15 longer. So one has to consider all of these. We 16 can't just dedicate and that's why you see very low, 17 less data under conditions which have very long 18 durations.

19 Now I just want to mention that these 20 expressions are average behavior after median 21 material. Same thing for rod and gas stainless steel.

22 Now as you mentioned that the slop e of the 360 NG was 23 different, what we have done is we have used a single 24 expression to represent all grades of steel and this 25 number, the fatigue limit we chose what studies in NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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44 1 Japan have established. And Jaske and O'Donnell in 2 1978 pointed this out that the current design curve 3 for stainless steel was not consistent with the 4 experimental data.

5 DR. WALLIS: I want to check this about 6 oxygen. You say it's worse to have less oxygen?

7 MR. CHOPRA: Pardon me?

8 DR. WALLIS: N goes down when you have 9 less oxygen?

10 MR. CHOPRA: In stainless steel, life goes 11 down dissolved oxygen is low.

12 DR. WALLIS: But these it goes the other 13 way?

14 MR. CHOPRA: No. The oxygen, there's a 15 constant factor --

16 DR. WALLIS: In the one before, the carbon 17 and low-alloy steels?

18 MR. CHOPRA: Yes. Now in carbon and low-19 alloy steel it's the high oxygen which is more 20 damaging.

21 DR. WALLIS: Then it doesn't make -- okay, 22 okay. That's right. Okay. Because I thought it was 23 the other way around. That's a negative --

24 MR. CHOPRA: The strain rate term is a 25 negative.

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45 1 DR. WALLIS: That's right. I was crawling 2 through that and then I was trying to go back to 3 before.

4 MR. CHOPRA: Actually, this whole term is 5

6 DR. WALLIS: I understand that. Just 7 before, but the other with the stainless steel, the 8 low oxygen is bad.

9 MR. CHOPRA: Right.

10 DR. WALLIS: Okay, that's what I'm trying 11 to --

12 MR. CHOPRA: I just mentioned that we 13 established a single curve and this we selected from 14 what was proposed by these studies.

15 Now we have the specimen data. We know 16 how to predict what will happen with specimens.

17 DR. WALLIS: What effect does this have on 18 welds of dissimilar metals?

19 MR. CHOPRA: Welds have different --

20 DR. WALLIS: All together different?

21 MR. CHOPRA: Yes.

22 DR. WALLIS: Is there some basis for that?

23 MR. CHOPRA: It depends on the data.

24 DR. WALLIS: You're not addressing that?

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46 1 design curves for these grades or types of structural 2 steel.

3 CHAIRMAN ARMIJO: For example, a welded 4 stainless steel is like a cast stainless steel, a weld 5

6 MR. CHOPRA: I think the behavior is very 7 similar. But --

8 CHAIRMAN ARMIJO: If it's similar, there's "9 a difference.

10 MR. CHOPRA: Because in some cases there 11 may be difference. We are just looking at here the 12 rod products.

13 CHAIRMAN ARMIJO: Stainless.

14 DR. WALLIS: Is there any effect of 15 fluence on this?

16 MR. CHOPRA: Irradiation? I'm sorry I, I 17 didn't get that?

18 DR. WALLIS: Is there any effect of 19 fluence?

20 MR. CHOPRA: We're not studying that.

21 There is an effect, but that's not -- in the design 22 curve --

23 DR. WALLIS: It's all synergistic.

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47 1 which are not considered in their design.

2 We have the data for specimens. Now to 3 use it to come up with a design curve for components, 4 I mention that they apply this adjustment factor of 20 5 on life and this factor is made up of effects of 6 material availability, data scatter, size, surface 7 finish, loading history.

8 In the current code, these are the 9 subfactors which are defined in the basis document.

10 Loading history was not considered, a total of 20 11 adjustment factors. In our study, based on the 12 distribution I showed for individual materials, this 13 subfactor can vary anywhere from a minimum of 2.1 to 14 2.8. These numbers are taken from studies in the 15 literature. Size can have an effect, minimum 1.2, 1.4 16 and so on. So we see a minimum of 6, maximum of 27.

17 When we take a large number, for example, 20, what we 18 are basically saying is I have a very bad material 19 which is very poor in fatigue resistance. I have 20 rough surfaces and I have the worse loading history.

21 So we used a Monte Carlo simulation and 22 using these as a log normal distribution to simulate 23 what would be the best adjustment needed to define the 24 behavior of components.

25 CHAIRMAN ARMIJO: So the present study, NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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48 1 you've agglomerated the date for carbon steels and 2 austenitic stainless steels and all these factors are 3 all pushed together.

4 MR. CHOPRA: Right.

5 CHAIRMAN ARMIJO: But you've separated 6 them. Are they different?

7 MR. CHOPRA: No, these are not the effects 8 of materialability is here and that depends on the 9 material. But effects of surface finish of the 10 component, size of the component or loading history 11 means random loading, high stress cycle followed by 12 low stress cycles. These -- in the current data, 13 these effects are not included. So somehow I need to 14 include these effects to come up with a design curve 15 which would be applicable to a real actual reactor 16 component.

17 Now the question is 20 was selected with 18 some basis. Is this reasonable because quite often, 19 this is. what is being questioned. There may be 20 conservatism in this which we need to eliminate. So 21 we are trying to see what possible conservatism might 22 be there in this margin or the adjustment factor of 23 20.

24 DR. BONACA: Twenty was arbitrarily taken 25 as a bounding number, right?

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49 1 Where did you get the 27?

2 MR. CHOPRA: I just took from the 3 literature what people have observed, effect of 4 surface-- surface finish is very well documented.

5 Depending on the average surface finish, an autonomous 6 value of surface finish, they have a harmless 7 reduction in light. So I can use typical finish for 8 grinding or milling operation and so on. It's well 9 documented. We can come up with what would be a 10 typical fabrication process, minimum and maximum. So 11 that's how we came up with this number.

12 DR. WALLIS: What is the basis of the 13 numbers? Is it trying to bound the data or bound the 14 95th percentile?

15 MR. CHOPRA: To come up with a design 16 curve which will be applicable to components.

17 DR. WALLIS: What's the basis of this? Is 18 there a rationale?

19 MR. CHOPRA: Right, 95 percent.

20 DR. WALLIS: Ninety-five, 99, 95?

21 MR. CHOPRA: Ninety-five?

22 DR. WALLIS: Why is 95 good enough?

23 MR. CHOPRA: Well --

24 DR. WALLIS: Why not 99?

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50 1 analysis to see what are the probabilities.

2 CHAIRMAN ARMIJO: I think 95/5 basis is 3 sort of a typical basis we've used in a lot of other 4 studies on failure data. But the reason that 95/5 is 5 okay is we've already done risk studies with fatigue 6 cracks initiating and growing to failure and growing 7 to leakage and the fact of a 95/5 probability of 8 fatigue crack initiation still keeps you in acceptably 9 low probability of getting a failure.

10 DR. WALLIS: Okay, so it's related to the 11 overall --

12 CHAIRMAN ARMIJO: Overall margin, yes. If 13 it were just a 95/5 to failure it would be an 14 unacceptable criteria.

15 DR. WALLIS: If the consequence were much 16 worse, you'd need to have a --

17 CHAIRMAN ARMIJO: Yes.

18 MR. BANERJEE: Can you expand a bit more 19 by what you mean by this log normal distribution?

20 MR. CHOPRA: We assumed that the effects 21 of all of these parameters have a log normal.

22 MR. BANERJEE: Of some mean?

23 MR. CHOPRA: Right. And I took these two 24 ranges as the 5th and 95th percentile of that 25 distribution.

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51 1 MR. BANERJEE: So what happens if you 2 chose a different distribution? Does it make any 3' difference to the results?

4 MR. CHOPRA: We have tried three 5 different, I think Bill tried and this gets the best 6 --

7 MR. BANERJEE: Best in what sense?

8 MR. CHOPRA: Very consistent result.

9 There's not much difference between normal and log 10 normal was not much difference. And log normal -- you 11 want to --

12 DR. SHACK: It's basically sort of an 13 arbitrary engineering judgment question. Experience 14 has indicated that when we have enough data, these 15 things do seem to be distributed log normally.

16 We generally don't have enough data, 17 actually, to determine the distribution. So we have 18 sort of just made the engineering judgment that the 19 log normal is close enough.

20 As John was explaining --

21 MR. BANERJEE: It doesn't affect the 22 results.

23 DR. SHACK: It doesn't affect the results 24 very much. What we're trying to do is to bound the 25 data in some reasonable fashion because the NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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52 1 consequence is not core damage when we're done. The 2 fact that we're not highly precise on this is not 3 something that concerns us, but we. think we've built 4 in sufficient conservatism to account for these 5 variables in a sensible way without going overboard.

6 And the fact that these affects can be 7 considered as independent is also something we don't 8 have data on. We have to sort of work on an 9 engineering judgment basis. So the Monte Carlo 10 simulation that we do assumes the log normal 11 distribution, assumes the independence.

12 MR. CHOPRA: I want to add one more, quite 13 6ften, actually in the welding research that WRC 14 Bulletin by industry, they are suggesting that in this 15 margin of 20, we can use a factor of 3 to offset 16 environment. This kind of analysis can suggest or 17 show that 3 number is very high. We do not have that, 18 at least what is the possible --

19 DR. KRESS: Is it a theoretical basis for 20 assuming the log normal? There may be, you know. You 21 can look at the physical phenomena and --

22 DR. SHACK: Well, the loading, probably --

23 DR. KRESS: Loading you would think would 24 be log normal. I'm not sure about the effects of the 25 other things.

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53 1 DR. SHACK: The log normal turns out to be 2 slightly more conservative than the normal and so 3 those were my -- if I don't have enough data to define 4 a distribution -

5 DR. KRESS: You might as well use --

6 DR. SHACK: I pick one or the other, sort 7 of on some sort of engineering judgment. The 8 differences are not very large between the two and we 9 just pick the log normal.

10 DR. WALLIS: If you know the distribution, 11 why do you need -- if you know 'the equation for the 12 distribution, why do you have *to do a Monte Carlo 13 analysis?

14 DR. SHACK: Because I'm taking a bunch of 15 random variables.

16 - DR. KRESS: That's the way you find the 17 mean, right?

18 MR. CHOPPA: There are four or five of 19 these things.

20 DR. SHACK: There are four or five 21 distributed variables.

22 DR. WALLIS: Easier to do it than to try 23 to go through the mathematics of predicting.

24 DR. SHACK: Yes, it's easier. Yes, I 25 could do it the other way, right.

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54 1 DR. KRESS: Is the 95 value four times the

2. mean?

3 DR. SHACK: No.

4 DR. KRESS: It has to be if it's log 5 normal.

6 DR. WALLIS: Four times the mean on a 7 constant A would be horrendous.

8 DR. KRESS: You've got to find the mean 9 value.

10 DR. WALLIS: Mean value is about five.

11 CHAIRMAN ARMIJO: Let's move on.

12 MR. CHOPRA: Doing this simulation, we get 13 these curves where this dash curve is now for the 14 specimen, the distribution of A for the specimen and 15 solid would be the distribution for the real 16 component. And we see that the median value has 17 shifted by about 5.3.

18 And 95 of 5th percentile is a factor of 19 12. So we can say that in this factor of 20, there is 20 some conservatism and we can use adjustment factor of 21 12 on life instead of 20.

22 DR. WALLIS: Where did 20 come from?

23 MR. CHOPRA: It's in the design basis 24 document of the current code.

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55 1 wise men?

2 CHAIRMAN ARMIJO: Many years ago.

3 MR. CHOPRA: Basically, that's what it 4 was.

5 MR. BANERJEE: Not so bad.

6 MR. CHOPRA: The design has several --

7 yes.

8 I've covered -- there is some conservatism in the 9 fatigue evaluations and often this conservatism is 10 used to offset environmental effects and there are two 11 sources of conservatism, in the procedures themselves, 12 the way we define design stresses and design cycles or 13 this adjustment factors of 2 and 20.

14 I showed there's not much margin, only 1.7 15 in this factor of 20, but the current code procedures 16 --

17 DR. WALLIS: Is there enough to account 18 for environmental effects?

19 MR. CHOPRA: No, environmental effects can 20 be as high as a factor of 15.

21 DR. WALLIS: Yes.

22 MR. CHOPRA: Or carbon C would be even 23 higher.

24 DR. WALLIS: These are all reactor data 25 you've got, right?

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56 1 MR. CHOPRA: Those are -- unless you 2 define the operating transient conditions. In certain 3 conditions those may be possible, but again, it's up 4 to the designer to define what are the conditions 5 during a transient, mean strain rates, temperatures 6 and so forth.

7 MR. BANERJEE: But I'm wondering whether 8 in your database you have anything which you've 9 evaluated from N reactor data or reactor data. Do you 10 have any information at all?

11 MR. CHOPRA: There are some components and 12 so on and I list a few examples where there have been 13 some studies. And I'll show you near the end of this.

14 DR. SHACK: The trouble with doing this 15 with field data is it's hard to control variables like 16 knowing that the strain range and because that has 17 such a strong effect on it. Unless you know that 18 accurate, it's hard to back out the result.

19 MR. CULLEN: Bill Cullen, Office of 20 Research. I'd like to explore Dr. Banerjee's question 21 a little more to find out what's behind it.

22 Are you concerned about irradiation 23 effects which really do not come into play for 24 pressure boundary? Or are you concerned about the 25 actual aqueous environment and its characteristics?

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57 1 I'm not sure -- what is the basis?

2 MR. BANERJEE: Well, the basis is more --

3 it would be nice to see some validation under field 4 conditions. There are always sort of surprises 5 between the lab and what happens in the field and even 6 if this sort of validation is not all that thorough, 7 a couple of data points would set your mind at rest 8 that it's not some unexpected factor that comes in.

9 It's more like -- I have a concern always 10 of going from the lab to a real field situation. It's 11 not for any specific issue, not like radiation or 12 combination of factors or boron plus temperature in 13 fatigue cycles which are slow. All these things may 14 or may not be there but just a general question, more 15 a general question.

16 MR. CULLEN: I understand the general 17 question. I'm a little concerned about your word 18 about there always are surprises when you go from the 19 laboratory to the actuality.

20 MR. CHOPRA: Maybe that's too strong.

21 MR. CULLEN: A little bit.

22 (Laughter.)

23 DR. WALLIS: Oftentimes, surprises may be 24 small.

25 MR. CULLEN: Thank you.

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58 1 MR. BANERJEE: I don't mean to say that 2 this stuff should not be used or anything. Right.

3 MR. CHOPPA: I mentioned that in fatigue 4 evaluations the procedures are quite conservative, but 5 the code allows us to use improved approaches, for 6 example, finite element analysis, fatigue monitoring 7 to define the design stresses and cycles more 8 accurately. So most of this conservatism can be 9 removed with better methods for defining these design 10 conditions.

11 So in that case, there is a need to 12 address the effect of environment explicitly in these 13 procedures.

14 Now the two approaches which we can use 15 either come up with new set of design curves or use 16 some kind of correction factor, Fen. Now since 17 environmental effects depend on a whole lot of 18 parameters, temperature, strain rate and so on, either 19 we come up with several sets of design curves to cover 20 the possible conditions which occur in the reactor or 21 field conditions or if you use a bounding curve, it 22 would be very conservative for most of the conditions.

23 Whereas this correction factor, F en 24 approach is relatively simple. You can -- it's very 25 flexible. You can calculate the environmental effects NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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59 1 for a specific condition. And this is what is being 2 proposed in this reg. guide.

3 The correction factor is nothing, and this 4 was proposed in 1991 by the Japanese. A correction 5 factor is nothing but a ratio of fatigue life and air 6 versus life and water. So we have these expressions 7 I showed you in the previous slides and we can then 8 calculate Fen for different steels, carbon steel, low-9 alloy steel, and below a strain threshold there's no 10 environmental effects, so the correction factor would 11 be one.

12 Other than that, we use these expressions, 13 actual conditions, temperature, strain rates and so on 14 to calculate the correction factor. To incorporate 15 environmental effects, we take the usage, partial 16 usage factors obtain for specific transients in air, 17 U1, U2 and so on, multiplied by the corresponding 18 correction factor and we get usage factor in the 19 environment.

20 Now to calculate usage factors in air, we 21 should use design curves which are consistent with or 22 conservative with respect to the existing data. And 23 as has been pointed out.quite a few years back, the 24 current code curve. for stainless steel is not 25 consistent with the current existing data and should NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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60 1 not be used for obtaining usage. And I just want to 2 show before I get to that, these are the expressions 3 for nickel allows. Correction factor, again, as a 4 function of these three variables. And usage and air 5 would be obtained from the curve for austenitic 6 stainless steels.

7 Now I mentioned that the current design 8 curve for austenitic stainless steel is not consistent 9 with the data. I plotted the fatigue data for 316, 10 304 stainless in air, different temperatures and this 11 dashed curve is the curve, current code mean curve.

12 This is the mean curve which was used to obtain the 13 design curve.

14 DR. WALLIS: Where is your design curve?

15 MR. CHOPRA: Design curve would be what 16 you adjust this curve for mean curve correction.

17 DR. WALLIS: Your recommended curve would 18 actually bound the data, wouldn't it?

19 MR. CHOPRA: This is the best -- actually, 20 this data, the curve is based on austenitic stainless 21 steel.

22 DR. WALLIS: I thought you were 23 recommending a bounding curve with this factor.

24 MR. CHOPRA: I'm just trying to show that 25 the current --

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61 1 DR. WALLIS: What's your design curve?

2 You should show that, shouldn't you?

3 MR. CHOPRA: These are mean curves.

4 DR. SHACK: This is air data, mean curve.

5 If we put a design curve on here, we could have a 6 design curve in air and a design curve in --

7 DR. WALLIS: There's all this air data.

8 Are you going to get to your -- it's so far down the 9 road, I can't -- okay.

10 CHAIRMAN'ARMIJO: I think he's just trying 11 to show the difference between the two sets of means.

12 MR. CHOPRA: That the current means --

13 DR. WALLIS: You do show the effect of the 14 F factors yet.

15 MR. CHOPRA: No. I'm just trying to show 16 --

17 DR. WALLIS: We've just been talking about 18 19 DR. SHACK: What he's trying to 20 demonstrate here is that the F factor requires him to 21 take the ratio in air. He's got to have the right air 22 curve.

23 MR. CHOPRA: And the current mean curve 24 for air, for austenitic stainless steel, is not 25 consistent with the data.

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62 1 Now I'd like to mention one thing, it's 2 been suggested that this curve, the data may be 3 different from the mean curve because of the way 4 fatigue life has been defined or the way we conduct 5 experiments. I can assure you that this difference in 6 the mean curve and the data is not due to any artifact 7 of test procedures or the way the fatigue life is 8 defined in terms of failure or 25 percent load drop.

9 DR. WALLIS: What occurs to me is the ASME 10 code mean curve was a mean curve to something.

11 MR. CHOPRA: Right.

12 DR. WALLIS: And it was presumably through 13 other data.

14 MR. CHOPRA: This curve, the current code 15 curve was based on very limited data. Now we have 16 much more. So I'm just showing that the data which 17 has been obtained since then is not consistent with 18 what we have.

19 DR. WALLIS: You have a much broader data 20 base.

21 MR. CHOPRA: Right.

22 DR. WALLIS: Okay, that's why yours is 23 better?

24 (Laughter.)

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63 1 change the current code curve. The current code curve 2 is not consistent with --

3 DR. WALLIS: It must have been based on 4 something.

5 MR. CHOPRA: And that data is somewhere in 6 here, up here. But since then we have much more data.

7 DR. WALLIS: Either that or steels have 8 been getting weaker.

9 MR. CHOPRA: Actually, that is the reason.

10 Mostly like because of the strength of the steel, 11 probably these curves were obtained on steel which was 12 stronger.

13 DR. WALLIS: Wait a minute --

14 MR. CHOPRA: Possible difference.

15 MR. CULLEN: Bill Cullen, Office of 16 Research again. Omesh, if you could go back to that, 17 I'd like to also point out that the curves on which 18 the original ASME code were based I think the data 19 only went out- to a factor of about, fatigue life of 20 106 or something.

21 MR. CHOPRA: Not even 6.

22 MR. CULLEN: So you've got two orders of 23 magnitude extrapolation there that we're doing now to 24 illustrate. But the other thing again is those tests 25 were all done at room temperature and you're showing NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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64 1 data from a wide variety of temperatures up to and 2 including operational.

3 MR. CHOPRA: Stainless does not --

4 MR. CULLEN: Doesn't show much difference, 5 right. To me, that's kind of the point. It all hangs 6 together on the lower curve.

7 MR. CHOPRA: This difference is genuine.

8 We need to use a different curve. And we have now 9 proposed a design curve for air for austenitic 10 stainless steels, the solid line. The current dashed 11 line is the current code of 10 6 and the high cycle 12 extension in the code. And the solid line curve is 13 based on the Argonne model plus adjustment factors of 14 12 on life and 2 on stress. It's not 20 and 2. It's 15 12 and 2.

16 DR. WALLIS: Now the kink that you have 17 here at 106 doesn't appear in the previous curve you 18 showed.

19 MR. CHOPRA: The design curve extends only 20 up to 106.

21 DR. WALLIS: So you've just extrapolated 22 it here in your figure?

23 MR. CHOPRA: Yes, because now there is a 24 need to go all the way to 1011.

25 DR. WALLIS: But you're saying mean curve, NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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65 1 so where. do you stop at 106?

2 CHAIRMAN ARMIJO: Two different things 3 here, hold on.

4 MR. FERRER: This is John Ferrer. I think 5 originally the stainless steel curve went out to 106.

6 Later, they got more data at high cycles and the data 7 was clearly showing that there was a drop off and so 8 they -- this is an artifact of fairing the two curves 9 together and the new correction we're doing really is 10 straightening out what they should have straightened 11 out to begin with.

12 DR. WALLIS: Well, it's a curve, it can't 13 be straightened out.

14 (Laughter.)

15 MR. FERRER: Fur the earlier slide was the 16 man curve through the data. Now we are talking about 17 the code curve which would include these factors.

18 DR. WALLIS: Okay.

19 MR. GURDAL: There is still a curve A, B 20 and C.

21 My name is Robert Gurdal. I'm AREVA, 22 Lynchburg, Virginia. Those curves is because before 23 just now there are three curves, there is A, B and C 24 and they are not indicated there. I just wanted to be 25 sure everybody knows.

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66 1 The reason you have the lower one which is 2 called a curve C --

3 MR. CHOPRA: But the region which we are 4 talking about is this 106 to 10 --

5 MR. GURDAL: You go above 16, you have a 6 curve A, curve B and curve C.

7 MR. CHOPRA: I have plotted that.

8 MR. GURDAL: The correct curve is curve A 9 which is the top one.

10 DR. WALLIS: So it's C on this figure and 11 it's A on the previous figure.

12 MR. GURDAL: Maybe, it could be.

13 DR. WALLIS: Maybe. It probably doesn't 14 matter that much.

15 MR. GURDAL: And the C is for the heat 16 affected zone compared to the A.

17 DR. WALLIS: This is the A in this one.

18 MR. GURDAL: That one could be the A, 19 because it does not have the kink.

20 MR. CHOPRA: This is the mean curve.

21 MR. GURDAL: Oh, that's the mean curve.

22 Sorry about that. But the design curve, if you go to 23 the design, there is a curve continuing without any 24 disconnection.

25 DR. WALLIS: Without any king, yes. Okay.

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67 1 MR. GURDAL: And that's the A. This one 2 is a C.

3 MR. CHOPRA: But the region we are talking 4 about is this.

5 MR. GURDAL: Okay, but the question was 6 about 106.

7 MR. CHOPRA: Which needs to be corrected.

8 DR. WALLIS: Okay, we've resolved that, I 9 think. Thank you. That's very good.

10 CHAIRMAN ARMIJO: Which gets to the point, 11 your design curve treats the weld heat affected zones 12 or the base material, everything as the same as 13 opposed to the code.

14 MR. CHOPRA: Yes, I think so.

15 MR. FERRER: I think so. In the code, I 16 think the previous gentleman was talking about their 17 -- in the high cycle regime, there are three separate 18 curves proposed by ASME that extend past the 106 19 cycles.

20 In our proposal we've just bounded that 21 with one curve.

22 MR. CHOPRA: We also have generated design 23 curves for carbon and low-alloy steels based on the 24 same approach using the Argonne models and adjustment 25 factors of 12 and 2. This is for carbon steel and NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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68 1 next is for low alloy.

2 Now current code curve for these is only 3 10 6and now this is the -current code curve and an.

4 extension has been proposed by a subgroup, fatigue 5 strength. This was proposed a few years back and it's 6 still not approved by the ASME code committees. We 7 are -- we have another approach to define extension of 8 this curve beyond 106 cycle. I just wanted to give a 9 couple of slides to show that.

10 What the subgroup fatigue strength 11 proposed was extension of the curve which is based on 12 load control data and the data extends only up to 106 13 and they use maximum effect of mean stress and they 14 propose extension which is expressed by applied stress 15 amplitude given in terms of life with an exponent of 16 - .05 which means 5 percent decrease in life, in stress 17 every decade. And since the data only extends up to 18 5 times 10 6, extrapolation to 10 may give 19 conservative estimates.

20 Another way of extending this curve would 21 be to use the approach with Manjoine had proposed a 22 few years back where the high-cycle fatigue is 23 represented by elastic strain with life blots and if 24 we use existing data which we have extending up to 10ll 25 cycles for these various speeds, we get a slope of-NEAL R.GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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69 1 007. Manjoine proposed -. 01 and we can use this 2 expression where the exponent is smaller and which is 3 consistent with the data and this would be for the 4 mean curve.

5 Now we take this adjusted for mean stress 6 correction using Goodman relation which is a 7 conservative approach and actually if we do that this 8 exponent would be .017. So it's slightly lower than 9 what is being proposed by the subgroup fatigue 10 strength, but we can use this expression and that's 11 *what we have used to define that extension to the 12 curve.

13 DR. WALLIS: When you make these 14 proposals, did you negotiate something with ASME or 15 did you just say this is what we use --

16 MR. CHOPRA: This has been presented to 17 them.

18 DR. WALLIS: There wasn't any give and 19 take. It was just -- you deduced this from your data?

20 MR. CHOPRA: I attended the subgroup 21 fatigue strength and all our work has been presented 22 there.

23 DR. WALLIS: But the proposal is 24 essentially yours. It isn't some compromise proposal.

25 It's your proposal.

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70 1 MR. CHOPRA: This was proposed by Manjoine 2 a few years back, so this is nothing new.

3 DR. WALLIS: All these green curves are 4 Argonne curves, proposed by Argonne?

5 MR. CHOPRA: No, the best fit curves are 6 what we have defined.

7 DR. WALLIS: Right, so they're not 8 something which has been negotiated and agreed on or 9 anything like that?

10 CHAIRMAN ARMIJO: It's certainly been 11 discussed.

12 DR. WALLIS: It's been discussed. IT's 13 been presented. ASME hasn't come around and said yes, 14 you guys are right.

15 DR. SHACK: One thing to think about for 16 the carbon and low-alloy steels, there's really in air 17 there's no disagreement over the mean curve. The 18 shape may shift just a smidgen, but the only real 19 difference between this design curve and the current 20 is they use a factor of 12 instead of 20. Then you do 21 have the discussion over how to extend it.

22 The environmental effect is a --

23 DR. WALLIS: It's the big one.

24 DR. SHACK: That's the big one.

25 CHAIRMAN ARMIJO: In the reg. guide, does NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

(202) 234-4433 WASHINGTON, D.C. 20005-3701 (202) 234-4433

71 1 this curve really extend out to 1011 or does it -- is 2 it truncated at 10 7 , since there seem to be a big 3 difference.

4 MR. CHOPRA: The proposal is up to 1011.

5 CHAIRMAN ARMIJO: Up to 16', but compared 6 to the ASME code for this particular steel, your curve 7 is nonconservative.

8 MR. CHOPRA: Well, this is --

9 CHAIRMAN ARMIJO: You predict a much 10 longer life.

11 MR. CHOPRA: This is based on the data we 12 have.

13 CHAIRMAN ARMIJO: Right, but nobody has 14 data out to 1011.

15 MR. CHOPRA: No.

I CHAIRMAN ARMIJO:

16 It's a less conservative 17 18 DR. WALLIS: You have a C. You have a 19 constant C or --

20 CHAIRMAN ARMIJO: Right.

21 DR. WALLIS: I'm surprised it isn't 22 completely flat to a green curve.

23 MR. CHOPRA: Made up of two. I mentione d 24 that extension is a different slope.

25 DR. WALLIS: Do they ever have ib cycles NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., NW.

(202) 234-4433 WASHINGTON, D.C. 20005-3701 (202) 234-4433

72 1 in a nuclear environment?

2 MR. FERRER: Vibration --

3 DR. WALLIS: Shaking things that shake.

4 MR. CHOPRA: So the method to apply the 5 correction would be to use for carbon low-alloy steel 6 you can use either the current code design curves or 7 the curves I've mentioned to reduce some conservatism.

8 As you see, it's -- they're based on 9 adjustment factors of 12, rather than 20.

10 For austenitic stainless steels and nickel 11 alloys, we use a new design curve for austenitic 12 stainless steels. And in the appendix to NUREG, there 13 are certain examples given to determine some of the 14 parameters.

15 For example, lab data shows quite often 16 people don't know how to calculate, how to define the 17 strain rates. Lab data shows average strain rate 18 always is a conservative approach.

19 And similarly, if we have a well-defined 20 linear transient temperature change, that can be 21 represented by average temperature and it could be 22 okay.

23 Now this one shows two more slides and 24 I'll be done. There was a question that lab data does 25 not represent the feed. There are certain reports NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

(202) 234-4433 WASHINGTON, D.C. 20005-3701 (202) 234-4433

73 1 where some operating reports where some operating 2 experience and component test results have been 3 published.

4 This is EPRI report, 1997, and gives a 5 complete chapter, a couple of them, giving examples of 6 corrosion fatigue effects on nuclear power plant 7 components.

8 Similarly, studies in Germany, MPA and 9 other places have shown the conditions which lead to 10 what they call strain-induced corrosion cracking.

11 This was demonstrated for BWR environments. And there 12 are examples, even these examples are component test 13 results. We support the lab data.

14 I want to just show the results of one 15 particular test, component test, recent tests, again, 16 sponsored by EPRI where they used tube u-bend tests 17 tested in PWR water at 240. And I'm just plotting the 18 results for a given strain amplitude what was the 19 fatigue life they measured.

20 In earth environment, these are the 21 triangles. So that serves as a baseline you would 22 expect in air. Then they tested in PWR water in two 23 conditions: a strain rate of .01 percent per second 24 and diamonds are .005 percent per second. And this 25 would give me for this strain amplitude a life in air NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

(202) 234-4433 WASHINGTON, D.C. 20005-3701 (202) 234-4433

74 1 of 12,500. This is about 36,000. This is 1700. And 2 you can determine for a component test what is the 3 environmental factor.

4 In this test, inert environment cracks 5 were on the OD. And they were biaxial conditions.

6 And the water, they were on the ID. And nearly 7 uniaxial. So since there was a conversion, there's a 8 question whether this number is accurate.

9 There's another way we can determine the 10 baseline life. They have a very well-defined strain 11 rate effect between these two. I applauded the 12 component test results with the lab data, exactly the 13 same slope and we know somewhere there's a threshold.

14 That would be the life in air. So I've got a number 15 8,000; 12,000. I use an average of 10. Gives me a 16 reduction of 5.8 for one strain rate; 2.8.

17 And the Fen we have presented, give you 18 5.5 and 3.6. Ii think these are very reasonable 19 comparisons from a real component test.

20 MR. BANERJEE: So the test was done 21 outside the reactor, right?

22 MR. CHOPRA: This is a component test, 23 where they took an actual u-bend tube and strained it.

24 So it's not a small specimen. They are testing a real 25 component -- it demonstrates that lab data is NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

(202) 234-4433 WASHINGTON, D.C. 20005-3701 (202) 234-4433

75 1 applicable to actual component test conditions.

2 CHAIRMAN ARMIJO: Did you compare any of 3 the other component tests that you referenced in the 4 previous slide with your data to see how your data 5 predicts?

6 MR. CHOPRA: Some of the earlier, no, we 7 have not.

8 MR. BANERJEE: Do you have any idea of the 9 -- is there anything which happened in a reactor where 10 you have the strain history or something for a period 11 of time?

12 MR. FERRER: I think the answer to that is 13 it's very difficult to have the exact data on the 14 strain history in an actual operating event. We've 15 tried to estimate it and the best you can do is 16 estimate it. I think Omesh presented some references.

17 I think the EPRI one which attributed some of the 18 cracking to environment, but you couldn't prove it 19 absolutely because you just don't have the exact 20 temperature measurements and the strain measurements 21 at the location of your cracks.

22 MR. BANERJEE: But you can estimate them, 23 right? Based on those estimates, what does it look 24 like?

25 MR. FERRER: If you go back to the NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

(202) 234-4433 WASHINGTON, D.C. 20005-3701 (202) 234-4433

76 1 reference EPRI report, you know, I think based on 2 their estimates they attribute some of it to 3 environmental, but I say those estimates are very 4 crude. They're not nearly as controlled as the lab 5 data and if you look at fatigue, the -- at the low 6 cycle end, the small change in stress gives you a 7 fairly large change in the number of cycles if you 8 look at the shape of the curve.

9 And so it's not that easy. There are some 10 estimates, but they're more judgmental than accurate 11 calculations.

12 MR. BANERJEE: But the evidence or 13 supports -- what you're saying --

14 MR. FERRER: Well, there's some evidence.

15 What you'll hear from -- probably from ASME is the 16 overall operating experience doesn't show that there's 17 a big problem there.

18, MR. BANERJEE: Okay.

19 CHAIRMAN ARMIJO: Okay. That's it?

20 MR. CHOPRA: Yes.

21 CHAIRMAN ARMIJO: Any other questions from 22 the Committee?

23 MR. GONZALEZ: I would like to go back to 24 the reg. guide to present a summary of the three 25 regulatory positions.

NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

(202) 234-4433 WASHINGTON, D.C. 20005-3701 (202) 234-4433

77 1 Regulation position 1, we are endorsing 2 that we will calculate fatigue using air with ASME 3 code analysis procedures plus use the ASME code air 4 curves for new ANL modern air curves. This is for 5 carbon and alloy steels only.

6 Then we will calculate the Fen using the 7 appendix A of the NUREG for carbon and alloy steels 8 and this will be applied to calculate the 9 environmental uses factor.

10 But we're given the option of using the 11 ASME curve or the new air curve from the ANL model.

12 Or austenitic stainless steel, we will calculate the 13 fatigue use factoring there with the ASME code 14 analysis procedure, plus the new ANL model air 15 stainless steel curve.

16 We'll use the -- also the Fen equation for 17 stainless steel and then calculate the environmental 18 usage factor.

19 For nickel chrome alloys, will be Alloy 20 600, 690. You will use again the ASME code analysis 21 procedure plus the new ANL model air stainless steel 22 curve. As the reason was it was explained before was 23 because of the new data.

24 And if the Fen specifically for nickel 25 alloys and calculate the usage factor -- the NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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78 1 environmental fatigue usage factor.

2 In summary, Reg. Guide 1.207 will endorse 3 the use of a new air curve for austenitic stainless 4 steels and also will endorse the Fen methodology. It 5 will give guidance on incorporating the environmental 6 correction factor, the fatigue design analysis and 7 this is described in Appendix A of the NUREG report 8 and also the NUREG report will describe in detail the 9 technical basis.

10 That's it. Any more questions?

11 CHAIRMAN ARMIJO: Okay, *ny questions?

12 We're scheduled for a break about now, but we're a 13 little bit ahead of schedule. I don't know if we can 14 reconvene in 15 minutes or do we have to wait until 15 3:35?

16 We'll just take a 15-minute break. Be 17 back at 3:25. Is that right? 3:25, thank you.

18 (Off the record.)

19 CHAIRMAN ARMIJO: Okay, we've got --

20 incredibly we're about five minutes ahead of schedule, 21 so that's good.

22 So Mr. Gonzalez, would you like to 23 continue?

24 MR. GONZALEZ: This is our second part, 25 second presentation. It's in the resolution to public NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

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NEC-JH_28

Title:

Advisory Committee on Reactor Safeguards 549th Meeting Docket Number: (n/a)

Location: Rockville, Maryland Date: Thursday, February 7th, 2008 Work Order No.: NRC-2007 Pages 1-346 NEAL R. GROSS AND CO., INC.

Court Reporters and Transcribers 1323 Rhode Island Avenue, N.W.

Washington, D.C. 20005 (202) 234-4433

96 much done. The question I have: is there any question for the staff here? Any questions for Mr.

Hopenfeld from members?

MEMBER ARMIJO: I'd like to ask with respect to the last presenter's comments about new data. Is the staff familiar with .-- no, I'm asking the staff if they're aware of the new data that you referred to MR. FAIR: Hi. I'm John Fair with Division of Engineering who did a lot of the reviews on environmental fatigue.

Yes, we are. The new data is the latest Argonne data that was being applied in new design certifications. Basically the criteria they're using ins license renewal was criteria that was developed quite a while back, and we made a decision at that time that we would, as criteria, we would* maintain that criteria because there were a lot of applications in process. So we didn't want to keep changing the rules as these people were putting in new applications. And a lot of the criteria had changed and was massaged over the years.

Actually if you go back and look at the latest criteria we're applying to new reactors, it's not as conservative as the old criteria because we NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., N.W.

(202) 234-4433 WASHINGTON, D.C. 20005-3701 www.nealrgross.com

97 changed the basis for deriving the curves. So if you go and look at the Fen factors themselves using the new criteria, they'll generally be lower.

MEMBER ARMIJO: Okay. Thank you.

VICE CHAIRMAN BONACA: Any other questions?

If not, Mr. Chairman, I'll turn the meeting back to you.

CHAIRMAN SHACK: Okay. Well, it's five minutes late. I'd like to take a break now. I thank the presenters, staff and the industry, for a good presentation, I think, very informative and Mr.

Hopenfeld for his comments.

We're slated for 15 minutes. So we'll be back at ten of.

(Whereupon, the foregoing matter went off the record at 10:35 a.m. and went back on the record at 10:55 a.m.)

CHAIRMAN SHACK: The next topic is a draft final Revision 1 to Regulatory Guide 1.45, "Guidance on Monitoring and Responding to Reactor Coolant System Leakage, and, Sam, I think you're going to take us through that.

MEMBER ARMIJO: Right. Thank you, Mr.

Chairman.

NEAL R. GROSS COURT REPORTERS AND TRANSCRIBERS 1323 RHODE ISLAND AVE., NW.

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~ 'Het',and Mas's.T by E. R. G. ECKERT Professor of Mechanical Engineerinug::..,

and Director of the Heat Transfer Laor'tory u:,

JI University of Minnesota WITH PART A, HEAT CONDUCTION AND APPENDIXrTOF PROPERTY, VAUES.

K:

by ROBERT M., MRAKE,J.

Professor and Chairman.

Mechanical Engineering Department Princetoniý 4 4 II

,IINTRODUCTIONTOHERAS

. Second Edition.

4; 1 II N

iI*~ .4 J

L .McGRAW-HILL BOOK COMPANY, INC.

chard C New York Toronto London In sselae 1959 I' Ii Ip4 I

212 HEAT TRANSFER BY CONVECTION H. Hausen' gave the expression for the average Nusselt number:

Mud = 0.1 161(Red) - 1251(Pr)ý [II +( d15] 4 I

where pB is the viscosity at bulk liquid temperature and 'U the viscosity at tube-wall temperature. Apart from the latter, the property values are to be inserted at tB. This formula takes into account the conditions in the intake region. It also satisfactorily reproduces the values in the transition zone Red = 2,300 to 6,000. This relation is expected to be especially applicable to fluids for which the variation of viscosity is the 1N I

I I

I d

FiG. 8-9. Local Nusselt numbers for flow through a tube near the entrance witib simultaneous development of the flow and temperature field. [From TV. Linke and I

I

• H. Kunze, Allgem. Wdnnetech., 4:73-79 (1953).]

1 H. Hausen, Z. Vet. dent. Ingr., Beih. Verfahrenstech., no. 4, 1943, pp. 91-98.

I I

I I

Boundary Layer Theory BYE tt~

......................................................................

Dr HERMANN SCHLICHTING It" Professor at the Engineering University of Braunschweig -

Director of the Aerodynamische Versuchsanstalt Gdttingen Head of the Institute for Aerodynamics of the Deutsche Forschungsanstalt fuir Luft- und Raumfahrt, Braunschweig, Geinany

-.

Translated by I 14 I....): Dr J. KESTIN

. F . ""Professor

- at Brown University in Providence, Rhode Island.:

I~~2 Fourth Edition I ii McGRAW-HILL BOOK COMPANY, INC.

NEW YORK TORONTO LONDON VERLAG G. BRAUN - KARLSRUHE

• * "t!'.*

• " 'L*

• ' ;,?!.

I.

S

i

'.!.

c. The rough plate

,a commmr6n'mximnum. at hid ;z-- 0-5. Further small local maxima occur at,. .- h.d: i on"t*

1i-0,:The:.minimabetween them occur at -- hAd ; 0-2_ 0"8, and 135. Deperiding or_143,5 x<_Ioo ft qof th6 cavity it' may sometimes happen: that regular vortex patterns are fornmed

ýinAit 3 or 4-5 x"20ft' enee betwee toI,the different -values of drag. As seen from the symmetry of the curves about*Ihid ca*vities of up to.- d/h 0"1 give the same increase in drag as corresponding small protul rase In .,drag;¶f,'A I first. termj is th~e r ring .stress on',- 1, f the heih 'I

• rf&rthk "'Ihi fue was*v-arieid b theunnel.' ýFrcra bc diemensinle,.

idhby (21.34) oughnessele~ment, d over the hih Fig. 21.14. Resistance coefficient of cCircular cavities of.varying depth in a flat wll'ai asmeasiired by Wicghardt [54]

d The flow pattern wlhich exists behind an obstacle placed in the boundary layer neari*a lar ribs arraned-. liffers markedly from that behind an obstacle placed in the free stream. This circuimsl aid circular croiss ,e merges clearly from an experiment performed by. H. Schlichting [38] and Rillustr

'

wall and others., F'i. 21.15. The experiment consisted in the measurement of the velocity: field.behind

.ted in Fig. 21.3 pheres placed on a smooth flat surface. The pattern of curves of constant velocity' clearl[j*

kind of negative wake effect. The smallest velocities have been measured in the freelgal which no-spheres are present over the whole length of the plate; on the other hand.,he 1 Il v.oi elocities have been measured behind the rows of spheres where precisely the snailler _31j ki1d.

-7777r lines) 'as",

Fig. 21.15. Curves of constant velocity in the flow field behind a row of spheres (full:

rose in boundace thi~scase by H. Schlichting [38], and accompanying it the secondary flow (broken lines) ini de the measured the ofh t]

kratio :the boundary layer behind sphere (.1), as calculated by F. Schultz-Grunow [46]. In.theneigh-bourhood of the wall, the velocity behind the spheres is larger than that in the gaps. Theeipheres i-flow.

produce a "negative wake effect" which is explained by the existence of secondafy allcurves hav 4 mm' .

Diameter of spheres it j%."

- II Heat NEC-JH_31 "q, Transfer Fifth Edition J. P. Holman Professor of Mechanical Engineering Southern Methodist University McGraw-Hil Book Company New York St. Louis San Francisco Auckland BogotA Hamburg Johannesburg London Madrid Mexico Montreal New Delhi Panama Paris Sao Paulo Singapore Sydney Tokyo Toronto QC 3,2 EA p,

2 MW L L -LQ

226 Empirical and practical relations for forced-convection heat transfer Fig. 6-1 Total heat transfer in terms of bulk-temperature :r-difference. i K..

F low __

dx L

plicated problems may sometimes be solved analytically, but the solutione-U when possible, are very tedious. For design and engineering purposes_

empirical correlations are usually of greatest practical utility. In thi,7 section we present some of the more important and useful empirica:

relations and point out their limitations. .,

First let us give some further consideration to the bulk-temperatui concept which is :important in all heat-transfer problems involvingflei inside closed channels. In Chap. 5 we noted that the bulk temperatý7i .

represents energy average or "mixing cup" conditions. Thus, for the ti, flow depicted in Fig. 6-1 the total energy added can be expressed in terns a bulk-temperature difference by q = rhwP(Tb. Tbi Y~'(~

-

provided c, is reasonably constant over the length. In some differený: I length dx thedifference temperature heat added or indq terms can beofexpressed either incoefficient the heat-transfer terms of a bu1-dq = rhcdTb = h(2Tr) dx (T, - Tb) where T,, and Tb are the wall and bulk temperatures at the particuLii location. The total heat transfer can also be expressed as q = hA(T,, - Td)a,-

where A is the total surface area for heat transfer. Because both T Tb can vary along the length of the tube, a suitable averaging p ,

must be adopted for use with Eq. (6-3). In this chapter most G.__-

attention will be focused on methods for determining h, the conv heat-transfer coefficient. Chapter 10 will discuss different methoid, taking proper account of temperature variations in heat exchangersr For fully developed turbulent flow in smooth tubes the following tion is recommended by Dittus and Boelter [11:i Nud =0.023 Red8 Pr" The properties in this equation are evaluated at the fluid bulk temperat and the exponent n h'as the 'following'values: .

'7W 1,7

Film condensation inside horizontal tubes 413 may be restructured as 34 3

= C [p(p - pr)gk MP 4 sin (bAP.

jx2 4m L mk we may solve for h as 413 S=,C [(p- pcg 4..sin (9-23) fe now define a new dimensionless group, the condensationnumber Co, as p ) 3Co (9-24) l that Eq. (9-23) can be expressed in the form Co = C4 /3 (4*sin:0 IP1r3 Ref71t 3 (9-25)

Tor a vertical plate A/PL = 1.0, and we obtain, using the constant from Eq.

q.*-10),

Co = 1.47 Rep73 for Rer < 1800 (9-26) 44or a horizontal cylinder A/PL = Ir and Co = 1.514 Re-u13 for Ref < 1800 (9-27)

When turbulence is encountered in the film, an empirical correlation by TKirkbride (2] may be used:

Co = 0.0077 Re.'f for Ref > 1800 (9-28) 9-4 Film condensation inside horizontal tubes Our discussion of film condensation so far has been limited to exterior surfaces, where the vapor and liquid condensate flows are not restricted by some overall flow-channel dimensions. Condensation inside tubes is of

.considerable practical interest because of applications to condensers in refrigeration and air-conditioning systems, but unfortunately these phe-nomena are quite complicated and not amenable to a simple analytical treatment. The overall flow rate of vapor strongly influences the heat-transfer rate in the forced convection-condensation system, and this in turn is influenced by the rate of liquid accumulation, on the walls. Because of the

'complicated flow phenomena involved we shall present only two empirical relations for heat transfer and refer to reader to Rohsenow [371 for more

Xmplete i1normatioft.

Chato [381 obtained the following expression for condensation of refriger-ants at low vapor velocities inside horizontal tubes:

g;-i* r =(, bSV(P - p,,gk.*.;A '

0y15

VERMONT YANKEE NEC-JH_32 NUCLEAR POWER CORPORATION 185 OLD FERRY ROAD, PO BOX 7002, BRATTLEBORO, VT 05302-7002 (802) 257-5271 August 20, 2001 BVY 01-66 U.S. Nuclear Regulatory Commission ATTN: Document Control Desk Washington, D.C. 20555

Subject:

Vermont Yankee Nuclear Power Station License No. DPR-28 (Docket No. 50-271)

Vermont Yankee 2001 Summary Reports for In-service Inspection and Repairs or Replacements In accordance with Article IWA-6000 of Section XI of the ASME Boiler and Pressure Vessel Code, Vermont Yankee (VY) hereby submits the Owner's Report for In-service Inspections (Form NIS-]) and the Owner's Report for Repairs and Replacements (Form NIS-2). These reports describe the in-service examinations, repairs and replacements performed during the period from December 4, 1999 to May 20, 2001 (including Refueling Outage 22). VY's third ten-year interval began September 1, 1993.

We trust that the information provided is adequate; however, should you have questions or require additional information, please contact Mr. Jim DeVincentis at (802) 258-4236.

Sincerely, VERMONT YANKEE NUCLEAR POWER CORPORATION AGautam Sen Licensing Manager Attachments cc: USNRC Region 1 Administrator USNRC Resident Inspector - VYNPS AD~l USNRC Project Manager - VYNPS Vermont Department of Public Service Inspection Agency - A*rkwright

I I

SUMMARY

OF VERMONT YANKEE COMMITMENTS I

BVY NO.: 01-66 The following table identifies commitments made in this document by Vermont Yankee.

I Any other actions discussed in the submittal represent intended or planned actions by Vermont Yankee. They are described to the NRC for the NRC's information and are not regulatory commitments. Please notify the Licensing Manager of any questions regarding this document or any associated commitments.

I COMMITMENT COMMITTED DATE OR "OUTAGE" I

None N/A I I

I I

I I

I I

VYAPF 0058.04 AP 0058 Original Page 1 of 1 I

I I

I I

Vermont Yankee Nuclear Power Corporation 2001 Form NIS-1 Owner's Summary Report for Inservice Inspections December 4, 1999 through May 20, 2001 Reviewed by: PlntIsrvice f~cI o -,ordinafo~r20 Approved by:

System Enginekring Manager Page 1 of 20

FORM NIS-1 OWNER'S REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code Rules

. Owner Vermont Yankee Nuclear Power Corporation, 185 Old Ferry Road, PO Box 7002, Brattleboro VT 05302-7002 (Name and Address ot Owner)

2. Plant Vermont Yankee Nuclear Power Station. P.O. Box 157, Governor Hunt Road, Vernon, VT 05354-0157 (Name and Address ot Plant)

3. Plant Unit 1 4. Owner Certificate of Authorization (if required) DPR-28

5. Commercial Service Date 11/30/1972 6. National Board Number for Unit NONE

7. Components Inspected - SEE ATTACHED PAGES 2 THROUGH 13.

8. Examination Dates 12/04/1999 to 05/20/2001 9. Inspection Interval from 09/1/1993 to 08/31/2003

10. Applicable Editions of Section XI 1986, no Addenda, 1992 w/1992 Addenda (IWE) and 1995 w/1996 Addenda (ASME Appendix VIII) 3

11. Abstract of Examinations Including a list of examinations and a statement concerning status of work required for current interval - SEE ATTACHED PAGES 2 THROUGH 21.

12. Abstract of Conditions Noted - SEE ATTACHED PAGES 22 THROUGH 25.

13. Abstract of Corrective Measures Recommended and Taken - SEE ATTACHED PAGES 22 THROUGH 25.

I We certify that the statements made in this report are correct and the examinations and corrective measures taken conform to the rules of the ASME Code,Section XI.

Certificate of Authorization No. (If applicable) P 8 Ex ation Date 3/21/2012 I yd A ,'i&

f Date 2 20AA Signed /*-Pf ReBy a

f w~ner CERTIFICATE OF INSERVICE INSPECTION I, the undersigned, holding a valid commission issued by the National Board of Boiler and Pressure Vessel Inspectors and/or the State or Province of Vermont and employed by Factory Mutual Insurance Co. of Johnston RI have inspected the components described in this Owner's Report during the period December 4, 1999 to May 20, 2001 and state to the best of my knowledge and belief, the Owner has performed examinations and taken corrective measures described in this Owner's Report in accordance with the requirements of the ASME Code,Section XI.

By signing this certificate neither the inspector nor his employer makes any warranty, expressed or implied, concerning the examinations and corrective measures described in this Owner's Report. Furthermore, neither the Inspector nor his employer shall be liable in any manner for any personal injury or property damage or a loss of any kind arising from or connected with this inspection.

Commissions VT-345 Inspector's Signature National Board, State, Province, and Endorsements Date * - J/ " .,20ol_ I Page 2 of 20

- -- m -- m- - - m - m mm - m m m m m FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 Components Inspected/Abstract of examinations Sections 7 and 11 ASME Category Component ID Exam Type System ID DrawingNo. ExaminationResults B-D N2F UT Nuclear Boiler ISI-RPV-103 Acceptable B-D N2F-IR UT Inner Radius Nuclear Boiler ISI-RPV-103 Acceptable B-D N2G UT Nuclear Boiler ISI-RPV-103 Acceptable B-D N2G-IR UT Inner Radius Nuclear Boiler ISI-RPV- 103 Acceptable B-D N2H UT Nuclear Boiler ISI-RPV- 103 Acceptable B-D N2H-IR UT Inner Radius Nuclear Boiler ISI-RPV-103 Acceptable B-D N2J UT Nuclear Boiler ISI-RPV- 103 Acceptable B-D N2J-IR UT Inner Radius Nuclear Boiler ISI-RPV-103 Acceptable B-D N2K UT Nuclear Boiler ISI-RPV-103 Acceptable B-D N2K-IR UT Inner Radius Nuclear Boiler ISI-RPV- 103 Acceptable Acceptable -

Automated Inner Radius examination in accordance with B-D N4A-IR UT Inner Radius Feedwater ISI-RPV- 103 General Electric Nuclear Energy document GE-NE-523-A71-0594-A, Revision 1 and VY Calculation VYC-1005 Acceptable -

Automated Inner Radius examination inaccordance with B-D N4B-IR UT Inner Radius Feedwater ISI-RPV- 103 General Electric Nuclear Energy document GE-NE-523-A71-0594-A, Revision 1 and VY Calculation VYC-1005 Page 3 of 20

FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 Components Inspected/Abstract of examinations Sections 7 and 11 ASME Category Component ID Exam Type System ID DrawingNo. Examination Results Acceptable -

Automated Inner Radius examination inaccordance with B-D N4C-IR UT Inner Radius Feedwater ISI-RPV- 103 General Electric Nuclear Energy document GE-NE-523-A71-0594-A, Revision I and VY Calculation VYC-1005 Acceptable -

Automated Inner Radius examination inaccordance with B-D N4D-IR UT Inner Radius Feedwater ISI-RPV- 103 General Electric Nuclear Energy document GE-NE-523-A71-0594-A,Revision I and VY Calculation VYC-1005 B-F Ni1A WELD PT Nuclear Boiler ISI-RPV-103 Acceptable B-F NiBIWELD PT Nuclear Boiler ISI-RPV-103 Acceptable B-F N12A-SE PT Nuclear Boiler ISI-RPV-103 Acceptable Acceptable -

B-F N1B-SE PT Nuclear Boiler ISI-RPV-103 Examination performed as follow up to indication removal during RFO-21 B-F N2F-SE UT, PT Nuclear Boiler ISI-RPV-103 Acceptable B-F N2G-SE UT, PT Nuclear Boiler ISI-RPV- 103 Acceptable B-F N2H-SE UT, PT Nuclear Boiler ISI-RPV-103 Acceptable B-F N2J-SE UT, PT Nuclear Boiler ISI-RPV-103 Acceptable B-F N2K-SE UT, PT Nuclear Boiler ISI-RPV-103 Acceptable Page 4 of 20 M M mM M m M M--- M M - M M -M M

m - m - m - -

m- - -m--m -- - m-m FORM NIS-I OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 Components Inspected/Abstract of examinations Sections 7 and 11 ASME Category Component ID Exam Type System ID DrawingNo. Examination Results B-F N6B-SE UT, PT Nuclear Boiler ISI-RPV-103 Acceptable B-F N8A-SE UT, PT Nuclear Boiler ISI-RPV-103 Acceptable B-F N8B-SE UT, PT Nuclear Boiler ISI-RPV-103 Acceptable O1A-N/W Acceptable -

B-G-1 Recirculation Pump P-18-IA VT-1 Nuclear Boiler ISI-RPV- 104 IDR # 01-09 generated for Bolting corrosion/plating concern - See Sections 12 and 13 02A-N/W Acceptable -

B-G-1 Recirculation Pump P-18-1A VT-I Nuclear Boiler ISI-RPV- 104 IDR # 01-09 generated for Bolting corrosioniplating concern - See Sections 12 and 13 03A-N/W Acceptable -

B-G-1 Recirculation Pump P- 18-I A VT- 1 Nuclear Boiler ISI-RPV-104 IDR # 01-09 generated for Bolting corrosion/plating concern - See Sections 12 and 13 04A-N/W Acceptable -

B-G- I Recirculation Pump P-18-1A VT-I Nuclear Boiler ISI-RPV-104 IDR # 01-09 generated for Bolting corrosion/plating concern - See Sections 12 and 13 05A-N/W Acceptable -

B-G-1 Recirculation Pump P-18-IA VT- I Nuclear Boiler ISI-RPV- 104 IDR # 01-09 generated for Bolting corrosion/plating concern - See Sections 12 and 13 06A-N/W Acceptable -

B-G-1 Recirculation Pump P-18-I A VT-I Nuclear Boiler ISI-RPV-104 IDR # 01-09 generated for corrosion/plating concern - See Bolting Sections 12 and 13 07A-N/W Acceptable -

B-G-1 Recirculation Pump P-18-IA VT-I Nuclear Boiler ISI-RPV- 104 IDR # 01-09 generated for corrosion/plating concern - See Bolting Sections 12 and 13 Page 5 of 20

FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 Components Inspected/Abstract of examinations Sections 7 and 11 ASME Category Component ID Exam Type System ID DrawingNo. Examination Results 08A-N/W Acceptable -

B-G- 1 Recirculation Pump P-18-IA VT- I Nuclear Boiler ISI-RPV-104 IDR # 01-09 generated for Bolting corrosion/plating concern - See Sections 12 and 13 09A-N/W Acceptable -

B-G- I Recirculation Pump P- 18-IA VT- I Nuclear Boiler ISI-RPV- 104 IDR # 01-09 generated for Bolting corrosion/plating concern - See Sections 12 and 13 1OA-NIW Acceptable -

B-G- 1 Recirculation Pump P-18-IA VT- I Nuclear Boiler ISI-RPV- 104 IDR # 01-09 generated for Bolting corrosion/plating Sections 12concern and 13 - See 11 A-N/W Acceptable -

B-G-I Recirculation Pump P-1 8-1A VT-I Nuclear Boiler ISI-RPV- 104 IDR # 01-09 generated for Boltingcorrosion/plating concern - See Bolting Sections 12 and 13 12A-N/W Acceptable -

B-G- 1 Recirculation Pump P-18-IA VT- I Nuclear Boiler ISI-RPV- 104 IDR # 01-09 generated for Boltingcorrosion/plating concern - See Bolting Sections 12 and 13 13A-N/W Acceptable -

B-G- 1 Recirculation Pump P-18-IA VT- I Nuclear Boiler ISI-RPV- 104 IDR # 01-09 generated for Bolting corrosion/plating concern - See Sections 12 and 13 14A-N/W Acceptable -

B-G- I Recirculation Pump P-18-IA VT- I Nuclear Boiler ISI-RPV- 104 IDR # 01-09 generated for concern- See Boltingcorrosion/plating Bolting Sections 12 and 13 15A-N/W Acceptable -

B-G- I Recirculation Pump P-18-IA VT- I Nuclear Boiler ISI-RPV-104 IDR # 01-09 generated for corrosion/plating concern - See Bolting Sections 12 and 13 Page 6 of 20 M M---i M M M- M --M M M M M M M

- m-- m -mrn-- - - rnm n - rn m rn -

FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 Components Inspected/Abstract of examinations Sections 7 and 1I ASME Category ComponentID Exam Type System ID DrawingNo. ExaminationResults 16A-N/W Acceptable -

B-G-1 Recirculation Pump P-18-IA VT- I Nuclear Boiler ISI-RPV-104 IDR # 01-09 generated for Bolting corrosion/plating concern - See Sections 12 and 13 B-J CS4B-F3ADW UT, PT Core Spray ISI-5920-9206 Acceptable B-J CS4B-MF5 UT, PT Core Spray ISI-5920-9206 Acceptable B-J CS4B-MF5B UT, PT Core Spray ISI-5920-9206 Acceptable B-i CS4B-MF6A UT, PT Core Spray ISI-5920-9206 Acceptable B-J FW20-Fl UT/FAC Feedwater ISI-FDW-PART 5A Acceptable -

Code Case N-560 examination B-i FW20-F1B UT/FAC Feedwater ISI-FDW-PART 5A Acceptable -

Code Case N-560 examination B-J FW20-F3B UT/FAC Feedwater ISI-FDW-PART 5A Acceptable -

Code Case N-560 examination B-J SL 11-F28 PT Standby Liquid Control ISI-SLC-PART 4 Acceptable B-J SLI 1-F29 PT Standby Liquid Control ISI-SLC-PART 4 Acceptable B-K 270 DEG RPV BRKT PT Nuclear Boiler ISI-RPV-103 Acceptable B-K RPV SUPPORT SKIRT MT Nuclear Boiler ISI-RPV-103 Acceptable B-K RR-34 PT Nuclear Boiler ISI-5920-6802 Sh.2 Acceptable B-K RR-35 PT Nuclear Boiler ISI-5920-6802 Sh.2 Acceptable B-O 26-03SH PT Control Rod Drive ISI-RPV-104 Acceptable B-O 34-39SH PT Control Rod Drive ISI-RPV-104 Acceptable Page 7 of 20

FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 Components Inspected/Abstract of examinations Sections 7 and 11 ASME Category Component ID Exam Type System ID DrawingNo. Examination Results C-C ACSP-H22 MT Standby Gas Treatment ISI-5920-9200 Acceptable C-C ACSP-H23 MT Standby Gas Treatment ISI-5920-9200 Acceptable C-C RHR-H192 MT Residual Heat Removal ISI-RHR-PART 11 Sh.4 Acceptable C-C RHR-H98 MT Residual Heat Removal ISI-5920-9208 Acceptable C-C RHR-HD25 PT Residual Heat Removal ISI-RHR-PART 16 Sh.1 Acceptable -

Successive examination C-F-2 CR4A-S5 UT, MT Control Rod Drive ISI-5920-9528 Acceptable C-F-2 CR6A-S57 UT, MT Control Rod Drive ISI-5920-9527 Acceptable C-F-2 CR6-S 10 UT, MT Control Rod Drive ISI-5920-9527 Acceptable C-F-2 CR6-$22 UT, MT Control Rod Drive ISI-5920-9527 Acceptable C-F-2 CR6-$26 UT, MT Control Rod Drive ISI-5920-9527 Acceptable C-F-2 CS1B-S30 UT, MT Core Spray ISI-5920-9210 Acceptable C-F-2 CT27-$30 UT, MT Core Spray ISI-5920-9210 Acceptable C-F-2 FW17-S5 UT, MT Feedwater ISI-FDW-PART 5A Acceptable High Pressure Coolant C-F-2 HP15A-S101 UT, MT Injection ISI-HPCI-PART 5 Acceptable C-F-2 RHl4-T373 UT, MT Core Spray ISI-5920-9208 Acceptable C-F-2 RH1B-S47 UT, MT Residual Heat Removal ISI-5920-9285 Acceptable C-F-2 RH2B-S1 13 UT, MT Residual Heat Removal ISI-5920-9285 Acceptable Page 8 of 20

-Mm-M =- M M-- M M = - - MM -

- - m -m-n - m-m - - - - mn - mmm FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 Components Inspected/Abstract of examinations Sections 7 and 11 ASME Category Component ID Exam Type System ID Drawing No. Examination Results C-F-2 RH2B-S 115 UT, MT Residual Heat Removal ISI-5920-9285 Acceptable C-F-2 RH3B-S170 UT, MT Residual Heat Removal ISI-5920-9288 Acceptable C-F-2 RH3D-S200 UT, MT Residual Heat Removal ISI-5920-9288 Acceptable C-F-2 RH3D-S206 UT, MT Residual Heat Removal ISI-5920-9288 Acceptable C-F-2 RH3D-T182 UT, MT Residual Heat Removal ISI-5920-9288 Acceptable C-F-2 RH7-S284 UT, MT Residual Heat Removal ISI-5920-9287 Acceptable C-F-2 RH9-S314 UT, MT Residual Heat Removal ISI-RHR-PART 11 Sh.4 Acceptable C-F-2 RH9-S320 UT, MT Residual Heat Removal ISI-RHR-PART I1 Sh.4 Acceptable C-FAUG CS2A-S62 UT, MT Core Spray ISI-5920-9211 Acceptable C-FAUG CS2A-S64 UT, MT Core Spray ISI-5920-9211 Acceptable C-FAUG CS2A-S65 UT, MT Core Spray ISI-5920-9211 Acceptable C-FAUG CS2A-S67 UT, MT Core Spray ISI-5920-9211 Acceptable C-FAUG CT1-S54 UT, PT Core Spray ISI-CST-PART 4 Acceptable C-FAUG CTl-S56 UT, PT Core Spray ISI-CST-PART 4 Acceptable

__ -____U_______

C-FAUG 3-S______3 RC3-S13 __ UT,________Cooling UT, MT Reactor Core Isolation __ SI-5920-9255 ISI-5920-9255 _ Acceptable____

Acceptable RatrCoreolting C-FAUG RC3-S14 UT, MT Reactor Core Isolation ISI-5920-9255 Acceptable

__________

___ ______

_________ ___ ______ _________Cooling_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

Page 9 of 20

FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 Components Inspected/Abstract of examinations Sections 7 and I1 ASME Category Component ID Exam Type System ID DrawingNo. ExaminationResults C-NAUG MSlD-F9 UT, MT Main Steam 5920-FS-Il Acceptable C-NAUG MS2D-F1 UT, MT Main Steam 5920-FS-I1 Acceptable D-A RSW-H171 VT-1 Residual Heat Removal ISI-SW-PART 9 Acceptable Acceptable -

D-A RSW-H261 VT- I Residual Heat Removal ISI-SW-PART 9 IDR # 01-01 generated for arc strikes - See Sections 12 and 13 Acceptable -

D-A RSW-HD261B VT-1 Residual Heat Removal ISI-SW-PART 9 IDR# 01-01 generated for arc strikes - See Sections 12 and 13 Acceptable -

General Visual Examination as required by ASME Subsection IWE has been 100% completed for the first period of the first IWE Interval.

5920-13, 5920-41, IDR # 01-07 generated for pitting E-A Class MC Containment General Visual Class MC Containment 6202-200 and general corrosion inthe Vent Header. IDR # 01-07 generated for pitting and general corrosion inthe Vent Header. Also, IDRf# 01-08 was generated for general corrosion and material loss in Penetrations X-207A through X-207H. See Sections 12 and 13 E-A Vent Line Areas (X-5B) VT-I Class MC Containment 5920-13 Acceptable E-A Vent Line Areas (X-5C) VT- 1 Class MC Containment 5920-13 Acceptable E-A Vent Line Areas (X-5D) VT-I Class MC Contaimnent 5920-13 Acceptable E-A Vent Line Areas (X-5E) VT- I Class MC Containment 5920-13 Acceptable Page 10 of 20

= m= Mm M - = =-- M = = = M = - m

m - --- m-m- n mm- m - - m - mm m FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 Components Inspected/Abstract of examinations Sections 7 and 11 ASME Category Component ID Exam Type System ID DrawingNo. Examination Results E-C Drywell Seal Area VT-I Class MC Containment 6202-2 Acceptable E-C Drywell Seal Area VT-3 Class MC Containment 6202-2 Acceptable E-G Pen. X-200A VT-I Class MC Containment 6202-208 Acceptable E-G Pen. X-200B VT-I Class MC Containment 6202-208 Acceptable E-G V16-19-5A VT- I Class MC Containment 5920-675 Acceptable E-G V16-19-5B VT- I Class MC Containment 5920-675 Acceptable E-G V16-19-5C VT-i Class MC Containment 5920-675 Acceptable E-G V16-19-5D VT-I Class MC Containment 5920-675 Acceptable E-G V 16-19-5E VT-I Class MC Containment 5920-675 Acceptable E-G V16-19-5F VT- 1 Class MC Containment 5920-675 Acceptable E-G V16-19-5H VT- I Class MC Containment 5920-675 Acceptable F-A ACSP-H22 VT-3 Standby Gas Treatment ISI-5920-9200 Acceptable F-A ACSP-H23 VT-3 Standby Gas Treatment ISI-5920-9200 Acceptable F-A CS-HD60A VT-3 Core Spray ISI-5920-9210 Acceptable Acceptable -

F-A FDW-HD39 VT-3 Feedwater ISI-FDW-PART 5A IDR,#01-12 generated for debris/corrosion - See Sections 12 and 13 F-A H-P-44-1B VT-3 High Pressure Coolant ISI-HPCI-PART 13A Acceptable age1 0 Injection Page 11I of 20

FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 Components Inspected/Abstract of examinations Sections 7 and 11 ASMIE Category Component ID Exam Type System ID Drawing No. ExaminationResults F-A HCI-1VT-3High Pressure Coolant F-A HPCI- I VT-3 Ig eecoon ISI-4PCI-PART 2 Acceptable F-A HPCI-2 VT-3 High Pressure Coolant ISI-HPCI-PART 2 Acceptable Injection F-A RHR-H129 VT-3 Residual Heat Removal ISI-5920-9288 Acceptable F-A RHR-H191 VT-3 Residual Heat Removal ISI-RHR-PART 11 Sh.4 Acceptable Acceptable -

F-A RHR-H192 VT-3 Residual Heat Removal ISI-RHR-PART I1 Sh.4 IDR # 01-02 generated for gouge -

See Sections 12 and 13 F-A RHR-H83 VT-3 Residual Heat Removal ISI-RHR-PART 11 Sh.4 Acceptable F-A RHR-H98 VT-3 Residual Heat Removal ISI-5920-9208 Acceptable F-A RHR-HD127C VT-3 Residual Heat Removal ISI-5920-9285 Acceptable F-A RHR-HD127E VT-3 Residual Heat Removal ISI-5920-9285 Acceptable -

Successive Examination F-A RHR-HDI27G VT-3 Residual Heat Removal ISI-5920-9285 Acceptable F-A RHR-HD188A VT-3 Residual Heat Removal ISI-5920-9288 Acceptable F-A RPV SUPPORT SKIRT VT-3 Nuclear Boiler ISI-RPV-103 Acceptable F-A RR-15 VT-3 Nuclear Boiler ISI-5920-6802 Sh.2 Acceptable F-A RR- 16 VT-3 Nuclear Boiler ISI-5920-6802 Sh.2 Acceptable F-A RR-17 VT-3 Nuclear Boiler ISI-5920-6802 Sh.2 Acceptable F-A RR-2 VT-3 Nuclear Boiler ISI-5920-6802 Sh.2 Acceptable Page 12 of 20

= m = M M M M m=- = M = = -M = = m

FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 Components Inspected/Abstract of examinations Sections 7 and 11 ASME Category Component ID Exam Type System ID DrawingNo. Examination Results F-A RR-35 VT-3 Nuclear Boiler ISI-5920-6802 Sh.2 Acceptable Acceptable -

F-A RR-44 VT-3 Nuclear Boiler ISI-5920-6802 Sh.2 IDR # 01-04 generated for setting -

See Sections 12 and 13 Acceptable -

F-A RR-52 VT-3 Nuclear Boiler ISI-5920-6802 Sh.2 IDR # 01-05 generated for setting -

See Sections 12 and 13 Acceptable -

F-A RR-7,8 VT-3 Nuclear Boiler ISI-5920-6802 Sh.2 IDR # 01-06 generated for setting -

See Sections 12 and 13 F-A RSW-H167 VT-3 Residual Heat Removal ISI-SW-PART 1 Sh.2 Acceptable F-A RSW-H171 VT-3 Residual Heat Removal ISI-SW-PART 9 Acceptable F-A RSW-H172 VT-3 Residual Heat Removal ISI-SW-PART 6 Sh. I Acceptable F-A RSW-H241 VT-3 Residual Heat Removal ISI-SW-PART I Sh.2 Acceptable F-A RSW-H261 VT-3 Residual Heat Removal ISI-SW-PART 9 Acceptable F-A RSW-HD261B VT-3 Residual Heat Removal ISI-SW-PART 9 Acceptable F-A SDV-N-R02 VT-3 Control Rod Drive ISI-5920-9527 Acceptable F-A SDV-N-RO5 VT-3 Control Rod Drive ISI-5920-9527 Acceptable F-AUG ACSP-H203 VT-3 Standby Gas Treatment ISI-5920-9201 Acceptable F-AUG ACSP-HD203E VT-3 Standby Gas Treatment ISI-5920-9201 Acceptable F-AUG ACSP-HD203F VT-3 Standby Gas Treatment ISI-5920-9201 Acceptable Page 13 of 20

FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 Components Inspected/Abstract of examinations Sections 7 and 11 ASME Category Component ID Exam Type System ID Drawing No. Examination Results F-AUG RHR-HD25 VT-3, PT Residual Heat Removal ISI-RHR-PART 16 Sh.1 Acceptable N/A ACSP-H201B VT-3 Standby Gas Treatment ISI-AC PART 5 Acceptable NNS HPCI-HD28 VT-3 High Pressure Coolant ISI-HPCI-PART 4 Sh. 1 Sh.1_AcceptableAcceptable Injection+___ISI-HPCIPART_4 NNS RHR-HD235 VT-3 Residual Heat Removal VYI-RHR-PART 7B Acceptable Page 14 of 20 M- M M M-- M- M M M M- M M- -

- n--mm -- -mm - m m m-m m m -

FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 Components Inspected/Abstract of examinations Sections 7 and 11 Quantity Inspected Quantity Previously Quantity Scheduled, Percentof Third Code Category Inspected, Third Third Interval Interval Complete 2001 Outage Interval B-A 0 15 16 94%

B-D 10 38 58 83%

B-F 12 19 35 89%

B-G-1 16 152 288 58%

B-G-2 0 77 109 71%

0 60%

B-J (Code Case N-560 was first used for 15 (Previously 64% of the standard (Code Case N-560 selection) selection during RFO-22) ASME Category B-J 25% selection had been completed)

B-J (These are ASME Category B-J, Item B9.40 socket welds which are 2 14 23 70%

not included in the Code Case N-560 selection. They are selected in accordance with Category B-J 25%

criteria.)

B-K 4 3 10 70%

B-L-2 0 0 Per approved Relief Request N/A No. B-i B-M-2 0 26 Per approved Relief Request N/A No. B-2 Page 15 of 20

FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 Components Inspected/Abstract of examinations Sections 7 and 11 yQuantity Inspected Quantity Previously Quantity Scheduled, Percentof Third CodeCatgor 200 OuageInterval 2001 Outage Interd Third Interval Interval Complete B-N-1 0. 2 .Each Period N/A B-N-2 0 Partial I N/A B-O 2 4 7 86%

C-A 0 3 4 75%

C-B 0 6 8 75%

C-C 4 12 20 80%

C-F-2, 20 43 72 88%

D-A 3 6 11 82%

E-A 20% 80% 100% 100%

E-C 100% 100% 100% 100%

E-D E-G F-A 33 59 119 77%

Page 16 of 20

- m m m m -u m-- - m m -m m-m -

FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 ABSTRACT OF CONDITIONS NOTED/CORRECTIVE MEASURES TAKEN Sections 12 and 13 Code Category Item Identification Conditions Noted and CorrectiveMeasures Taken VT-I examination of Recirculation pump P-18-1A bolting identified possible missing protective thread coating (the bolting was examined in place, under tension). The examination also revealed corrosion on 0 1 A-N/W through 16A- the exposed bolting. Inservice Discrepancy Report # 01-09 was generated to request Engineering B-G-1 N/W evaluation of these indications. Technical Evaluation (TE) 2001-034 was generated and contains: The pump casing cover/body bolting will perform its design function with the as-noted surface conditions.

Margin exists in the 2 1/2' diameter cap screws for future corrosion. No additional and/or augmented inspections other than planned inservice inspection is required.

VT-3 examination of rigid strut support RHR-H 192 revealed a gouge on the pipe clamp. Inservice Discrepancy Report # 01-02 was generated to request Mechanical Design Engineering evaluation of this C-C RHR-H 192 condition. Technical Evaluation 2001-015 was issued containing: a) The gouge does not extend behind the lugs therefore the lugs have full bearing surface on the clamp. b) The reduction of a maximum of 1/32" of depth on an 8" deep clamp is insignificant (<1%). These indications were determined to be caused during initial installation/fabrication.

VT-3 examination of rigid frame support RSW-H 172 revealed a crack in the concrete wall adjacent to the base plate. Inservice Discrepancy Report # 01-03 was generated to request Mechanical Design Engineering evaluation of this condition. Technical Evaluation 2001-017 was issued containing: a) The D-A RSW-H172 support and the associated anchor bolts will perform their intended design functions in the as-found condition. b) The cracking has been determined to be a surface hairline crack that is the result of normal aging and/or as-expected normal shrinkage cracking of the concrete. This area is monitored in accordance with Vermont Yankee procedure PP 7030 "Structures Monitoring Program" which implements 10 CFR 50.65.

Page 17 of 20

FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 ABSTRACT OF CONDITIONS NOTED/CORRECTIVE MEASURES TAKEN Sections 12 and 13 Code Category Item Identification ConditionsNoted and CorrectiveMeasures Taken VT-3 examination of spring hanger RSW-H261 revealed several arc strikes and a poor weld profile on integrally attached pipe lugs (these lugs are used in common with spring hanger RSW-HD261B - see below). Inservice Discrepancy Report # 01-01 was generated to request Mechanical Design Engineering evaluation of these conditions. Technical Evaluation 2001-014 was issued containing: a)

D-A RSW-H261 None of the arc strikes contained cracking b) The maximum recordable depth of any arc strike was .03" c) No overstress conditions exist, d) The weld in question is an "extra weld" not called for in the engineering qualification of the pipe lug (the lug is only required to be welded on 2 sides, this weld is on the third (not required) side. These indications were determined to be caused during initial installation or modification.

VT-3 examination of spring hanger RSW-HD261B revealed several arc strikes and a poor weld profile on integrally attached pipe lugs (these lugs are used in common with spring hanger RSW-H261 - see above). Inservice Discrepancy Report # 01-01 was generated to request Mechanical Design Engineering evaluation of these conditions. Technical Evaluation 2001-014 was issued containing: a) None of the arc D-A RSW-HD261B strikes contained cracking b) The maximum recordable depth of any arc strike was .03" c) No overstress conditions exist. d) The weld in question is an "extra weld" not called for in the engineering qualification of the pipe lug (the lug is only required to be welded on 2 sides, this weld is on the third (not required) side. These indications were determined to be caused during initial installation or modification.

During General Visual examination general corrosion and material loss was found. Inservice Penetrations X-207A Discrepancy Report # 01-08 was generated to request Mechanical Design Engineering evaluation of this E-A Pthratonsh X-20condition. Technical Evaluation (TE) 2001-025 was generated and contains: The condition is acceptable as found as there is significant margin remaining to code minimum wall thickness accompanied by a low expected rate of galvanic corrosion in the inerted containment.

Page 18 of 20 m-m M-m

= M = -= -m-m M = M M m

/-- I- --- - m - - - I m m -

FORM NIS-1 OWNER'S DATA REPORT FOR ENSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 ABSTRACT OF CONDITIONS NOTED/CORRECTIVE MEASURES TAKEN Sections 12 and 13 Code Category Item Identification ConditionsNoted and CorrectiveMeasures Taken During General Visual examination pitting and general corrosion in excess of the allowable values provided by Mechanical Design Engineering were found. The corrosion and pitting are accompanied by loss of coating. There was also significant standing water in Vent Header bowl H. Inservice Discrepancy E-A Vent Header Report # 01-07 was generated to request Mechanical Design Engineering evaluation of this condition.

Technical Evaluation (TE) 2001-025 was generated and contains: a) The observed pitting in the Vent Header is acceptable. b) The standing water was removed and the source was identified and corrected prior to drywell closeout.

VT-3 examination of anchor FDW-HD39 revealed debris in the form of paint chips, insulation and minor corrosion in a required 1/16" gap between a trunion and the base plate. Inservice Discrepancy Report # 0 1-12 was generated to request Mechanical Design Engineering evaluation of this condition.

F-A FDW-HD39 Technical Evaluation (TE) 2001-038 was generated and contains: The as found condition of the support/anchor is acceptable with the exception of the identified debris. The trunions are not "bound up" restricting thermal growth/movement of the pipe, and the debris does not adversely impact the overall function of the support/anchor. The debris was subsequently cleaned from the anchor.

VT-3 examination of spring hanger RR-44 revealed a spring can setting that was out of tolerance by greater than +/- 5% provided by Mechanical Design Engineering. Inservice Discrepancy Report # 01-04 was generated to request Mechanical Design Engineering evaluation of this condition. Technical Evaluation (TE) 2001-021 was generated and contains: The setting was determined to have not affected the supports structural or functional capability. It was noted that the support was "adjusted" as far as possible, i.e., there was no more thread remaining on the rod at the adjustment nut. This condition will be revisited during the next extended refueling outage (RFO-23) to determine if any further action would be warranted.

Page 19 of 20

FORM NIS-1 OWNER'S DATA REPORT FOR INSERVICE INSPECTIONS As Required by the Provisions of the ASME Code rules Vermont Yankee Nuclear Power Corporation Vermont Yankee Nuclear Power Station Owner Certification: DPR-28 Commercial Service Date: 11/30/72 ABSTRACT OF CONDITIONS NOTED/CORRECTIVE MEASURES TAKEN Sections 12 and 13 Code Category Item Identification ConditionsNoted and Corrective Measures Taken VT-3 examination of spring hanger RR-52 revealed a spring can setting that was out of tolerance by greater than +/- 5% provided by Mechanical Design Engineering. Inservice Discrepancy Report # 01-05 was generated to request Mechanical Design Engineering evaluation of this condition. Technical Evaluation (TE) 2001-022 was generated-and contains:The setting was determined to have not affected the supports structural or functional capability. It was noted that the support was "adjusted" as far as possible, i.e., there was no more thread remaining on the rod at the adjustment nut. This condition will be revisited during the next extended refueling outage (RFO-23) to determine if any further action would be warranted.

VT-3 examination of spring hanger RR-7, 8 revealed a spring can setting that was out of tolerance by greater than +/- 5% provided by Mechanical Design Engineering. Inservice Discrepancy Report # 0 1-06 was generated to request Mechanical Design Engineering evaluation of this condition. Technical F-A RR-7, 8 Evaluation (TE) 2001-023 was generated and contains: a) The setting was determined to have not affected the supports structural or functional capability. b) The spring cans are capable of performing their intended design function in the as-found, as-leftcondition. This condition will be revisited during the next extended refueling outage (RFO-23) to determine if any further action would be warranted.

Page 20 of 20 M M M M n M = = - = M = = = M = -

I_

I I

Vermont Yankee Nuclear Power Corporation I 2001 Form NIS-2 Owner's Summary Report for I Repairs or Replacements i December 4, 1999 through May 20, 2001 Reviewed by:

Plant Inservice lns6eqtýon Qjordinator Approved by:

System Engi eering ManageY Page 1 of 14

FORM NIS-2 OWNER'S REPORT FOR REPAIRS OR REPLACEMENTS As required by the Provisions of the ASME Code Section XI

1. Owner Vermont Yankee Nuclear Power Corporation Name Date I

185 Old Ferry Road, PO Box 7002, Brattleboro VT 05302-7002 Sheet 2 of 14 Address

2. Plant Vermont Yankee Nuclear Power Station Unit I Name P.O. Box 157, Governor Hunt Road, Vernon, VT 05354-0157 N/A Address Repair Organization P.O. No., Job No., etc.

3. Work Performed by Vermont Yankee Nuclear Power Corporation Type Code Symbol Stamp N/A I

Name Authorization No. N/A 185 Old Ferry Road, PO Box 7002, Brattleboro VT 05302-7002 Expiration Date N/A Address

4. Identification of System See attached table, pages 4 through 14

5. (a) Applicable Construction Code B.3 1.1 1967 Edition, No " Addenda, No Code Case (b) Applicable Edition of Section XI Utilized for Repairs or Replacements 1986 Edition No Addenda

6. Identification of Components Repaired or Replaced and Replacement Components See attached table, pages 4 through 14

7. Description of Work See attached table, pages 4 through 14 I

8. Tests Conducted See attached table, pages 4 through 14 I

9. Remarks See attached table, pages 4 through 14 Page 2 of 14

I Form NIS-2 (cont.) Sheet 3 of 14 CERTIFICATE OF COMPLIANCE We certify that the statements made in the report are correct and these repairs/replacements conform to the rules of ASME Code,Section XI.

Type Code Symbol Stamp N/A Certificate of Authorization mber N/A Expiration Date N/A Date J ,202'* '

Si-ned t,

7? DaeY26 7

'Di ý, # ce Prnesilnt,

-neering CERTIFICATE OF INSERVICE INSPECTION I, the undersigned, holding a valid commission issued by the National Board of Boiler and Pressure Vessel Inspectors and the State or Province of Vermont and employed by Factory Mutual Insurance Co. of Johnston RI have inspected the components described in this Owner's Report during the period December 4, 1999 to May 20, 2001 and state to the best of my knowledge and belief, the Owner has performed examinations and taken corrective measures described in this Owner's Report in accordance with the requirements of the ASME Code,Section XI.

By signing this certificate neither the inspector nor his employer makes any warranty, expressed or implied, concerning the examinations and corrective measures described in this Owner's Report. Furthermore, neither the inspector nor his employer shall be liable in any manner for any personal injury or property damage or a loss of any kind arising from or connected with this inspection.

_--_ _ _ _ _ Commissions VT-345 Inspector's Signature National Board, State, Province, and Endorsements Date _ _ _ _ _20o, Page 3 of 14

FORM NIS-2 OWNER'S REPORT FOR REPAIRS OR REPLACEMENTS As required by the provisions of the ASME Code,Section XI, 1986 Edition, No Addenda Vermont Yankee Nuclear Power Plant Unit 1 P.O. Box 157, Vernon, VT, 05354 Construction Code B31.1, 1967 Edition, No Addenda, No Code Case Cmponent SNational Other Identification Repaired Component System Name of Manufacturer Board (Wr Order No., Minor Year epaed, ASME Code Description of Test eIdentiication Manufacturer SerialNumber Modification,Design Change, Built Stamped Work Conducted Number Number etr.) Replacement Byron Jackson 691-N-0362 N/A WO 99-009059-000 1972 Repaired N/A Repaired Pump System P-7-1B SW interSals Leakage RRU-8 HVAC H. K. Porter M-24442 N/A WO 00-001039-000 1972 Repaired N/A Repaired Leak .nSystem RRU-8HVACinservice N/A - repair RCWHantBCWN/A N/A MM 99-05 1972 ReardNAstructural Structural Hanger made to Plant RCW-H88 RBCCW Fabricated WO 97-008451-020 Repaired N/A Modifications components components only N/A - repair made to MM 99-05 Structural Hanger Plant RCW-H89 RBCCW Fabricated N/A N/A WO 97-008451-020 1972 Repaired N/A Modifications structural components only Fairbanks 38D87001 1TDS N/A MM 2000-001 1972 Replaced Cooling System Morse M12 WO 99-004206-000 WatereBellows inservice Obstacle Removal N/A - repair Obstacl Re I a made to SFP SFCPlant SFPC N/A N/A ~MM MM/99N0641972 99-064 17 Reard NA FuelireRack in Support of Spent structural (Spent Fuel Pool) Fabricated WO 00-000377-000 Fuel Rack cnSpotonpn installation NA/Aonrtrl only

______________

Replaced Stem and WO 00-002526-000 Bonnet, Moved System N/A WO 00-001746-002 1972 Repaired N/A Pakng ToRepair sLea V 13-16 RCIC Walworth SMB-00 WO 00-00 1746-002 Packing To Repair Leakage Leak SR-10-80B RHRSW Consolidated - C31419 N/A WO 00-001935-000 1972 Replaced N/A Replaced Relief System Dresser Valve Leakage TK43762 N/A WO 00-001943-000 1972 Replaced N/A Replaced Relief System SR-10-80A RHRSW Consolidated -

Dresser Valve inservice Page 4 of 14

- m - - - ---- -- m - - - - -- -

m m - m-m m m - - m m m - m -

FORM NIS-2 OWNER'S REPORT FOR REPAIRS OR REPLACEMENTS As required by the provisions of the ASME Code,Section XI, 1986 Edition, No Addenda Vermont Yankee Nuclear Power Plant Unit 1 P.O. Box 157, Vernon, VT, 05354 Construction Code B31.1, 1967 Edition, No Addenda, No Code Case Component Equipment Syidem Name of Manufacturer National Board Other Identification (Work Order No., Minor Year Repaired, Replaced, or ASME Code Description of Test Number Identification Manufacturer Serial Number Modification,Design Change, Built Replaced, Stamped Work Conducted rNumber etc) Replacement Performed Weld System DG-1-1A DG Fairbanks 38D870012 ITDS N/A WO 99-004206-002 1972 Repaired N/A Repair To Eroded inservice Drset Mo02 ARebtave Snste VG-9B CAD Target Rock Model 9 75E002 N/A WO 00-002632-000 1972 Repaired N/A Rebuilt Valve System

______________Leakage Plant Rpae eto f Sse 3"SW-5E SW Fab t N/A N/A WO 00-003663-000 1972 Replacement N/A Replaced Section of System Fabricated 3"SW-5E Piping Leakage Perform Weld Build V70-1A SW Walworth Model # N/A WO 96-012700-000 1972 Repaired N/A Up Valve Body in System 5341 WE Hinge Pin Area Replaced Leakage System TK-3-125-10-19 HCU Liquidonics 200L-8.2-5 N/A WO 00-005412-001 1972 Replaced N/A Replaced System Accumulator Tank Functional General Replaced System HCU eleric P/N 921D59G2 N/A WO 00-005412-000 1972 Replaced N/A Acmlator Tk utin TK-3-125-06-35 Electric Accumulator Tank Functional Fairbanks 38D70006TDS Machined Eroded System DG-1-1B DG Morse M12 N/A WO 00-005379-000 1972 Repaired N/A Area Faces.

On Flange Functional System V70-7A SW Crane Cat. 487 1/2 N/A WO 00-003806-000 1972 Replaced N/A Replaced Valve Leakage P-7-A N/ Rebilt ump System SW Byron Jackson 691-N-0361 N/A WO 00-004971-001 1972 Repaired N/A Rebuilt Pump Leakage P-7-JA SW Walworth Mod. # 5202WE N/A WO 95-005089-000 1972 Replaced N/A Replaced Valve System V70-101

____________ ________Leakage TK-3-125-14-35 HCU Liquidonics 200L-8.2-5 N/A WO 00-005489-000 1972 Replaced N/A Replaced System MMv 2000-0'1 Accumulator Leakage System Mod. # 5275WE N/A MM 2000 1972 Replaced N/A Replaced Valves Leakage V70-1 11A/B SW Walworth SLC Powell N/A N/A WO 00-004934-000 1972 Replaced N/A Replaced Valve System VI 1-12A Page 5 of 14

FORM NIS-2 OWNER'S REPORT FOR REPAIRS OR REPLACEMENTS As required by the provisions of the ASME Code,Section XI, 1986 Edition, No Addenda Vermont Yankee Nuclear Power Plant Unit 1 P.O. Box 157, Vernon, VT, 05354 Construction Code B31.1, 1967 Edition, No Addenda, No Code Case Component National Other Identification Repaired, System Name of Manufacturer Board (Work Order No., Minor Year Replaced, or ASME Code Descriptionof Test Equipment Replacement Stamped Work Conducted Number Identification Manufacturer SerialNumber Number Modification,Design Change, Built RCIC Walworth SMB-00 N/A WO 96-011053-000 1972 Replaced N/A Replaced Valve System V13-15 Leakage CRD BW/IP N/A N/A WO 99-008912-000 1972 Repair N/A Replaced Valve System LCV-3-33D Seats Leakage LCV-3-33C CRD BW/IP N/A N/A WO 99-008911-000 1972 Repair N/A Replaced Valve System Seats Leakage Tested in accordance CC SB 6-97B PCAC Chalmers Allis 00616-11 N/A WO 99-011315-000 1972 Replaced N/A BoltsFlange Replaced with Operations SB-16-19-7B Procedure OP 4202 691-N-0362 N/A WO 00-004971-001 1972 Repaired N/A Rebuilt Pump System P-7-1B SW Byron Jackson Leakage Mod. # Weld Buildup of System V70-1A SW Walworth 5341 WE N/A WO 96-012700-000 1972 Repaired N/A Valve Body in Hinge Pin Area Leakage Fairbanks 38D87001 lTDS Sse DG-1-lA DG N/A WO 99-008333-000 1972 Replaced N/A Replaced Bolting System Morse M12 . ý inservice 38D870012TDS N/A WO 99-009496-000 1972 Replaced N/A Replaced Bolting inservicem DG-1-1A DG Fairbanksse N/A - repair Fairbanks 38D70006TDS elcd upr made to DG-l-lB DG Morse0M12 S N/A WO 99-011236-000 1972 Replaced N/A Replaced Support ructural components only Page 6 of 14 M - M M M M- M M M- M M M M M - n

- -m-m m m-- m m m - - m -

FORM NIS-2 OWNER'S REPORT FOR REPAIRS OR REPLACEMENTS As required by the provisions of the ASME Code,Section XI, 1986 Edition, No Addenda Vermont Yankee Nuclear Power Plant Unit 1 P.O. Box 157, Vernon, VT, 05354 Construction Code B31.1, 1967 Edition, No Addenda, No Code Case Component National Other Identification Repaired, ent EupetBoard ystem Name of Manufacturer (Work Order No., Minor Year Rpaeo ASME Code Descriptionof Test Identification Manufacturer SerialNumber Modification,Design Change, Built R Stamped Work Conducted Number Number etc) Replacement N/A - repair Fairbans 38D7006TDSReplaced Support made to maet DG-1-1B DG Fairbanks Morse 38D70006TDS M 12 N/A WO 99-010257-000 1972 Replaced N/A R lac Clamp p structural components I only DG-I-IB DG Fairbanks Mre12N/A 38D70006TDS WO 99-009229.000 1972 Replaced N/A Replaced Bolting System inevc Morse M12 mnservice Fairbanks 38D70006TDS N/A WO 99-009500-000 1972 Replaced N/A Replaced Bolting n iSystem DG-1-1B DG Morse M12 inservice Liquidonics 200L-8.2-5 N/A WO 00-006190-000 1972 Replaced N/A Replaced System TK-3-125-22-35 HCU 0 6dReplaced System HCU Liquidonics 200L-8.2-5 N/A WO 00-006191-000 1972 Replaced N/A Accumulator Leakage TK-3-125-14-31 Rebuilt Pump 691-N-0361 N/A WO 00-006381-000 1972 Replaced/ N/A Assembly - System P-7-1A SW Byron Jackson Repaired Replaced With Leakage Spare S-3-IB SW R. P. Adams 106047 N/A WO 00-006231-000 1972 Repaired N/A Opened and System Cleaned Strainer Leakage RBCCW thermal Small Bore Piping MM 2000-042 Stress at P-18-lA/B NB Byron Jackson 671-S-1108 N/A MM000 -042 1972 Repaired N/A Modifications To System WO 00-001839-000 Small Bore Piping Leakage Recirc. Pumps I At P-18-1A/B Repaired Relief SR-16-19-77 N2 Kunkle Valve L-3072 N/A WO 00-000596-000 1972 Repaired! N/A Valve - Replaced System Co. Replaced With Spare Leakage Mod. # elcdBont Sse 3003 9 N/A WO 00-006999-000 1972 Replaced N/A Replaced Bonnet system V 11-41 SLC Powell 11 3003 WE IIIStuds and Nuts 1Leakage Page 7 of 14

FORM NIS-2 OWNER'S REPORT FOR REPAIRS OR REPLACEMENTS As required by the provisions of the ASME Code,Section XI, 1986 Edition, No Addenda Vermont Yankee Nuclear Power Plant Unit 1 P.O. Box 157, Vernon, VT, 05354 Construction Code B31.1, 1967 Edition, No Addenda, No Code Case Component System Name of Manufacturer National Other Identification Repaired, Equimen Board (Work OrderNo., Minor Year Replaced, ASME Code Description o Test Identification Equipment Manufacturer Serial Number Board Modification,Design Change, Built Stamped Work Conducted Number Number eta) Replacement RV-10-21A/B RHR s e Mod. # 1685 N/A 00-04 1972 Repaired N/A Removed Valves System Dresser WO 000-0400 Rv- 10-210a/B Leakage Consolidated WO 00-001935-004 SR-10-80 A&B RHRSW TK43762 N/A WO 00-001943-005 1972 Repaired N/A Rebilt Vale Sem Dresser WO 00-007021-000 Relief Valve Leakage General Mod. #ReledCnrl Ssm CRD-06-31 CRD N/A WO 00-004225-002 1972 Replaced N/A Replaced Control System Electric 7RDB 144BG I Rod Drive Leakage General Mod. #Reacdonr! Ssm CRD-06-11 CRD N/A WO 00-004225-003 1972 Replaced N/A Replaced Control System Electric 7RDB144BGI _Rod Drive Leakage General Mod. # Replaced Control System CRD- 14-31 CRD Eeti 7RB4 GI N/A WO 00-004225-004 1972 Replaced N/A lo rv ekg Electric 7RDBI44BGI N/A Rod DriveControl Replaced Leakage System CRD-142-27 CR General Mod. # N/A WO 00-004225-004 1972 Replaced CRD-42-27 CRD General Geetric Mod.

MRdBI# N/A WO 00-004225-005 1972 Replaced N/A Replaced Replaced Control Control System System Electric 7RDBI44BGI _Rod Drive Leakage General Mod. #Reledonrl ysm CRD-18-39 CRD eleric Electric 4 7RDB144BG 1_ N/A WO 00-004225-006 1972 Replaced N/A Replaced Rod DriveControl System Leakage General Mod. # elce otol Sse CRD-26-15 CRD Eeti 7R l4Bl N/A WO 00-004225-007 1972 Replaced N/A RelcdCnol Ssm Electrc____144BG Rod Drive Leakage CRD-34-31 CRD General Mod. # N/A WO 00-004225-008 1972 Replaced N/A Replaced Control System Electric 7RDB144BG1 Rod Drive Leakage General Mod. #ReledCnrl Ssm CRD-34-39 CRD Eeti 7R l4B N/A WO 00-004225-009 1972 Replaced N/A RelcdCnrl Ssm Elcri RD14B1Rod Drive Leakage SLC Union Pump P-C274713 N/A WO 99-009881-000 1972 Replaced N/A Replaced Stuffing System P-45-IA Co. Box Studs and Nuts Functional Union Pump Replaced Cylinder System P-45-IA SLC Co. P-C274713 N/A WO 00-006269-000 1972 Replaced CI N/A Flange land Tie Studs Nuts Functional P-45-IB SLC Union Pump P-C274714 N/A WO 98-011881-000 1972 Replaced N/A Replaced Stuffing System Co. Box Studs and Nuts Functional Page 8 of 14 M M M M M M M M--- M M M = = -

- -- m - - m m - - m-m - - - m - -

FORM NIS-2 OWNER'S REPORT FOR REPAIRS OR REPLACEMENTS As required by the provisions of the ASME Code,Section XI, 1986 Edition, No Addenda Vermont Yankee Nuclear Power Plant Unit 1 P.O. Box 157, Vernon, VT, 05354 Construction Code B31.1, 1967 Edition, No Addenda, No Code Case Component System Name of Manufacturer National Other Identification Repaired, Equipment S Board (Work OrderNo., Minor Year Replaced, or Identification Manufacturer Serial Number Number odfication, Design Change, Built Replacemend Stamped Work Conducted Number Number ~~etc.)Relcmn Union Pump Replaced Cylinder P-45-IB SLC P-45-1 SLC P-C274714 N/A WO 00-006266-000 1972 Replaced N/A Flange and Tie NutsStuds System Co. Functional TK-3-125-18-43 HCU Eentric PIN N/A WO 00-005707-000 1972 Replaced N/A Replaced HCU System Electric 921D595G2 Piston Accumulator Leakage Liquidonics Mod. # N/A WO 00-005708-000 1972 Replaced N/A Replaced HCU System TK-3-125-02-27 HCU 200L-8.2-5 Piston Accumulator Leakage SR 086 RHR H Desr Dresser Mod. #

9352774 N/A WO 97-002364-000 1972 Replaced N/A Replaced Valve System Leakage SR-10-86A HPCI-HD 103FN (Snubber S/N ADH- N/A -

301-1597 removed for Repaired/ Replaced and replacement functional testing and HPCI Anchor Darling ADH-301-1598 N/A WO 00-001027-000 1972 Repaied N/A Repla and replacem returned to stock S/N Replaced Rebuilt Snubber of snubber ADH-301-1598 only installed)

RHR-H 185 (Snubber N/A -

S/N 32198 removed for Fig 200 Repaired/ Replaced and replacement functional testing and rurntionaltotstok RHR Grinnell Fi.N/A WO 00-000995-000 1972 Rpie! N/A Rpae n elcmn returned n stock S/N _S/N 32198 Replaced Rebuilt Snubber of snubber 26351 installed) only RR-3 (Snubber N/A -

S/N 32197 removed for Repaired/ Replaced and replacement functional testing and NB Grinnell Miller Model N/A WO 00-000993-000 1972 Replaced N/A Rebuilt Snubber ofsnubber returned to stock SRN onu 26347 installed) only CS-HD54A (Snubber N/A -

S/N 32195 removed for Repaired/ Replaced and replacement functional testing and CS Miller Fig. 201 N/A WO 00-000991-000 1972 Replaced N/A Rebuilt Snubber ofsnubber returned to stock S/N 26348 installed) I only Page 9 of 14

FORM NIS-2 OWNER'S REPORT FOR REPAIRS OR REPLACEMENTS As required by the provisions of the ASME Code,Section XI, 1986 Edition, No Addenda Vermont Yankee Nuclear Power Plant Unit I P.O. Box 157, Vernon, VT, 05354 Construction Code B31.1, 1967 Edition, No Addenda, No Code Case Component N e MauctrrNational Other Identification Repaired, Nameof Manufacturer Board (Work OrderNo., Minor Year Replaced,'or ASME Code Descriptionof Test Equipment System Identification Manufacturer Serial Number Modification,Design Change, Built Stamped Work Conducted Number Number etc) Replacement RHR-H197A N/A-(Snubber S/N 32196 Repaired removed for functional RHR Lynair Fig. 200 N/A WO 00-000989-000 1972 N/A Replaced and replacement testing and returned to Replaced Rebuilt Snubber of snubber stock S/N,26349 only installed)

RR-35 (Snubber N/A -

S/N 322003 removed Repaired/ Replaced and replacement for functional testing NB Miller Fig. 200 N/A WO 00-000906-000 1972 Replaced N/A Rebuilt Snubber ofsnubber and returned to stock S/N 30034 installed) only MSSRV ~Replaced/Rebuilt Sse MSSRV NB Target Rock 249 N/A PO VY009397 1972 Repaired/ N/A Main Steam Safety System S/N 249 Replaced Relief Valve Leakage MSSRV Repaired/Replaced/Rebuilt Sse S/N 250 NB Target Rock SN50Replaced 250 N/A PO VY009397 1972 Repaired! N/A Main Steam Safet Ma inSeam Safety System Leakage Relief Valve

, Replaced/Rebuilt MSSRV S/N NB Target Rock 67-HH-I-14 N/A PO VY009397 1972 Repaired! N/A Main Steam Safety System 67-HH- 14 Replaced Relief Valve Leakage MSSRV SIN RepairedReplaced/Rebuilt System BL 111134 N NB Target Rock BL 1134 N/A PO VY009397 1972 Repaied Replaced N/A Main Steam Safety ReliekValv Leakage Relief Valve MSSRV S/N Repaired/ Replaced/Rebuilt System BL-1137 NB Target Rock BL- 1137 N/A PO VY009397 1972 Replaced N/A Main Steam Safety Leakage Relief Valve Installed Alternate P0-04 Patent 192 Rea00e0N/A Cooling System To Hydrostatic to CoClndb0 SFPC Fant N/A N/A WO1972 Repaired N/A Standby Fuel Pool and System Fuel Pool Cooling Cooling System leakage System System Design Change Page 10 of 14 M M M-- M M M M M- M M M M M M i

FORM NIS-2 OWNER'S REPORT FOR REPAIRS OR REPLACEMENTS As required by the provisions of the ASME Code,Section XI, 1986 Edition, No Addenda Vermont Yankee Nuclear Power Plant Unit 1 P.O. Box 157, Vernon, VT, 05354 Construction Code B31.1, 1967 Edition, No Addenda, No Code Case Component System Name of Manufacturer National Other Identification Repaired, ASME Code Description of Test Equipment IenicainMnfcue SeilNumber Board (Work Order No., Minor Year Replaced, or Nme Identification Manufacturer Serial Number Number Modification,etc-)

Design Change, Built Replacement Stamped Work Conducted Mod. # 75E001 Valve internals NG-13A CAD Target Rock SMN N/A WO 00-004690-000 1972 Repaired N/A inspection and Bonnet Tack Weld Mod. # 75E001 Valve internals System NG-13B CAD Target Rock SIN 3 N/A WO 00-004691-000 1972 Repaired N/A inspection and Sem Bonnet Tack Weld Leaage Valve internals NG- 11B CAD Target Rock Mod 3 N/A WO 00-004687-000 1972 Repaired N/A inspection and System S/N 3 Bonnet Tack Weld Leakage Mead. # 75E001 Valve internals System NG-12B CAD Target Rock SMN4 N/A WO 00-004689-000 1972 Repaired N/A inspection and Sem Bonnet Tack Weld Leakage Mod. # 75 E001 Valve internals System NG- 12A CAD Target Rock S/N 2 N/A WO 00-004688-000 1972 Repaired N/A inspection and Sem Bonnet Tack Weld Leakage Valve internals NG-I 1A CAD Target Rock Mod. # 75E001 N/A WO 00-004393-000 1972 Repaired N/A inspection and Leakame Bonnet Tack Weld L a P-81D RHRSW Byron Jackson Mod. # VTP N/A WO 01-001098-000 1972 Replaced N/A Replaced Pump System SoN 691-N-0366 Rotating Assembly Leakage DG Fairbanks 38D87001 ITDS N/A WO 01-000804-000 1972 Replaced N/A Replaced the Collar System DG-l-lA Morse M12 Stud Assembly inservice Fairbanks 38D70006TDS Replaced Broken System DG-1-1B DG Morse M T2 N/A WOO1-001 101-000 1972 Replaced N/A OCS Scavenging inservice Air Piping Stud SA Kunkle Valve N/A N/A WO 00-000597-000 1972 Replaced N/A Replaced Relief System SR-72-3A Co. Valve Leakage Mod. #

RV-2-71A NB Target Rock 67F-000-15 N/A WO 00-004226-000 1972 Replaced N/A Replaced Relief System 6X10 Valve Leakage Page 11 of 14

FORM NIS-2 OWNER'S REPORT FOR REPAIRS OR REPLACEMENTS As required by the provisions of the ASME Code,Section XI, 1986 Edition, No Addenda Vermont Yankee Nuclear Power Plant Unit 1 P.O. Box 157, Vernon, VT, 05354 Construction Code B31.1, 1967 Edition, No Addenda, No Code Case Component National Other Identification Repaired, Name of Manufacturer Board (Work Order No., Minor Year Replaced, o ASME Code Description of Test Equipment System Identification Manufacturer SerialNumber Modification, Design Change, Built Stamped Work Conducted Number Number eta) Replacement Mod. Re RV-2-71B NB Target Rock 67F-000-15 N/A WO 00-004720-000 1972 Replaced N/A Replaced Relief System 6X I0 Valve Leakage Mod. # R RV-2-71C NB Target Rock 67F-000-15 N/A WO 00-004721-000 1972 Replaced N/A eplaced Relief System 6X I0 Valve Leakage Mod. #

RV-2-71D NB Target Rock 67F-000-15 N/A WO 00-004722-000 1972 Replaced N/A Replaced Relief System 6X I0 Valve Leakage Mod. # 3707 R SV-2-70A NB Dresser RA-RT21 N/A WO 00-004230-000 1972 Replaced N/A Replaced Safety System S/N BL 1137 Relief Valve Leakage Mod. # 3707 R SV-2-70B NB Dresser RA-RT21 N/A WO 00-004745-000 1972 Replaced N/A eplaced Safety System Relief Valve Leakage S/N BL 1134 V70-71C RBCCW Honeywell Mod. # 8105 N/A WO 00-007152-000 1972 Repaired N/A Repaired Plug and System Stem Leakage Mod. #

V2-80D MS Rockwell 1612JMMY N/A WO 01-001729-000 1972 Repaired N/A Repaired Valve System S/N 123 Seat Leakage Mod. #ReardVle Ssm CS Rockwell 770 # N/A WO 01-001806-000 1972 Repaired N/A Repaired Valve System V14-13A 770 JMMY Internals Leakage Repaired/ Replaced Mod. #tDVSS P-18-IB RBCCW Byron Jackson S/N N/A WO 00-001839-000 1972 Repaired/ N/A Spool on Seal Heat System 671-S-1109 Replaced Exchanger Cooling Leakage III___ Unit Page 12 of 14

- - = = = M - = M =-- = - -

FORM NIS-2 OWNER'S REPORT FOR REPAIRS OR REPLACEMENTS As required by the provisions of the ASME Code,Section XI, 1986 Edition, No Addenda Vermont Yankee Nuclear Power Plant Unit 1 P.O. Box 157, Vernon, VT, 05354 Construction Code B31.1, 1967 Edition, No Addenda, No Code Case Component National OtherIdentification Repaired, Equipment System Name of Manufacturer Board (Work OrderNo., Minor Year Replaced, ASME Code Description of Test Numben Identification Manufacturer SerialNumber N rModification,Design Change, Built ReplaceStamped work Conducted NmeNubretc.) Replacement Tested in accordance PCAC Atwood and Mod. # 20751H N/A WO 00-004108-001 1972 Repaired N/A Installed Disc Nut with V16-19-5G Morril Spacer Shim Operations Procedure OP 4202 V3-114-38-35 HCU N/A N/A WO 01-001886-000 1972 Repaired N/A Repaired Valve System Electric Internals Leakage Mod. #

CV-3-127-38-35 HCU Hammel-Dahl 2500ASA- N/A WO 01-001886-001 1972 Repaired N/A Replaced Teflon System 999Z1204 Seat Ring Disc Leakage Mod. # C44099 V13-131 RCIC Walworth S/N N/A WO 00-006789-000 1972 Repaired N/A Repaired Valve System 530I BSB-WE Internals Leakage Nibco Fig. T-134 N/A WO 01-001934-000 1972 Replaced N/A Replaced Valve System V70-319B SW

_______Leakage_

V70-319D SW Nibco Fig. T-134 N/A WO 01-001935-000 1972 Replaced N/A Replaced Valve V13-6 RCIC Enertech Mod. # DRV-2 N/A WO 01-001732-000 1972 Repaired N/A Replaced Valve System Enertech Mod. # DRV-2 N/A WO 01-001733-000 1972 Repaired N/A Replaced Valve System V13-7 RCIC

___________ ________Spring Leakage N/A WO 01-001740-000 1972 Repaired N/A Replaced Valve System V23-3 HPCI Enertech Mod. f DRV-B Spring Leakage Mod. f DRV-B N/A WO 01-001748-000 1972 Repaired N/A Replaced Valve System V23-4 HPCI Enertech

____ ________ ______ Spring Leakage_

Mod. f Repaired Leak in System RRU-8 HVAC H. K. Porter 41-523-H N/A WO 00-001039-010 1972 Repaired N/A Service Water S/N M-24442 1 Supply Connection Leakage Page 13 of 14

FORM NIS-2 OWNER'S REPORT FOR REPAIRS OR REPLACEMENTS As required by the provisions of the ASME Code,Section XI, 1986 Edition, No Addenda Vermont Yankee Nuclear Power Plant Unit 1 P.O. Box 157, Vernon, VT, 05354 Construction Code B31.1, 1967 Edition, No Addenda, No Code Case Component National Other Identfication Repaired, ASME Code Description of Test Equipment Id~ ubrModification, Manufacturer Board (Work Order Design Change, Year No., Minor Replacedeor Replaced, WrkCoorte Number Identification Manufacturer Serial Number e a Built Replacement Drywell Seat and Mod. # MM 2000-010 Repaired/ Replaced DW Seal gDrywell CB General Electric N/A 1972 N/A and Protective N/A r S Mark I WO 00-001840-000 Replaced Coatig Mak I _____ _______________ _________Coating Repairs _____

Page 14 of 14

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NEC-JH_33 Entergy Nuclear Operations, Inc.

Vermont Yankee OR P.O! Box 0250

ý--.Enteigy

%2=%MFk7 320 Governor Hunt Road Vernon, VT 05354 Tel 802 257 7711 February 5, 2008 BVY 08-008 ATTN: Document Control Desk U.S. Nuclear Regulatory Commission Washington, DC 20555-0001

References:

1), Letter, Entergy to USNRC, "Vermont Yankee Nuclear Power Station, License No. DPR-28, License Renewal Application," BVY 06-009, dated January 25, 2006

2) Letter, Entergy to USNRC, "Update of Aging Management Program Audit Q&A Database," BVY 07-079, dated November 14, 2007

3) Letter, USNRC to Entergy, "Update on Extension of Schedule for the Conduct of Review of the Vermont Yankee Nuclear Power Stationi License Renewal Application," NVY 07-157, dated November 27, 2007

4) Letter, ý;Entergy to USNRC, "License Renewal Application, Amendment 33," BVY 07-082, dated December 11, 2007

5) Letter, , Entergy to USNRC, "License Renewal Application, Amendment 34," BVY 08-002, dated January 30, 2008

Subject:

Vermont Yankee Nuclear Power Station License No. DPR-28 (Docket No. 50-271)

License Renewal Application, Amendment 35 On January 25, 2006, Entergy Nuclear Operations, Inc. and Entergy Nuclear Vermont Yankee, LLC (Entergy) submitted Reference (1), the License Renewal Application (LRA) for the Vermont Yankee Nuclear Power. Station (VYNPS). -

VYNPS submitted Reference (2) following an NRC audit of the VYNPS Aging Management Program and subsequently received Reference (3), which included an NRC Request for Additional Information. References (4) and (5), respectively, provided the initial response to Reference (3) and later clarifications to that response. Additional clarification and details regarding recirculation nozzle Cumulative Usage Factor (CUF) and water chemistry effects are provided in Attachments 1 and 2 to this letter. VYNPS information meeting the NRC's position on Extended Power Uprate (EPU) operating experience evaluation for Aging Management Programs is also discussed below.

VYNPS had not yet entered operation at EPU levels at the time Reference (1) was submitted. EPU power ascension began in March of 2006. To ensure that operating experience at EPU levels is properly addressed by aging management programs, Entergy will perform an evaluation of operating experience at EPU levels prior to the period of extended operation. In addition to VYNPS operating experience, the evaluation will include operating experience from other BWR plants operating at EPU levels.

A117-7

I BVY 08-008.

Docket No. 50-271 Page 2 of 3 This is a new commitment, and has been entered as Commitment #51 on the VYNPS License Renewal Commitment List, Revision 9 (Attachment 3).

Should you have any questions concerning -this submittal, please contact Mr. David Mannai at (802) 451-3304.

I declare under penalty of perjury that the foregoing is true and correct. I Executed on February 5, 2008.

Sincerely, . "I Si ice President Vermont Yankee Nuclear Power Station : Additional Information Regarding Recirculation Nozzle CUF : Additional Information Regarding Water Chemistry Effects : License Renewal Commitment List, Revision 9 cc: Mr. James Dyer, Director.

U.S. Nuclear Regulatory Commission Office O5E7 Washington, DC 20555-00001 Mr. Samuel J. Collins, Regional Administrator, Region 1 U.S. Nuclear Regulatory Commission 475 Allendale:

Road King of Prussia, PA 19406-1415 Mr. Jack Strosnider, Director U.S. Nuclear Regulatory Commission Office T8A23 Washington, DC 20555-00001 Mr. Jonathan Rowley, Senior Project Manager U-S. Nuclear Regulatory Commission 11555 Rockville Pike I

MS-O-11F1 Rockville, MD 20853

BVY 08-008 Docket No. 50-271 Page 3 of 3 Mr. Mike Modes USNRC RI 475 Allendale Road King of Prussia, PA 19406 Mr. James S. Kim, Project Manager U.S. Nuclear Regulatory Commission Mail Stop O-8-C2A Washington, DC 20555 USNRC Resident Inspector Entergy Nuclear Vermont Yankee, LLC P.O. Box 157 Vernon, Vermont 05354 Mr. David O'Brien, Commissioner VT Department of Public Service 112 State Street - Drawer 20 Montpelier, Vermont 05620-2601 Diane Curran, Esq.

Harmon, Curran, Spielberg & Eisenberg, LLP 1726 M Street, N.W., Suite 600 Washington, DC 20036

BVY 08-008 Attachment 1 Vermont Yankee Nuclear Power Station License No. DPR-28 (Docket No. 50-271)

License Renewal Application Amendment 35 Additional Information Regarding Recirculation Nozzle CUF

VERMONT YANKEE NUCLEAR.POWER STATION LICENSE RENEWAL APPLICATION AMENDMENT 35 ATTACHMENT 1 Additional Information Regiarding Recirculation Nozzle CUF NRC Request:

Demonstrate why the confirmatory analysis for the feedwater nozzle bounds the geometry of the recirculation outlet nozzle.

Response

The feedwater nozzle was chosen for the confirmatory analysis since it has the largest number of, and most severe, transients and the highest calculated fatigue usage of the three nozzles which used the VY fatigue analysis approach. The analysis of the feedwater nozzle is bounding for the recirculation outlet nozzle since the calculated usage factors. and thermal transient stresses are significantly less than those for the feedwater nozzle.

As pointed but during the January 8, 2008 presentation to the NRC Staff, the recirculation outlet nozzle has a different geometry (i.e., "skewed") as compared to the other nozzles.

However, the feedwater nozzle configuration remains conservative and bounding when compared to the recirculation. outlet nozzle configuration for the followihg reasons:

The previous comparisons of nozzle corner stress factors from BWRVIP-108, which included evaluation of a recirculation outlet nozzle, demonstrate that the recirculation outlet nozzle configuration does not provide results that are significantly different from the other nozzle configurations.

• The transients experienced by the recirculation outlet nozzle are significantly less severe and less numerous than the transients that affect the feedwater nozzle.

• The most significant thermal transient (improper start causing reverse flow) was modeled directly in the Finite Element Model due to its unique characteristics.

• In-the nozzle corner, the thermal stresses are small compared to the pressure stresses.

• The previous analyses for all three nozzles for.VY yielded significantly lower fatigue usage for the recirculation outlet nozzle compared to the feedwater nozzle.

• Industry experience for the BWR fleet has repeatedly demonstrated that the recirculation outlet nozzle fatigue usage is significantly lower than feedwater nozzle fatigue usage.

BVY 08-008 Docket No. 50-271 Attachment I Page I ofI

-BVY 08-008 Attachment 2.

Vermont Yankee Nuclear Power Station License No. DPR-28 (Docket No. 50-271)

License Renewal Application Amendment 35 Additional Information Regarding Water Chemistry Effects

VERMONT YANKEE NUCLEAR POWER STATION LICENSE RENEWAL APPLICATION AMENDMENT 35 ATTACHMENT 2 Additional Information Regardinq Water Chemistry Effects NRC Request:

Describe how water chemistry effects were accounted for in the evaluation of environmentally assisted fatigue.

Response

Per Section X.M1 of NUREG 1801 (GALL Report) the environmentally assisted fatigue (EAF) evaluations used appropriate Fatigue Life Correction Factors (Fen) calculated using the methodology in NUREG/CR-6583 for carbon and low alloy steels and NUREG/CR-5704 for stainless steels.

For carbon and low alloy steels the Fen factor relationships are shown on page 69 of NUREG/CR-6583. As shown on page 60 of NUREG/CR-6583, the input values used to develop the Fen factors are sulfur content, strain rate, temperature, and dissolved oxygen content in the fluid. Input values for these parameters were chosen to maximize the Fen factors calculated for all components.

The Fen factor relationship for stainless steels is shown on page 31 of NUREG/CR-5704.

As shown on page 25 of. NUREG/CR-5704, the input values used to develop the Fen factors are strain rate, temperature, and dissolved oxygen content in the fluid. Similar to the carbon and low alloy steel calculations, the input values were chosen to maximize the Fen factors.

The inputs were selected as follows:

For the carbon and low alloy steel expressions, the transformed sulfur content parameter was set equal to the maximum value of 0.015 to maximize the effects of this parameter.

For all expressions, the transformed strain rate parameter was set equal to the minimum strain rate (i.e., lessthan 0.001%/sec) for all transients to maximize the effects of this parameter.

For all expressions, the transformed temperature parameter was computed using 550°F for all locations. This temperature envelopes normal operating temperatures to maximize the effects of this parameter, and is very conservative for feedwater temperature.

° For the transformed dissolved oxygen parameter, dissolved oxygen (DO) data was taken from recorded plant data for the feedwater line. For all other locations evaluated in the reactor coolant system, the EPRI BWRVIA code was used to determine DO levels. The EPRI BWRVIAmodel was used to determine DO at component locations at original licensed power (OLP) for both BWR normal water chemistry (NWC) and noble metal water chemistry (NMCA+HWC). Also, current licensed power with NMCA+HWC was evaluated.

BVY 08-008 Docket No. 50-271 Attachment 2 Page I of 2

I VERMONT YANKEE NUCLEAR POWER STATION LICENSE RENEWAL APPLICATION AMENDMENT 35 ATTACHMENT 2 I For the purposes of ensuring that the DO effects on Fe,, are conservative and bounding with respect to water chemistry, the F6 n values used accounted for variations in plant recorded feedwater DO data. It is noted that excursions observed in the plant data used I

are small in number and are of short duration. Approximately 13 years of recorded feedwater DO measurements, including excursions, were evaluated for input to the EAF analysis. A DO value (50 ppb) was used to calculate bounding Fen value for the feedwater I

piping. This represents the mean of the measured data plus one standard deviation.

For locations in the reactor coolant system, the BWRVIA model was run varying the DO content for the power/water chemistry conditions discussed above. The results of these I

sensitivity studies showed that the resulting variations in DO at component locations are significantly less than the changes input to the feedwater DO. The variation of feedwater DO (mean plus one standard deviation) was evaluated. This resulted in less than a 2%

I change in the bounding Fen used in the EAF analysis for the low alloy steel components in the beltline and lower sections of the reactor vessel. There is no effect on the.bounding Fen values from the input feedwater DO variations for the stainless steel components. I The Fen' factors are determined using several parameters and, collectively, these parameters were chosen to conservatively maximize their contribution. The Fen factors are bounding for each location, based on all of the input values. The bounding Fen factors for I

each location and material were used for all stress range pairs in the cumulative usage factor calculations.

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BVY 08-008 Docket No. 50-271 I

Attachment 2 Page 2 of 2 I I

NEC-JH_34 Entergy Nuclear Operations, Inc.

Vermont Yankee

--- Entergy P-0. Box 0250 320 Governor Hunt Road Vernon, VT 05354 Tel 802 257 7711 January 30, 2008 BVY 08-002 ATTN: Document Control Desk U.S. Nuclear Regulatory Commission Washington, DC 20555-0001

References:

1) Letter, Entergy to USNRC, "Vermont Yankee Nuclear Power Station, License No. DPR-28, License Renewal Application," BVY 06-009, dated January 25, 2006.

2) Letter, Entergy to USNRC, "Update of Aging Management Program Audit Q&A Database," BVY 07-079, dated November 14, 2007.

3) Letter, USNRC to Entergy, "Update on Extension of Schedule for the Conduct of Review of the Vermont Yankee Nuclear Power Station License Renewal Application," NVY 07-157, dated November 27, 2007.

4) Letter, Entergy. to USNRC, "License Renewal Application, Amendment 33," BVY 07-082, dated December 11, 2007.

5) Letter, Entergy to USNRC, ý"License Renewal Application, Amendment 31," BVY 07-066, dated September 17, 2007.

Subject:

Vermont Yankee Nuclear Power Station License No. DPR-28 (Docket No. 50-271)

License Renewal Application. Amendment 34 On January 25, 2006, Entergy Nuclear Operations, Inc. and Entergy Nuclear Vermont Yankee, LLC (Entergy) submitted the License Renewal Application (LRA) for the Vermont Yankee Nuclear Power Station (Reference 1).

In Reference (2), Entergy provided an update to the Aging Management Program (AMP)

Audit Q&A Database. In Reference (3), the NRC requested additional information relative to audit question number 387. This information was provided in Reference (4).

Subsequent to that submittal and a follow-up meeting with the NRC staff on January 8, 2008, Entergy agreed to perform additional analyses to support the original response. to this letter provides the results of those analyses. Attachment 2 provides an update to the Cumulative Usage Factor for the Core Spray nozzle forging blend radius that was previously submitted with Reference (5).

This letter contains no new regulatory commitments.

Should you have any questions concerning this submittal, please contact Mr. David Mannai at (802) 451-3304.

j~II BVY 08-002 Docket No. 50-271 Page 2 of 3 I declare under penalty of perjury that.the foregoing is true and correct.

Executed on January 30, 2008. I Sincerely, kfd A/ l<u vI

• Site Vice President Vermont Yankee Nuclear Power Station I

Attachments cc: Mr. James Dyer, Director U.S. Nuclear Regulatory Commission Office 05E7 Washington, DC 20555-00001 Mr. Samuel J. Collins, Regional Administrator, Region 1 U.S. Nuclear Regulatory Commission 475 Allendale Road King of Prussia, PA 19406-1415 Mr. Jack Strosnider, Director I U.S. Nuclear Regulatory Commission "Office T8A23 Washington, DC 20555-00001 Mr. Jonathan Rowley, Senior Project Manager U.S. Nuclear Regulatory Commission 11555 Rockville Pike MS-O-1 1F1 Rockville, MD 20853 /

Mr. Mike Modes USNRC RI 475 Allendale Road King of Prussia, PA 19406 Mr. James S. Kim, Project Manager U.S. Nuclear Regulatory Commission Mail Stop O-8-C2A Washington, DC 20555

'~II I

BVY 08-002 Docket No. 50-271 Page 3 of 3 USNRC Resident Inspector Entergy Nuclear Vermont Yankee, LLC P.O. Box 157 Vernon, Vermont 05354 Mr. David O'Brien, Commissioner VT Department of Public Service 112 State Street - Drawer 20 Montpelier, Vermont 05620-2601 Diane Curran, Esq.

Harmon, Curran, Spielberg & Eisenberg, LLP 172.6 M Street, N.W., Suite 600 Washington, DC 20036

BVY 08-002 I I

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I Attachment Vermont Yankee 1 I License No. Nuclear Power DPR-28 (Docket License Renewal Station No. 50-271) I Application Amendment I

34 RAI 4.3.3-2 Additional Information I I

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VERMONT YANKEE .NUCLEAR POWER STATION LICENSE RENEWAL APPLICATION AMENDMENT 34 ATTACHMENT 1 Vermont Yankee Feedwater Nozzle Confirmatory Analysis Results On January 8, 2008, the Office of Nuclear Reactor Regulation (NRR) staff and Entergy Vermont Yankee (VY) met in a public meeting to discuss VY's response to RAI 4.3.3-2 on environmentally assisted fatigue (EAF). After a formal presentation and dialogue with NRC staff, VY agreed to

-perform a confirmatory EAF analysis on the reactor pressure vessel (RPV) feedwater nozzle. This analysis would confirm the VY fatigue analysis approach by performing an alternate confirmatory analysis using ASME Code,Section III, Subsection NB-3200 [1] methodology to demonstrate available nozzle margins and acceptability of the VY approach. Table 1 provides the results of the confirmatory analysis and demonstrates that the existing VY fatigue analysis approach is acceptable.

Discussion The following items summarize the methods used in the VY confirmatory analysis [2],[3],[4]:

1. The feedwater nozzle was chosen for confirmation since it has the largest number and most complicated and severe transients, and the highest calculated fatigue usage of the three nozzles which used the VY fatigue analysis approach. The analysis of the feedwater nozzle is bounding for the core spray and recirculation outlet nozzles since the calculated usage factors are at least 70% less than those for the feedwater nozzle and the number and severity of thermal transients are less.

2. The confirmatory analysis performed a detailed ASME Code,Section III, Subsection NB-3200

[11 fatigue calculation. The same ANSYS finite element model (FEM) was used as for the current licensing basis fatigue analysis, and was also used in the existing environmental fatigue analysis. The same number and severity of design transients and the same water chemistry inputs were used as had been used in the existing environmental fatigue analysis. Thermal transient stresses were calculated directly using the FEM for all transients.

3. The same transient definitions and cycle counts for 60 years of operation, as defined in Reference [5] and used for the existing analysis [8], were used for computation of cumulative fatigue in the confirmatory analysis.

4. The limiting cross-sections previously evaluated for the feedwater nozzle (nozzle corner and safe end) were evaluated.

5. Primary plus secondary and total stress ranges for all events were calculated and a correction for elastic-plastic analysis (i.e., K,) was applied, where appropriate. Total stress intensity for each transient pair based on stress component differences was calculated per ASME Code,Section III, Paragraph NB- 3216.2 [1]. Stress ranges for primary plus secondary and primary plus secondary plus peak stress were calculated using all six components of stress (3 direct and 3 shear stresses). When more than one load set was defined for either of the event pair loadings, the stress differences were determined for all of the possible loading combinations, and the pair producing the largest alternating total stress intensity (including the effects of Ke) was used.

BVY 08-002 Docket 50-271

VERMONT YANKEE NUCLEAR POWER STATION LICENSE RENEWAL APPLICATION AMENDMENT 34 ATTACHMENT 1

6. For the fatigue usage calculation, stress intensities for the event pairs were re-ordered in order of decreasing primary plus secondary plus peak stress intensity, including a correction for the ratio of modulus of elasticity (E) from the fatigue curve divided by E from the analysis. A fatigue table was created to determine the number of cycles available for each of the events of an event pair, and to determine fatigue usage per ASME Code,Section III, Paragraph NB-3222.4e [1].

For each load set pair in the fatigue 'table, the allowable number of cycles was determined from the alternating stress, which is half of the corrected total stress intensity range, using the appropriate ASME Code, Section II1I [1] fatigue curve.

7. Per Section X.M1 of the GALL Report [61, environmental fatigue multipliers were calculated using the Fen relationships from NUREG/CR-6583 [71 for carbon and low alloy steels. The Fen factors are bounding for all transient pairs based on the highest temperature of each of the transient stress pairs.

The results of the confirmatory analysis and a comparison of the final CUF results from the existing EAF analysis are shown in Table 1 below.

Table 1 - VY Feedwater Nozzle 60 year EAF CUF Location Safe End Analysis EAF Analysis EAF CUF / Allowable 0.2560 /1.0000 I

[8]

Confirmatory Analysis [41 0.0994 /1.0000 I Nozzle Corner (Blend Radius)

EAF Analysis

[8]

Confirmatory 0.6392 /1.0000 0.3531 /1.0000 I

Analysis [4]

I

==

Conclusions:==

The existing EAF analysis for the VY feedwater, recirculation outlet, and core spray nozzles used a I

simplified fatigue analysis approach to calculate CUFs, including bounding Fen relationships. The confirmatory analysis used ASME Code,Section III, Subsection NB [1] methods and included more refined but still conservative Fen relationships.

I For the locations identified above, the EAF results, using either the existing or confirmatory analysis, show that the fatigue usage factors, including environmental effects, are well within I

allowable values for 60 years of operation.

The confirmatory analysis for the feedwater nozzle, which used ASME Section III [1) code methods, confirms the adequacy of the existing VY fatigue analysis approach for all three nozzles.

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BVY 08-002 I

Docket 50-271 I

VERMONT YANKEE NUCLEAR POWER STATION LICENSE RENEWAL APPLICATION AMENDMENT 34 ATTACHMENT 1

References:

1. American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel (B&PV)

Code,Section III, Rules for Construction of Nuclear Power Plant Components, Division 1-Subsection NB, Class 1 Components, 1998 Edition including 2000 Addenda.

2. Structural Integrity Associates Calculation No. VY-19Q-301, Revision 0, "Design Inputs and Methodology for ASME Code Confirmatory Fatigue Usage Analysis of Reactor Feedwater Nozzle".

3. Structural Integrity Associates Calculation No. VY-19Q-302, Revision 0, "ASME Code Confirmatory Fatigue Evaluation of Reactor Feedwater Nozzle".

4. Structural Integrity Associates Calculation No. VY-19Q-303, Revision 0, "Feedwater Nozzle Environmental Fatigue Evaluation".

5. Entergy Design Input Record (DIR) Rev. 1, EC No. 1773, Rev. 0, "Environmental Fatigue Analysis for Vermont Yankee Nuclear Power Station," 7/26/07.

6. NUREG-1801, Revision 1, "Generic Aging Lessons Learned (GALL) Report," U.S. Nuclear Regulatory Commission, September 2005.

7. NUREG/CR-6583 (ANL-97/18), "Effects of LWR Coolant Environments on Fatigue Design Curves of Carbon and Low-Alloy Steels," March 1998.

8. Structural Integrity Associates Calculation No. VY-1 60-302, Revision 0, "Fatigue Analysis

'of Feedwater Nozzle".

BVY 08-002 Docket 50-271

BVY 08-002 Attachment 2 Vermont Yankee Nuclear Power Station License No. DPR-28 (Docket No. 50-271)

License Renewal Application Amendment 34 Update to Core Spray CUF

,If VERMONT YANKEE NUCLEAR POWER STATION LICENSE RENEWAL APPLICATION AMENDMENT 34 ATTACHMENT 2 Update to Supplemental Information for Environmentally Assisted Fatigue Vermont Yankee Nuclear Power Station (VYNPS) provided the following information with Amendment 31 in response to License Renewal Commitment 27. The commitment specified addressing environmentally assisted fatigue by refining fatigue analyses to include the effects of reactor water environment to verify that the cumulative usage factors (CUFs) are less than 1. Entergy completed refinement of the fatigue analyses as specified in Commitment 27 in accordance with the clarifying details provided in the letter of July 30, 2007. The results indicated that the CUFs of the most fatigue sensitive locations will be less than 1.0 through the period of extended operation, considering both mechanical and environmental effects. Subsequent to the Amendment 31 submittal, the environmentally-adjusted CUF value for the Core Spray nozzle forging blend radius was updated to reflect new information, as shown in the revised table below. This table supersedes and replaces in its entirety the table submitted as part of Attachment 1 to BVY 07-066, dated September 17, 2007.

The following results of the refined fatigue analyses are the environmentally adjusted CUF values for 60 years of operation for the locations specified in NUREG/CR-6260.

VYNPS Cumulative Usage Factors for NUREG/CR-6260 Limiting Locations Material Overall*

Environmental Environmentally NUREG-6260 Location Multiplier (Fen) Adjusted CUF 1 RPV vessel shell/ bottom head Low alloy steel 9.51 0.08

• 2 RPV shell at shroud support Low alloy steel 9.51 0.74 3 Feedwater nozzle forging blend radius Low alloy steel 10.05 0.64 4 RR Class 1 piping (return tee) Stainless steel 12.62 0.74 5 RR inlet nozzle forging. Low alloy steel 7.74 0.50 6 RR inlet nozzle safe end Stainless steel 11.64 0.02 7 RR outlet nozzle forging Low alloy steel 7.74 0.08 8 Core spray nozzle forging blend radius' Low alloy steel 10.05 0.0432 0.1668 9 Feedwater piping riser to RPV nozzle Carbon steel 1.74 0.29 Effective multiplier for past and projected operating history, power level, and water chemistry.

BVY 08-002 Docket 50-271

pR3d REG&

UNITED STATES CO0 NEC-JH_35 NUCLEAR REGULATORY COMMISSION WASHINGTON, D.C. 20555-0001 SAFETY EVALUATION BY THE OFFICE OF NUCLEAR REACTOR REGULATION RELATED TO AMENDMENT NO. 229 TO FACILITY OPERATING LICENSE NO. DPR-28 ENTERGY NUCLEAR VERMONT YANKEE, LLC AND ENTERGY NUCLEAR OPERATIONS, INC.

VERMONT YANKEE NUCLEAR POWER STATION DOCKET NO. 50-271 Proprietary information pursuant to Title 10 of the Code of Federal RegulationsSection 2.390 has been redacted from this document.

Redacted information is identified by blank space enclosed within double brackets.

-190-implementation of the proposed EPU. Based on this, the NRC staff concludes that spent fuel storage at VYNPS will continue to meet the requirements of draft GDC-40, 42, and 66 following implementation of the proposed EPU. Therefore, the NRC staff finds the proposed EPU acceptable with respect to spent fuel storage. n 2.8.7 Additional Review Area - Methods Evaluation 2.8.7.1 Application of NRC-approved Analytical Methods and Codes The analyses supporting safe operation at EPU conditions are required to be performed using NRC-approved licensing methodology, analytical methods and codes. In general, the analytical methods and codes are assessed and benchmarked against measurement data, comparisons to actual nuclear plant test data and research reactor measurement data. The validation and benchmarking process provides the means to establish the associated biases and uncertainties. The uncertainties associated with the predicted parameters and the correlations modeling the physical phenomena are accounted for in the analyses. NRC-approved licensing methodology, topical reports and codes specify the applicability ranges. The generic licensing topical reports (LTR) covering specific analytical methods or code systems quantify the accuracy of the methods or the code used. The safety evaluation reports approving topical reports include restrictions that delineate the conditions that warrant specific actions, such as I

obtaining measurement data or obtaining further NRC approval. In general, the use of NRC-approved analytical methods is contingent upon application of these methods and codes withinI the ranges for which the data were provided and against which the methods were evaluated.

Thus, a plant-specific application does not entail review of the NRC-approved analytical methods and codes. 3 To implement the proposed EPU and maintain the current 18-month cycle, a higher number of maximum powered bundles are loaded into the core and the power of the average bundles is also increased, making the core radial power distribution flatter. Due to an increased two-phase pressure drop and higher coolant voiding, the flow in the maximum powered bundles decreases. This effect leads to a higher bundle power-to-flow ratio and higher exit void fraction.

Since the maximum powered bundles set the thermal limits, EPU operation reduces the margins to thermal limits.

Table 2.8.7-1 below shows the predicted operating conditions for the maximum powered bundles for VYNPS as shown in Table 6-2 of Attachment 3 to Reference 25. Figures 2.8.7-1 through 2.8.7-4 show plots for some of these parameters for VYNPS throughout the core cycle.

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Table 2.8.7-1 Ranges of Operational Experience Metric VYNPS Prediction U]

As shown, the VYNPS maximum exit void fraction is 87% and the core average bundle exit void fraction is 76%.

2.8.7.2 Applicability of Neutronic Methods 2.8.7.2.1 Methods Review Topics In Enclosure 3 to a letter dated March 4, 2004, (Reference 69) GE provided its evaluation of the impact of operation at higher void conditions on all of GE's licensing methodologies. The generic evaluation was also based on core thermal-hydraulic conditions that bound the EPU conditions (void fraction 90% or greater). Specifically, operation with a large number of bundles operating at high in-channel void fractions could potentially affect the following topics:

1. Assumptions made in the generation of the lattice physics data that establish the neutronic feedback (see SE Section 2.8.7.2.2).

2. Accuracy of the fuel isotopics generated considering the method employed in the lattice physics (see SE Section 2.8.7.2.2).

3. Assumptions made in the generation of the neutronic parameters in assuming 0%

bypass voiding, although voiding is present during some transients (see SE Section 2.8.7.2.2).

4. Applicability of the thermal-hydraulic correlations used to model physical phenomena (see SE Section 2.8.7.3).

UNITED STATES NUCLEAR REGULATORY COMMISSION.NECJH_6-WASHINGTON, D.C. 20555-0001oNoC 6 2 October 25, 2007 LICENSEE: Entergy Nuclear Operations, Inc.

FACILITY: Vermont Yankee Nuclear Power Station

SUMMARY

OF TELEPHONE CONFERENCE CALL HELD ON AUGUST 20, 2007, BETWEEN THE U.S. NUCLEAR REGULATORY COMMISSION AND ENTERGY NUCLEAR OPERATIONS, INC., CONCERNING THE VERMONT YANKEE NUCLEAR POWER STATION LICENSE RENEWAL APPLICATION The U.S. Nuclear Regulatory Commission (NRC or the staff) and representatives of Entergy Nuclear Operations, Inc. held a telephone conference call on August 20, 2007, to discuss the regulatory requirements stated in 10 CFR Part 54.21 (c)(1) as it relates to the Vermont Yankee Nuclear Power Station license renewal application. provides a listing of the participants and Enclosure 2 contains a summary of the issue discussed with the applicant.

The applicant had an opportunity to comment on this summary.

onathan G. Rowley, Project Manager License Renewal Branch B Division of. License Renewal Office of Nuclear Reactor Regulation Docket No. 50-271

Enclosures:

1. List of Participants

2. Summary of Discussion cc w/encls: See next page

I Vermont Yankee Nuclear Power Station I cc:

Regional Administrator, Region I Ms. Carla A. White, RRPT, CHP I

U. S. Nuclear Regulatory Commission 475 Allendale Road King of Prussia, PA 19406-1415 Radiological Health Vermont Department of Health P.O. Box 70, Drawer #43 I

108 Cherry Street Mr. David R. Lewis Pillsbury, Winthrop, Shaw, Pittman, LLP Burlington, VT 05402-0070 I 2300 N Street, N.W. Mr. David Mannai Washington, DC 20037-1128 Manager, Licensing Entergy Nuclear Operations I Mr. David O'Brien, Commissioner Vermont Yankee Nuclear Power Station Vermont Department of Public Service 112 State Street Montpelier, VT 05620-2601 P.O. Box 500 185 Old Ferry Road Brattleboro, VT 05302-0500 I

Mr. James Volz, Chairman Public Service Board Resident Inspector Vermont Yankee Nuclear Power Station I

State of Vermont U. S. Nuclear Regulatory Commission 112 State Street Montpelier, VT 05620-2701 P.O. Box 176 Vernon, VT 05354 I

Chairman, Board of Selectmen Town. of Vernon Director, Massachusetts Emergency Management Agency U

.P.O. Box 116 ATTN: James Muckerheide Vernon, VT 05354-0116 Operating Experience Coordinator 400 Worcester Road Framingham, MA 01702-5399 I Vermont Yankee Nuclear Power Station 320 Governor Hunt Road Vernon, VT 05354 Jonathan M. Block, Esq.

Main Street P.O. Box 566 I

Putney, VT 05346-0566 G. Dana Bisbee, Esq.

Deputy Attorney General Mr. John F. McCann I

33 Capitol Street Director, Licensing Concord, NH 03301-6937 Entergy Nuclear Operations, Inc.

440 Hamilton Avenue I

Chief, Safety Unit White Plains, NY 10601 Office of the Attorney General One Ashburton Place, 19th Floor Boston, MA 02108 Mr. John T. Herron Sr. Vice President I

Entergy Nuclear Operations, Inc.

1340 Echelon Parkway Jackson, MS 39213 I

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Vermont Yankee Nuclear Power Station cc:

Mr. Christopher Schwartz Mr. Norman L. Rademacher Vice President, Operations Support Director, NSA Entergy Nuclear Operations, Inc. Vermont Yankee Nuclear Power Station 440 Hamilton Avenue P.O. Box 0500 White Plains, NY 10601 185 Old Ferry Road Brattleboro, VT 05302-0500 Mr. Michael J. Colomb Director of Oversight Mr. Raymond Shadis Entergy Nuclear Operations, Inc. New England Coalition 440 Hamilton Avenue Post Office Box 98 White Plains, NY 10601 Edgecomb, ME 04556 Mr. William C. Dennis Mr. James P. Matteau Assistant General Counsel Executive Director Entergy Nuclear Operations, Inc. Windham Regional Commission 440 Hamilton Avenue 139 Main Street, Suite 505 White Plains, NY 10601 Brattleboro, VT 05301 Mr. Theodore Sullivan Mr. William K. Sherman Site Vice President Vermont Department of Public Service Entergy Nuclear Operations, Inc. 112 State Street Vermont Yankee Nuclear Power Station Drawer 20 P.O. Box 500 Montpelier, VT 05620-2601 185 Old Ferry Road Brattleboro, VT 05302-0500 Mr. Michael D. Lyster 5931 Barclay Lane Mr. James H. Sniezek Naples, FL 34110-7306 5486 Nithsdale Drive Salisbury, MD 21801 Diane Curran, Esq.

Harmon, Curran, Spielberg &

Mr. Garrett D. Edwards Eisenberg, L.L.P 814 Waverly Road 1726 M Street, NW, Suite 600 Kennett Square, PA 19348 Washington, DC 20036 Ms. Stacey M. Lousteau Ronald A. Shems, Esq.

Treasury Department Shems, Dunkiel, Kassel & Saunders, PLLC Entergy Services, Inc. 91 College Street 639 Loyola Avenue Burlington, VT 05401 New Orleans, LA 70113

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Vermont Yankee Nuclear Power Station cc: I Karen Tyler, Esq.

Shems, Dunkiel, Kassel & Saunders, PLLC 91 College Street Burlington, VT 05401 Sarah Hofmann, Esq.

Director of Public Advocacy I

Department of Public Service 112 State Street - Drawer 20 Montpelier, VT 05620-2601 Jennifer J. Patterson, Esq. I Office of the New Hampshire Attorney General 33 Capitol Street Concord, NH 03301 3

Matias F. Travieso-Diaz, Esq.

Pillsbury, Winthrop, Shaw, Pittman, LLP i 2300 N Street, NW Washington, DC 20037-1128 Matthew Brock, Esq.

Assistant Attorney General Office of the Massachusetts Attorney I General Environmental Protection Division One Ashburton Place, Room 1813 Boston, MA 02108-1598 Anthony Z. Roisman, Esq.

National Legal Scholars Law Firm 84 East Thetford Road Lyme, NH 03768 Mr. Oscar Limpias Vice President, Engineering Entergy Nuclear Operations, Inc.

1340 Echelon Parkway I

Jackson, MS 39213 I

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TELEPHONE CONFERENCE CALL VERMONT YANKEE NUCLEAR POWER STATION LICENSE RENEWAL APPLICATION LIST OF PARTICIPANTS AUGUST 20, 2007 PARTICIPANTS AFFILIATIONS Jonathan Rowley U.S. Nuclear Regulatory Commission (NRC)

Kenneth Chang NRC Stephen Hoffman NRC Michael Metell Entergy Nuclear Operations, Inc. (Entergy)

Gary Young Entergy Allen Cox Entergy David Lach Entergy David Mannai Entergy Michael Hamer Entergy Brian Ford Entergy Enclosure 1

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OPEN ITEMS VERMONT YANKEE NUCLEAR POWER STATION LICENSE RENEWAL SAFETY EVALUATION REPORT I

AUGUST 20, 2007 I The U.S. Nuclear Regulatory Commission (NRC or the staff) and representatives of Entergy Nuclear Operations, Inc. held a telephone conference call on August 20, 2007, to discuss the regulatory requirements stated in 10 CFR 54.21(c)(1) as it relates to the Vermont Yankee I

Nuclear Power Station (VYNPS) license renewal application (LRA).

Discussion summary: It is the NRC position that in order to meet the requirements of I 10 CFR 54.21 (c)(1), an applicant for license renewal must demonstrate in the LRA that the evaluation of the time-limited aging analyses (TLAA) has been completed. The NRC does not accept a commitment to complete the evaluation of the TLAA prior to entering the period of extended operation.

I Fatigue analyses based on a set of design transients and on the life of the plant are treated as TLAAs. The applicant made a commitment (license renewal. Commitment #27) to address I

environmentally assisted fatigue by refining fatigue analyses to include the effects of reactor water environment to verify that the cumulative usage factors are less than 1.0. The NRC could not accept this commitment.

I Based on the discussion, the applicant agreed to amend its LRA to demonstrate that the evaluation of the TLAA has beehi completed. The NRC's review of this TLAA evaluation will be documented in the final VYNPS safety evaluation report.

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Enclosure 2 I

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