ML20216D902
| ML20216D902 | |
| Person / Time | |
|---|---|
| Issue date: | 08/14/1997 |
| From: | NRC (Affiliation Not Assigned) |
| To: | |
| Shared Package | |
| ML20216D900 | List: |
| References | |
| BWRVIP-05, BWRVIP-5, NUDOCS 9709090456 | |
| Download: ML20216D902 (71) | |
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1 INDEPENDENT ASSESSMENT BY THE OFFICE OF NUCLEAR REACTOR REGULATION l
RELATED TO THE REVIEW 0F THE TOPICAL REPORT BY THE AQ1 LING WATER REACTOR VESSEL AND INTERNALS PROJECT:
"BWR REACTOR PRESSURE VESSEL SHELL WELD INSPECTION RECOMENDATIONS" (BWRVIP-05)
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.s TABLE OF CONTENTS TABLE OF CONTENTS................
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1.0 INTRODUCTION
I 1.1 Background............................
2 1.2 Description of Independent Staff Assessment 3
2.0 REVIEW 0F REACTOR PRESSURE VESSEL FABRICATION TECHNIQUES 4
2.1 Summary of fabrication Methods..................
4 2.1.1 Shielded Metal Arc Welding (SMAW)...............
5 2.1.2 Submerged Arc Welding (SAW)..................
5 2.1.3 Electroslag Welding (ESW)...................
5 2.1.4 Division of BWR Welds into Sub-Populations for Statistical Analysis 6
2.1.4.1 Hanufacturers of BWR RPVs................
6 2.1.4.2 Weld Categorization...................
7 2.2 Statistical Analysis of A Sampling Approach to RPV Weld Examinations 8
3.0 EVALUATION OF NON-DESTRUCTIVE EXAMINATION (NDE)..........
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4.0 RPV DEGRADATION MECHANISMS....................
14 5.0 DISCUS $10N OF LIMITING TRANSIENTS.................
15 6.0 EVALUATION OF PROBABILISTIC FRACTURE HECHANICS MODEL 16 6.1 Structure and Underlying Methodology..............
17 6.2 Technical Issues and Input Parameters 17 6.2.1 Flaw Size Distribution 17 6.2.2 Flaw Density
........................18 6.2.3 Flaw Initiation and Growth Data...............
18 6.2.4 Fracture Toughness
.....................19 6.2.5 Probability of Detection Function..............
20 6.2.6 Embrittlement function and Reference Temperature 20 6.2.7 Other Input Parameters 21 6.3 Importance Sampling 21 6.4 Evaluation of the BWRVIP's Vessel P0F Results 22 7.0 INTEGRATED PROBABILISTIC ASSESSMENT................
23 7.1 Frequency Estimation of Potential Initiating Events and Vessel Integrity Challenges......................
23 7.1.1 Inadvertent injection During Shutdown............
23 7.1.2 Hydrostatic Tests or Vessel Leak Tests
...........25 7.1.3 Loss of Reactor Water Cleanup System
............26 7.1.4 Intentional injection to Support Maintenance Activities...
26 7.1.5 Summary of Transient frequency Estimation..........
26 7.2 Background for Probability of Failure and Uncertainty Analyses.
27 7.2.1 Reforence Case Model s....................
28 7.2.2 Hean and Standard Deviation of Chemistry and Material Data 28 7.2.3 Mean and Standard Deviation of Neutron Fluence 29 7.2.4 Flaw Density and the Flaw Size Distribution.........
29 7.2.5 Residual Stresses and Crack Driving Forces for a Limiting Transient in BWR Vessels
..................31 7.2.6 Uncertairty Distribution for the Conditional Probability of Vessel Failure...........................
32 7.3 Conditional failure Probabilities
...............32 7.4 Uncertainty Analysi s......................
32 7.5 Sensitivity to Flaw Density and the Flaw Size Distribution...
33 7.6 Sensitivity to Inservice Inspection 33 4
t APPENDICES A REFERENCES.............................. A-1 B ACRONYMS AND INITIAllSMS,......................
B-1 C LIMITING TRANSIENTS C-1 TABLES TABLE 2-1 Processes Used to Fabricate Beltline Welds in BWR Reactor Vessels...........................
10 1ABLE 7-1 Three Reference Cases for Staff's BWRVIP-05 Fracture Analysis of RPVs
............................34 TABLE 7-2 Summary of Data from SAFT-UT Inspection of PVRUF Vessel Welds. 35 TABLE 7-3 NRC Staff's Flaw Size Distribution
.............36 TABLE 7-4 Upper 95 percent Confidence Bound Flaw Size Distribution from PVRUF Data
.........................36 TABLE 7-5 Values of Material Property Parameters Used in the BWR RPV Uncertainty Analysis
....................37 TABLE 7-6 Conditional Probability of Failure P(FlE) for CE Fabricated BWR Vessels...........................
38 TABLE 7-7 Conditional Probability of Failure P(FlE) for B&W Fabricated BWR Vessels...........................
39 TABLE 7-B Conditional Probability of failure P(F Vessels................IE) for CB&I Fabricated BWR
...........40 TABLE 7-9 Results of Uncertainity Analyses 41 TABLE 7-10 Results of Sensitivity Analyses...............
42 TABLE 7-11 NRC Staff's flaw Size Distribution With Adjustment for Inser.' ice Inspection
.........................43 TABLE 7-12 Results of Sensitivity Anklyses due to ISI
.........44 TABLE C-1 Sampling Review of LERs and ens............... C-7 FIGURES Figure 3-1 PISC I detection results for defect detection probability versus defect depth using 1974 ASME Section XI Code type procedures at sensitivity of 50% DAC 14 Figure 7-1 Probability of Detection and Correct Sentencing Curves from PISC-II Studies....,..................
45 Figure 7-2 Effect of Inservice Inspections with Differeing POD Curves on Flaw Depth Distribution
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'y INDEPENDENT ASSESSMENT BY THE OFFICE OF NUCLEAR REACTOR REGULATION RELATED TO THE REVIEW OF THE TOPICAL REPORT BY THE B0lllNG WATER REACTOR VESSEL AND INTERNALS PROJECT:
"BWR REACTOR PRESSURE VESSEL SHELL WELD INSPECTION RECOMMENDATION 1" (BWRVIP-05)
1.0 INTRODUCTION
By letter dated September 2B, 1995, as supplemented by letters dated June 24 and October 29, 1996, and May 16, June 4, and June 13, 1997 (References 1, 2, 3, 4, 5, and 6, respectively), the Boiling Water Reactor Vessel and Internals Project (BWRVIP), a technical committee of the BWR Owners Group (BWROG),
submitted the proprietary report, "BWR Vessel and Internals Project, BWR Reactor Pressure Vessel Shell Weld Inspection Recommendations (BWRVIP-05),"
whicn proposed to reduce the scope of inspection of the BWR reactor pressure vessel (RPV)he axial welds and zero percent of the circumferential welds. welds from percent of t Reference 2 provided supplemental information in response to two Nuclear Regulatory Commission (NRC) staff requests for additional information (RAls) dated April 2 and May 20, 1996 (References 7 and 8).
In Reference 3, the BWRVIP modified the BWRVIP-05 original proposal to increase the examination of the axial welds to 100 percent (from 50 percent) while still 3roposing to inspect essentially zero percent of the circumferential RPV siell welds (exce)t that the intersections of the axial and circumferential welds would have seen included; approximately 2 - 3 percent of the circumferential welds wauld be inspected under this proposal).
Reference 4 provided additional supplemental information in response to a third RAI dated May 20, 1997 (Reference 9).
References 5 and 6 provided the BWRVIP's VIPER computer code and detailed programming information for the VIPER code; the VIPER code was used by the BWRVIP in obtaining the results detailed in the BWR\\'IP-05 report.
The staff met with members of the BWRVIP and their consultants on several occasions to discuss issues related to the review of the BWRVIP-05 report.
These meetings.were summarized in the following meeting summaries: summary of July 18, 1995, meeting dated July 25, 1995; summary of March 19, 1996, meeting dated March 26, 1996; summary of October 15, 1996, meeting dated December 10, 1996; and, summary of January 16, 1997, meeting dated February 13, 1997 (References 10, 11, 12, and 13, respectively). Additionally, by letter dated April 18, 1997 (Reference 14), the BWRVIP requested to meet with the Commission on this issue. On May 12, 1997, the Commission was briefed by i
representatives of the BWRVIP and the staff on the issues related to the requirements for a full inspection of reactor pressure vessel shell welds.
The above meeting summaries, as well as the transcript of the May 12, 1997, Commission meeting (Reference 15) and the Commission's May 30, 1997, Staff Requirements Memorandum (SRM M970512B, Reference 16) are available in the Commission's Public Document Room, 2120 L Street, N.W., Washington, D.C.
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1.1 Background
In January 1991, the NRC published in the federal Register (56 FR 3796) a proposed Rule to amend Section 50.55a to Title 10 of the Code of rederal Regulations (10 CFR 50.55a), " Codes and Standards." One purpose of this amendment was to incorporate by reference a later edition and addenda to Section XI of the American Society of Mechanical Engineers (ASME) Code.
This included the 1989 Edition of the ASME Section XI, Division 1, and addenda through 1988.
In addition, the Rule proposed to create Section 50.55a(g)(6)(ii)(A) to 10 CFR 50.55a (10 CFR 50.55a(g)(6)(ii)(A)), " Augmented examination of reactor vessel," which required that all licensees perform an inspection of RPV welds in accordance with Scction XI of the ASME Code on an expedited schedule, and revoked all previously granted reliefs for RPV weld examinations.
As noted in the Statement of Considerations for the Rule, the primary reason for the augmented examination requirement was that few examinations had been performed up to that time on either boiling water reactor (BWR) or on pressurized water reactor (PWR) RPV shell welds.
Therefore, there was a concern on the part of the staff regarding the existence of manufacturing flaws in the RPV shell welds, and the initiation and propagation of flaws during service.
The staff had no assurance that ASME Section XI flaw acceptance criteria were being satisfied in the RPV shell welds, which had never been completely examined as part of a Section XI inservice inspection program. The augmented examination was needed for.BWRs because of evidence demonstrating a viable mechanism for initiating environmentally-assisted cracks in the RPV cladding, evidence that the cladding cracks can propagate into the ferritic steel of the RPV base material and, evidence that BWR reactor vessels may be embrittled more by neutron irradiation than would be predicted by Regulatory Guide 1.99, Revision 2, " Radiation Embrittlement of Reactor Vassel Materials."
The intent of 10 CFR 50.55a(g)(6)(ii)(A) was to require that licensees perform an expanded RPV shell weld examination, as specified in the 1989 Edition of ASME Section XI, on an "ex> edited" basis.
" Expedited," in this context, effectively meant during tie inspection interval when the Rule was approved or the first period of the next inspection interval.
The final Rule was published in the Federal Register on August 6, 1992 (57 FR 34666).
By incorporating into the regulations the 1989 Edition of the ASME Code, the staff was requiring that licensees perform volumetric examinations of "essentir.11y 100 percent" of the RPV pressare-retaining shell welds during all inspection intervals.
This represented an expansion by the ASME of the requirements from previous editions of Section XI of the ASME Code which, as far back as the Winter 1975 Addenda, had required examination of 100 percent of the RPV pressure-retaining nelds during the first inspection interval, then limited examinations in the intervals thereafter.
Requiring every licensee to perform an extensive volumetric examination of every RPV shell weld at least once during the service life of the RPV was therefore consistent with the philosophy that had been expressed in the ASME Code for more than 20 years, in those 20 years, however, BWR licensees had requested and been granted extensive relief from performing RPV shell weld examinations.
As a result, only a small percentage of the beltline welds in BWR RPVs had been examined,.
and no BWR licensee had completed an examination which would satisfy the philosophy and requirements of the NRC and the ASME Code.
Recognizing the small percentage of RPV welds that were being examined, the conflict between this small percentage of examinations and the ASME Code requirements, and the fact that inspection technology had evolved such that commercial systems were available to support the ASME-specified scope of RPV inspections, the previously granted reliefs were revoked in 1992 with the issuance of the Rule.
1.2 Description of Independent Staff Assessment The following provides an independent staff assessment of the BWRVIP-05 report.
The Commission, in SRM M970512B, requested that the staff consider a tiered approach in gathering additional baseline information and/or implementing the Rule (10 CFR 50.55a(g)(6)(ii)(A The SRM recommended that thestaff'sassessment(a)shouldaddresstheBWNIPproposaltoexamine100 percent of the axial welds which would include examinations of some circumferential weld lengths near the intersections of the weld types to determine if this proposal could provide an appropriate level of sampling'of the RPV welds, (b) should provide a comprehensive evaluation of the l
probabilistic analysis contained in the BWRVIP aroposed alternative in determining the acceptability of a proposed tecinical alternative and/or in pursuing changes to the rule, and, (c) should receive appropriate review, including review by the Advisory Committee on Reactor Safeguards (ACRS).
The staff has initiated a broader, risk-informed review of the BWRVIP-05 proposal since the May 12, 1997, Commission briefing.
One result of this effort was the identification of a transient at a foreign BWR of U.S. design in which the RPV was subjected to high pressure (7.9 MPa or 1150 psig) at a low temperature (26*C to 31'c or 79'F to B8'F).
This cold overpressure I
transient is not included as a design basis event for BWRs and was not considered in the BWRVIP-05 report which was focused only on design basis events.
However, the recent recognition of this transient has led the staff to determine that cold overpressure transients are of sufficient safety significance to be considered in this assessment.
Additionally, the staff has performed a preliminary review of 229 licensee event reports and 81 event notifications which involved potential BWR overcooling or overpressure events since 1980 (see Appendix C). Of the 310 events identified, 35 were identified as potential precursors to cold overpressure events of the type that occurred overseas.
These types of events are of particular interest because the fracture toughness of the RPV decreases at low temperatures resulting in greater potential for RPV failure.
Preliminary evaluations of the foreign event indicate conditional failure probabilities for axial and circumferential welds significantly higher than l
those associated with the transients assumed in the BWRVIP-05 report.
These preliminary staff evaluations indicate that the conditional failure probabilities for the circumferential welds, instead of being approximately 30 orders of magnitude less than the axial welds' conditional failure probabilities magnitude lower (as stated in the BWRVIP-05 report), are about four orders of Further work is being performed to more fully assess the risk associated with these events for both the axial and circumferential welds at fluence levels projected to be res;aed later in life at some plants.
This additional work includes further studies of the potential precursor events in order to better quantify the potential for cold overpressure events in BWRs, and additional probabilistic fracture mechanics analysis to both understand the sensitivities to various parameters and to support an uncertainty analysis.
This assessment provides, in Section 2.0, a description of the techniques used by the several vendors to fabricate the reactor vessels.
Section 3.0 contains a discussion of the history of non-destructive examinations, and Section 4.0 has a discussion of the two degradation mechanisms which have the potential to initiate RPV cracking or to cause existing flaws to grow. -In Section 5.0 and Appendix C there is a discussion of limiting transients of concern.
The probabilistic fracture mechanics model is discussed in Section 6.0, and an integrated probabilistic assessment is presented in Section 7.0.
Section 8.0 contains a summary of foreign inspection strategies and inspection capabilities, and the staff's interim conclusions are presented in Section 9.0.
A final Safety Evaluation Report (SER) will be completed and provide additional evaluations of non-design basis events. A final conclusion regarding the need for circumferential weld examinations, or other actions that may be suggested by the risk informed assessment, will be reached when the final SER is completed.
2.0 REVIEW 0F REACTOR PRESSURE VESSEL FABRICATION TECHNIQUES Reactor pressure vessels were fabricated to very high standards, as evidenced by preservice and inservice inspections to date. However, different processes were used to fabricate RPV welds under a number of varying conditions.
Additionally, several vendors constructed BWR RPVs.
The staff reviewed the various fabrication methods associated with RPV fabrication in order to support a statistical assessment of a sampling approach to RPV weld inspections, in particular, the staff focused on whether it could be assumed that, considering the various fabrication methods, that all RPV welds represent one statistical population, or if the population of RPV welds could be divided into well defined subpopulations.
Section 2.1 discusses the review of welding techniques and differences in the types of welds (Reference 17).
Section 2.2 discusses the statistical evaluation of a sampling approach to inspection.
2.1 Summary of Fabrication Methods A number or variables had to be considered when determining which weld process to use for fabrication of an RPV. One variable was economic impact.
During the late 1960's and early 1970's when RPVs were fabricated, shielded metal arc welding (SMAW) was the most expensive of the three weld processes.
Submerged arc welding (SAW) was less expensive than the SMAW process.
Electroslag welding (ESW) was the least expensive process of the three, however, it could only be used to fabricate axial weids. The weld processes used in fabricating BWRs are described below.
Relative strengths and weaknesses for each weld process are also outlined.
2.1.1 Shielded Metal Arc Welding (SMAW)
SMAW is a process wherein coalescence is produced by heating with an electric arc between a coated electrode and the plate or forging.
SMAW is a manual process in which the electrode is fed into the weld joint by the welder.
Shielding is obtained from decomposition of the electrode coating and the filler metal is obtained from the electrode. Using the SMAW process, weld metal can be deposited in any orientation.
Therefore, this process was used for some field-fabricated beltline welds.
Welding is interrupted frequently to change the electrode which results in a slower rate of deposition and smaller weld beads.
More than 100 passes could be needed to complete an SMAW weld.
Since the process is manual, there is a greater probability of weld defects.
These defects are expected to be relatively smaller than defects associated with other weld processes.
2.1.2 Submerged Arc Welding (SAW)
SAW is a process wherein coalescence is produced by heating with an electric arc or arcs between a bare metal electrode or electrodes and the plate or forging.
SAW is an automatic process in which the electrode is a spool of wire that is continuously fed into the weld joint at a constant rate.
The welding is shleided by a blanket of granular, fusible material that is on the plate or forging, and the filler material is obtained from the electrode or electrodes.
The granular material, or molten flux, that submerges the arc, serves to protect the weld from oxidation.
Since SAW is an automatic process, it is less subject to operator error and would be less likely to produce discontinuities such as cracks, entrapped slag or porosity.
However, competent welders should be able to produce quality SMAW welds that are equivalent to these produced using the SAW process.
The number of passes needed to complete an SAW weld could be 50 to 100.
2.1.3 Electroslag Welding (ESW)
ESW is a process that is unique in comparison to the other weld processes since the entire weld thickness is deposited in one pass.
ESW is an automatic process used to fabricate axial welds (not applicable to circumfirential welds).
An arc is present throughout the process, and the==" ed weld f1"x formsamoltenslagwhichprovidestheheatneededtomelttheveldrt'.
The slag floats above the molten weld metal and protects tu v..dment Iren, atmospheric contamination. One or more weld wires can be fed into the f.J.
so high weld deposition rates can be achieved. The molten weld metal and slag are contained in the weld region by water-cooled cop)er shoes that are braced against the plates being joined. The structure of tie ESW weld is as-cast throughout and must be heat treated to achieve the required properties.
The ESW process closely resembles a continuous casting operation.
Due to the one-pass nature of the ESW process, the probability of producing a defect is small, but any defects would be larger than those found in welds that are fabricated using the other weld processes.
It should be noted that there are welding variables such as heat and contamination that affect the quality of the resulting weld.
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2.1.4 Division of BWR Welds into Sub-Populations for Statistical Analysis it is important to note that each reactor vessel can have unique characteristics that distinguish it from other reactor vessels, even if the vessels were fabricated by the same vendor.
For example, differences in the fit up of the vessel shell courses or ' roundness" of the vessel can exist among vessels that were fabricated by the same vendor.
Accurately categorizing sub-populations of reactor vessel welds is difficult due to the many variables involved in the welding process. However the staff attempted to divide the welds according to available weld process Information.
2.1.4.1 Manufacturers of BWR RPVs Table 2-1 shows the BWR welds grouped by vessel fabricator and weld process.
The axial welds that were fabricated by Chicago Bridge and Iron (CB&I) are double V-groove welds made with a spacer bar.
The SAW process was used from the shell inner diameter (ID).
The spacer bar was removed by back gouging from the outer diameter (00), then the SMAW process was used to complete l
fabrication of the axial weld.
The circumferential welds that were fabricated by CB&I are typically single V-groove welds with a closed root gap.
in most cases, the SAW process was used f rom the shell OD.
Back gouging was used to remove the root pass in order to l
reach the sound metal. The SMAW process was used from the shell ID to complete fabrication of the circumferential weld..
The axial welds that were fabricated by Combustion Engineering (CE) are double U-groove welds with a closed root gap geometry.
The SAW process was used from both the shell ID and OD. After one side was completed, back gouging was used to reach the sound weld metal. The SMAW rocess was used as needed to shape the back gouged cavity.
The SAW process was used to complete fabrication of r
the axial weld.
The circumferential welds that were fabricated by CE are square groove welds with a carbon steel backing strip. The SAW process was used from the shell 00.
The backing strip was removed and back gouging was used to reach the sound weld metal.
The SMAW process was used from the shell ID to complete fabrication of the circumferential weld.
The ESW process was used for axial welds that were fabricated by Babcock &
Wilcox (B&W).
TheexceptionisgeBigRockPointaxialweldswhichwere fabricated using the SAW process The SAW process was used from the shell 10. The spacer bar was. removed by back gouging from the shell 0D.
The SAW process was used from the shell OD to complete fabrication of the axial weld.
The circumferential welds that were fabricated by B&W are a single V-groove welds with a backing strip. The SAW process was used from the shell OD.
The backing strip was removed and back gouging was used to reach the sound weld Big Rock Point is scheduled to be permanently shut down by Fall 1997, and is not considered in this assessment.
metal.
The SMAW process was used from the shell ID to complete fabrication of j
the circumferential weld.
Dresden unit 2 was the only BWR fabricated by New York Shipbuilding (NYS).
The ESW process was used to fabricate the axial welds, and the SAW process was used to fabricate the circumferential welds.
The NYS fabricated vessel is comparable to the B&W vessels.
Hope Creek was the only BWR fabricated by Hitachi.
The SMAW process was used to fabricate the axial welds, the circumferential welds, and a low pressure 2
coolant injection (LPCI) nozzle weld that is in the beltline.
The Hitachi fabricated vessel is comparable to the CB&I vessels.
1 2.1.4.2 Weld Categorization i
Sources of variance to consider when categorizing welds include: 1) whether the welds were field or shop fabricated, 2) details of weld repair procedures,
- 3) whether a single or double groove process was used to fabricate the weld,
- 4) details of back gouging procedures, 5) details of cladding procedures, and
- 6) stress relief processes.
l The BWRVIP reviewed plant-specific documents and reported that the following j
CB&l vessels were field completed:
1 Clinton Duane Arnold Limerick 1 & 2 Monticello Susquehanna 1 & 2 Vermont Yankee All other BWR vessels were shop fabricated.
Details on which CB&I welds were field or shop fabricated are not readily available.
This level of detail would require a plant-specific search. With regard to weld repair, locations of the repairs would vary; but all major repairs were required to be documented as part of the ASME Code procedures.
With regard to back chipping and weld refill, double and single groove welds with a backing strip or spacer bar require grinding, chipping, machining and/or air-arc gouging to remove root defects.
Lack of fusion and porosity were common defects that occurred because production of a fully fused root was a
not an objective of the weld procedure.
Back gouging the root to sound weld metal was.an important part of the weld procedure.
When air-arc gouging is used, an additional 0.03125 inches (0.79375 mm) of weld metal was required to be removed by mechanical means. This was required to assure that the arc-heat 3
hardened surface was removed. After the metal is removed, the weld groove is restored to the original groove dimensions. Weld repair may be required if j
the weld groove is not acceptable for continued welding.
The weld groove is surface inspected to assure removal of defects before welding.
With regard to back cladding, the axial and circumferential welds were back i
clad using SMAW, SAW, or the semi-automatic gas metal arc welding (GMAW), or some combination of weld processes.
The plate material and axial welds of individual shell courses were machine clad as one step.
The vessel assembly drawing usually specified more than one process.
In general, there are a minimum of two layers of back clad. The first layer is E309 type electrode and the remaining layers are E308L.
E309 and E308 type electrodes were used
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in older vessels.
For B&W fabricated vessels, the SMAW, SAW and GMAW processes were used for back cladding.
For CE and CB&I fabricated vessels, the SMAW process or a combination of the SMAW and GMAW processes were used for i
back cladding.
With regard to stress relief processes, the standard temperature and time used for stress relief was ll50'F (621'C) for a cumulative period of approximately 35 hours4.050926e-4 days <br />0.00972 hours <br />5.787037e-5 weeks <br />1.33175e-5 months <br />.
The guidelines for the amount of stress relief time are given by Section III of the ASME Code.
A shell course may have gone through multiple stress f elief cycles as a sub-assembly before experiencing another process cycle during treatment of the circumferential weld.
It should be noted that the ASME Code contains alternate requirements for stress relief temperatures that are lower than the minimum of 1100'F (593'C).
2.2 Statistical Analysis of A Sampling Approach to RPV Weld Examinations Of the total length of axial and circumferential welds in a BWR RPV, approximately 60 percent are circumferential welds and axial welds make up the remaining 40 percent.
The staff has concluded that no meaningful statement about the circumferential welds can be made based only on a statistical analysis of the axial welds. Any such statement must be based on the assumption that both the axial and circumferential welds are random samples from some parent population of welds.
If this assumption is made, an inference can be made about the expected number of defective circumferential welds based on an inspection of the axial welds.
For the best possible situation in which all axial welds are inspected and no defective welds are found, an inference can be made about the expected number of defective circumferential welds based on the inspection of the axial welds:
Denote the expected number of defective > circumferential welds by d, c
the proportion of defective welds in the parent population by p, the number of circumferential welds by a and the number of axial welds by n.
Assuming that circumferential welds-comprise 60 percent and axial welds 40 percent of the total weld length, then m = 1.5n and d - ap.
Further, with the standard confidence level of 95 percent, it can be concluded that p < 3/n.
Therefore, it follows that, with 95 percent confidence, d < 4.5.
Based on this bound for d, a bound on the probability that at least one of the circumferential welds is defective can be derived.
The distribution of the 1
number of defective circumferential welds is given by a Poisson distribution with mean d.
Therefore, the probability that at least one circumferential weld is defective is equal to 1 - exp(d).
It follows that this probability is less than 1 - exp(-4.5) = 0.99, with 95 percent confidence.
Based on the above evaluation, the staff has concluded that no useful inference regarding the condition of the circumferential welds can be made based on a 100 percent inspection of the axial welds only.. The staff also considered the effect of including a small percentage of the circumferential (8)
Note that for the purpose of this evaluation, " defective" is defined as containing-a crack-like indication that could severely jeopardize RPV integrity. -
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welds to in the inspection sample by assuming a sample size of 50 percent of total RPV welds (axial and circumferential).
Including the small percentage of the circumferential welds does not alter the above conclusion, i
The above statistical analysis is based on the assumption that both the axial and circumferential welds are random samples from the same parent population
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of welds.
In reality, as discussed in Section 2.1, the circumferential welds are not from the same population due to the different weld processes and other sources of variances (i.e., field vs. shop fabrication); therefore, a sampling approach is not valid.
An example of the above is the results of the Tennessee Valley Authority's Browns Ferry Nuclear Plant Unit 3 RPV shell welds augmented examination (Reference 17).
The augmented examination consisted of 15 axial welds and 5 circumferential welds.
The axial welds were ESW and the circumferential welds were SAW.
The axial welds contained no reportable indications.
circumferential welds contained over 100 recordable indications (3)The
, 15 of i
which required evaluation to be determined acceptable by ASME Code Section XI, IWB-3600 (1986 Edition).
Since all the recordable flaws were in the SAW type welds, this examination indicates that the SAW type welds have a greater propensity for flaws than the ESW type welds.
Hence, examination results from welds.',, axial welds are not representative of the SAW circumferential the ESW The above evaluation of a statistical sampling approach to RPV weld examinations indicates that relatively large sam)1e sizes e.g., on the order of 50%, do not allow a meaningful conclusion wit 1 regard to the possibility of a single or a small number of isolated unacceptable defects existing in the remainder of the population.
The practical implication of this evaluation is that a sampling approach cannot provide reliable protection against the possibility of an isolated, unacceptable conditions existing in the population of uninspected welds. An example of this situation would be a large defect existing in a substandard weld repair. This is an important conclusion relative to assessment of the RPV which must have an extremely high reliability. However, this evaluation should not be interpreted as meaning that no useful information can be gathered from inspection of a sample of welds.
For example, a sampling inspection can provide useful inferences regarding the possible existence of a more wide-spread mechanistic mode of degradation. An example of this might be a service induced cracking mechanism. Thus it is possible that additional insights may be available through a statistical assessment of the sample of welds that has been or will be inspected in BWRs under the VIP-05 proposal.
As indicated in Section 2.1,
" Reportable" indications are those that are required to be reported to l
(3) the NRC; " recordable" indications are those of a sufficient size to be recorded by the testing equipment.
(')
During a meeting on August 8, 1997, the staff was informed that the axial welds at Hatch did non-reportable indications and that the inspection methods used for the axial welds may have been different from those used for the circumferential welds.
The staff will need additional details in order to assess this new information. -
care must be taken in any such evaluation to make sure that differences in 1
sub-populations of welds are taken into account, and it may be difficult to define such sub-populations.
The staff intends to perform further evaluations and to pursue this area in greater detail with the industry.
TABLE 2-1 Processes Used to Fabricate Beltline Welds in BWR Reactor Vessels
PROCESS USED TO FABRICATE-----
PLANT FABRICATOR AXIAL WELDS CIRCUMFERENTIAL WELDS BIG ROCK POINT B&W SAW SAW BROWNS FERRY UNITS 1/2/3 B&W ESW SAW
^
DRESDEN UNIT 3 B&W ESW SAW PEACH BOTTOM UNITS 2/3 B&W ESW SAW QUAD CITIES UNITS 1/2 B&W ESW SAW BRUNSWICK UNITS 1/2 CB&I SAW SAW CLINTON CB&!
SAW SMAW DUANE ARNOLD CB&I SMAW SMAW i
GRAND GULF CB&l SMAW,
SAW LASALLE UNIT 2 CB&I SAW SAW LIMERICK UNIT 1 CB&I SMAW SAW MONTICELLO CB&I SMAW SMAW NINE MILE POINT UNIT 2 CB&I SAW SAW PERRY CB&l SMAW SAW RIVER BEND CB&I SMAW SMAW SUSQUEHANNA UNITS 1/2 CB&I SAW SMAW VERMONT YANKEE CB&I SMAW SMAW WNP UNIT 2 CB&I SAW SAW LIMERICK UNIT 2 CB&I SMAW SMAW COOPER CE SAW SAW FERMI UNIT 2 CE SAW SAW FITZPATRICK CE SAW SAW HATCH UNITS 1/2 CE SAW SAW LASALLE UNIT 1 CE SAW SAW 4
MILLSTONE UNIT 1 CE SAW SAW NINE MILE POINT UNIT 1 CE SAW SAW OYSTER CREEK CE SAW SAW PILGRIM CE SAW SAW _
TABLE 2-1 i
Processes Used to Fabricate Beltline Welds in BWR Reactor Vessels
PROCESS USED TO FABRICATE-----
PLANT FABRICATOR AXIAL WELDS CIRCUMFERENTIAL WELDS DRESDEN UNIT 2 NYS ESW SAW HOPE CREEK HITACHI SMAW SMAW*
g
~Also SMAW for a nozzle weld 3.0 EVALUATIONOFNON-DESTRUCTIVEEXAMINATION(NDE)
Reactor pressure vessels were constructed to ASME Section !!! requirements, which include the requirements for NDE fabrication inspection. Many reactor pressure vessels were constructed in the U.S. from the mid-1960s through the mid-1970s. To evaluate the effectiveness of the NDE performed during this time, several relevant studies of inspection reliability should be assessed.
The only significant study conducted in the 1960s was the work of the Pressure Vessel Research Committee (PVRC) of the Welding Research Council.
in the late 1960s, PVRC Specimen 201 (CB&l vessel 8" thick, butt weld, manual metal arc process) was inspected, using several techniques and procedures. This specimen contained a total of 10 flaws (3 cracks, 5 slags, and 2 lack-of-fusion flaws).
The effectiveness of the inspection procedure was computed as the ratio (in percent) of the number of flaws detected plus the number of false indications to the total number of flaws present. The average performance of the radiographic testing (RT) teams was only 26 percent, and ranged from a high of 33 percent to a low of 17 percent.
The performance of the ultrasonic testing (UT) teams from the unciad surface was about 50 percent, with a high of 65 percent and a low of 43 percent.
For UT from the clad surface, the UT average was 14 percent, with a high of 43 percent and a low of 0 percent (3 teams made na sfetections at all).
With advanced 1960s NDE technology, such as acoustic holography and focused probes, UT inspections from the clad surface achieved an average of 65 percent, with a high of 70 percent and a low of 60 percent.
This work was followed by inspections of PVRC Specimen 251J (CE vessel, butt weld, with submerged arc weld process), through the mid-1970s.
This plate contained 10 cracks and 5 slags. The RT results were similar to those described above, with an average of 21 percent (a high of 27 percent and a low of 15 percent).
The UT inspections of Specimen 251J produced similar performance to that for Specimen 201 and even included a specially developed PVRC procedures intended to address some problem areas identified in the earlier PVRC work.
The next significant reliability study was a round robin conducted under the international Plate Inspection Steering Committee (PISC) program (References 19 and 20).
The focus of PISC was on UT, and the PISC procedure was based on the ASME 1974 Code, using a recording threshold of 20 percent DAC (distance-amplitude-correction) and a reporting threshold of 50 percent DAC. Any UT signal exceeding the recording threshold must t'e recorded and those signals --
d exceeding the reporting threshold must be fully documented, dispositioned and 1
reported. The defect detection probability exceeded 90 percent for vertical
]
defects 50 mm (1.97 inches) defects up to 150 mm (5.91 inches) in through-w and larger in depth.
However, for sets of defects it was only 40 percent for
]
size.
The PISC studies continued with a second phase in the early to mid-1980s (Reference 21).
This PISC 11 database is the largest ever collected on the capabilities of NDE procedures and techniques for thick section steel.
The results for PISC Plate No. 3 show that UT procedures based on a 50 percent DAC sensitivity achieve on average, a 15 percent probability of detection and correct sentencing,g) for defects up to 60 mm (2.36 inches) in through-wall size. To put this in perspective, this is the sensitivity level specified in the ASME Code up until the 1986 edition; and the changes introduced with the 1986 edition were based on these PISC 11 results.
When the sensitivity was increased to 10 or 20 percent DAC, the effectiveness of NDE increased, on average, to 75 percent, for defects u) to 60 mm (2.36 1
inches) in through-wall size. This PISC II result is the
) asis for the ASME Code requirement going to a sensitivity of 20 percent DAC.
The PISC II studies also showed that effective performance for the near surface zone can be achieved only with special NDE techniques; and that a 70* dual longitudinal transducer is an effective tool for inspections of the near surface zone j
through the vessel cladding.
In summary, these PVRC and PISC studies showed that RT and UT inspections conducted through the mid-1980s, in accordance with ASME Code minimum requirements, were not very effective.
Inspections based on the procedures, equipment, and personnel requirements specified in ASME Section XI Code, i
Appendix VIII, are expected to perform significantly better than those based i
on ASME Code prior to Appendix VIII.
It is difficult to state quantitatively how much better, since Appendix VIII describes a screening test, and does not I
quantify the probability of detection (P00). The design basis for reactor pressure vessel testing specified in Appendix VIII is that an inspection result that combines a true value of 70 percent POD with a false call probability of 20 pei ant should have only a few percent chance of passing the performance demonstration test. This is a significantly higher performance than that demonstrated in the earlier studies cited above.
The industry is using highly effective performance plots of POD versus flaw size for input to the VIPER program (Fig. 2-11 of Reference 5).
This i
performance is significantly better than those reported in the PVRC and PISC studies.(including the results for special procedures).
Furthermore, the industry plots are the composite performance results for several separate and independent inspection techniques.
Up through the 1986 edition, the ASME Code called for vessel examinations using two inspection angles, each larger than l
30*, and separated by at least 15*.
Most inspections were conducted at 45*
t and 60*, including the PVRC and PISC studies.
For inspections based on the (5) -
Correct sentencing means that a defect is detected and is also correctly
~
sized as an acceptable defect.
If the defect is rejectable, it must be sized as rejectable, based on ASME Section XI, IWB-3500.,
3 A
1986 and later editions of the ASME Code, performance would be expected to conform to the PISC 11 results for 20 percent DAC. This is substantially less effective than the performance the BWRVIP is using for thA 'lIPER program.
The POD curves (Figure 2-11 of Reference 5) for use in thc. v!PER program are based on the assumptions that the techniques are complimentary and independent.
Regarding the NDE procedures industry is using in the performance demonstration tests, the staff does not know if these are complementary or independent techniques.
Furthermore, the PISC and other round robin studies have shown that the largest contributor to variability in POD performance is the human factor. These uncertainties are not reflected in the POD curves that industry is using (Figure 2-11 of Reference 5). The results shown there should be regarded as an upper bound on performance, and are not typical of inspections conducted using normal inservice inspection (ISI) techniques.
As an example, the PISC-II results on Plate No. 3 (shown in Figure 3-1), for a through-wall flaw size of 50 mm (1.97 inches) and a sensitivity of 20 percent DAC, the average performance is 75 percent with a variability frem 50 percent to 100 percent probability of detection and correct sizing. To represent inspection performance more realistically, results should be shown separately for the early (pre-1986 Code) inspections with a ser.sitivity of 50 percent DAC, those with a sensitivity of 20 percent DAC, and those based on special procedures.
The latter would be a good first approximation for those successfully passing the demonstration tests specified in Appendix VIII of the 4
ASME Code.
In summary, the POD curves industry is using in the VIPF9 program are considered to be an optimistic upper bound on POD perform nce.
It should be noted that assuming a high POD maximizes the benefit of inspection and is conservative for evaluating the ability of inspection to reduce the probability of RPV failures. However, such an assumption may not be conservative when assessing other inspection issues.
Furthermore, the staff's position is that the reference case assessment should be performed using best estimate (i.e., realistic) assumptions. More realistic P0D curves, based on round robin data (shown in Figure 3-1), should be used in the VIPER calculations.
The proper curve should be selected in Figure 3-1, depending on the era of the NDE testing being simulated.
l l
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+ Am, nase mien a th.c.,unsemian twiw e.mac l
0 set.: es==
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/ / /
j 8
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,, a#
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a to ao in too 3m a: s ca.o Defect site :.dimenstonn of the defect in the through wall section Thickness of,the plates a from 200 mm to 250 mm Figure 1 PIsc I detection results for defect detection probability versus defect depth using 1974 ASME Section XI Code type proceduros at sensitivity of 50% DAc (Reference 10).
b 4.0 RPV DEGRADATION MECHANISMS l
Based on approximately 700 reactor-years of U.S. operating experience of BWR vessels (approximately 1700 reactor-years world-wide), the BWRVIP identified that fatigue and stress corrosion cracking (SCC) are the two degradation c
mechanisms which have the potential to initiate RPV cracking or to cause existing flaws to grow. There are two fatigue mechanisms significant to the i
BWR RPV: system cycling fatigue and rapid cycling fatigue.
The system cycling fatigue, associated with plant start-up, shutdown, SCRAM and safety relief valve (SRV) blowdown, had been evaluated for the vessel shell as part of the vessel design and, because the resulting usage factors were around 0.1, found by the BWRVIP to be insignificant. The rapid cycling fatigue associated with plant operation when feedwater sparger and nozzle cracking was initiated was discounted because the vessel shell is remote from these nozzles.
Based on the above, the BWRVIP narrowed its investigation of the degradation mechanisms to only SCC, which was suggested by RPV cracking experience (e.g, cracks found in some clad and unclad steam generator shells of PWR vessels and in some recirculation nozzles of BWR vessels).
Cracking observed in the manually backclad region also suggested the involvement of SCC.
i >
The staff reviewed the qualitative arguments in the BWRVIP-05 report supporting the elimination of fatigue as a significant degradation mechanism, and determined that additional analysis was required.
As a result, the BWRVIP provided (Reference 2) a quantitative comparison of the crack growths due to these two mechanis The low alloy steel (LAS) SCC growth rate relation of da/dt-1.18X10'gs gK and the ASME Section XI reference bi-linear fatigue crack growth law for carbon and low alloy steels in water environment were used in this study.
The crack growth history for both mechanisms were presented for crack depths of 1.27, 2.54, and 7.62 cm (0.5, 1.0, and 3.0 inches).
The results indicate that, when the cracks were increased into the range of larger sizes of the Marshall distribution (e.g., 7.62 cm or 3.0 inches), the crack growth from stress corrosion is more dominant. This difference is even more pronounced if the staff's LAS SCC growth rate, which is approximately 10 times larger than the BWRVIP's value, is used in this comparison.
Hence, the staff has concluded that the fatigue crack growth does not need to be considered in this application.
5.0 DISCUSSION OF LIMITING TRANSIENTS The BWRVIP-05 report discusses the operating characteristics of a BWR with respect to design transients and their effects on the vessel.
These transients generally occur when a large steam region exists (i.e., a steam bubble is present). According to the BWRVIP-05 report, the most limiting operational transients with respect to the vessel are loss of feedwater or single safety relief valve (SRV) blowdown events for normal and upset conditions.
Normal operating temperature and pressure for a BWR RPV is 260*C (500*F) and 6.9 HPa (1000 psig).
During the hydrostatic test, RPV temperature and pressure is approximately 65.5*C to 93.3*C (150*F to 200*F) and 6.9 MPa (1000 psig) and is maintained on the pressure-temperature (P-T) curve for that particular vessel.
The most limiting transients for emergency and faulted operating conditions are any transient which causes or results in a rapid cooldown, and rapid depressurization, of the vessel shell welds. These transients are limiting for pre-existing cracks in the vessel shell welds.
In comparison, the BWRVIP-05 report states that the water-solid leak test condition, or hydrostatic test, is limiting for small flaws in RPV 10.
To assess the probability of failure (P0F) for rapid depressurization, the staff assumed a hypothetical transient and performed a parametric study using the VISA-II code.
The transient was characterized by a rapid depressurization to 10 psi (69 kPa) with the fluid temperature decreasing from iM*F to 206*F (288'C to 97'C) in 12 seconds (i.e., equivalent to a cooling rate of 1720*F or 956*C per minute). Use of this essentially instantaneous cooldown transient was meant to bound all real transients. The staff examined both axial and circumferential welds in this stug.
In one case, the staff intentionally increased the fluence to 5.0 X 10 n/cm2 and the initial RT for the axial weldto40*FtoincreasetheprobabilityofproducingvesseYo' failure.
However, the staff's results indicated no failure for all cases subjected to rapid depressurization transient.
As a result, the severe foreign transient discussed in Appendix C.2.1 was used in the staff's independent simulations reported in Section 7.
The BWRVIP-05 report was limited to design basis accident (DBA) events.
To provide a broader risk-informed assessment, the staff performed a sampling review of approximately 17 years of licensee event reports (LERs) and event notifications (ens) to determine if other events (i.e., shutdown events) could be potentially more limiting to the vessel (see Appendix C).
The staff's sampling review does not encompass all LERs and ens, which will be more fully reviewed in the staff's final SER. The staff also notes that some overpressure events during shutdown, as described in Appendix C.1, may not be reportable under 10 CFR 50.72.
The preliminary staff sampling review is discussed in Appendix C.2.
Table C-1 includes the results of the staff's sampling review of LERs and ens.
The table provides the maximum pressure and temperature reached at the end of the transient, where known.
Where the information was not provided, an NP is listed.
It should be noted that, although the transients discussed in Appendix C are examples of the types of events that could result in cold overpressure events, these events occurred at temperatures that are high relative to the reference temperature of the vessel weld material, as required by plant Technical Specifications (TS).
Therefore, these events did not, in and of themselves, represent significant challenges to the RPV.
This is more thoroughly discussed in Appendix C.
The staff's review also identified an actual low-temperature overpressure event that occurred at a foreign plant. A more detailed discussion of the staff's independent assessment of the potential for beyond design basis challenges to the RPV is presented in Section 7.1.
6.0 EVALUATION OF PROBABILISTIC FRACTURE HECHANICS MODEL Conventional vessel analysis codes, such.as VISA-II and FAVOR (References 22 and 23), are codes based on probabilistic fracture mechanics (PFM) methodology that perform millions of deterministic vessel simulations to determine the conditional P0F for a vessel subjected to a specific transient.
The vessel conditional P0F is the ratio of the number of failed vessels to the number of simulations.
For each simulation, the random variables (e.g., crack size, copper, nickel, and fluence) are assigned according to the prescribed distribution with the form and parameters of the distributions specified by the user.
Deterministic fracture mechanics analyses are then performed, and the vessel conditional P0F is determined.
The BWRVIP-05 report used the VIPER code (Reference 5) to perform the PFM analyses.
Section 6.1 describes the underlying methodology and the special features in the VIPER code.
Section 6.2 evaluates technical issues and input parameters. All major technical issues related to the VIPER code have been reviewed by the staff, with special attention to the flaw size distribution and the flaw density in vessel welds.
New information in these two areas has been accumulated in the past 5 years. The staff has used the new information to thoroughly address the projected flaw size distribution and density.
The staff's assessment of the VIPER's importance sampling technique is in Section 6.3 and the staff's assessment of the BWRVIP's P0F results is presented in Section 6.4.. - -
4 6.1 Structure and Underlying Methodology The structure and the underlying methodology in VIPER is similar to those of VISA-II and FAVOR.
In these type of codes, parameters which are not highly 4
variable (e.g., vessel and clad geometry, transient pressure and temperature definition, weld volume and orientation) are treated as deterministic.- Highly variable parameters (e.g, clad stress, fluence, copper and nickel contents, initial RT
, and fracture toughness) are treated as random variables.
The VIPER codeliffers from VISA-II or FAVOR by having implemented the following i
additional features to its core PFM methodology: (1) an importance sampling technique was added to the conventional Monte-Carlo simulation to significantly reduce the number of simulations needed to obtain convergent results, (2) a cladding SCC initiation curve was added to simulate crack growth in cladding, and (3) an SCC growth rate for.LAS was added to simulate further crack growth from cladding into base metal or shell welds.
The VIPER code assigned initiated and manufacturing flaws and sizes in accordance with the appropriate distributions.
Crack initiation in cladding
{
was first determined by comparing the service time with the predicted time for cladding crack initiation under a certain cladding stress according to the cladding SCC initiation law. Once initiation occurs, the flaw is assumed to crack through the clad instantly and line up with a manufacturing flaw in the LAS. Material properties (e.g., fluence, copper and nickel contents, RTm) are sampled for each flaw, and the flaws are grown according to the probabilistic crack growth law.
Inspections are performed for the inspection intervals and percentage of weld s)ecified, and the process continued until i
either the flaw grows to failure ()ased on fracture mechanics analysis), or until the end of vessel life is reached. The program continues to loop through the-analysis in this manner, until the specified number of iterations have been completed.
Like the VISA-II code, VIPER can define POD curves and treat the percent inspection of the vessel as an input parameter.
Detailed discussions regarding these three areas and the POD curves used in the VIPER are discussed in the following sections, along with some other basic technical issues and input parameters.
1 6.2 Technical Issues and Input Parameters 6.2.1 Flaw Size Distribution For cases without preservice insmection (PSI), the VIPER code-used the following function to describe tie flaw size distribution:
N(x) = (6.94)exp(-6.94x).......................
(1) where x is the crack length in inches and N(x) has been normalized. The staff used the following flaw size distribution derived from the pressure vessel research user facility (PVRUF) data (Reference 24). The distribution, in the cumulative distribution form, is:
i F(.0787) =.9054 and, F(x) = 1 - (.1616)exp(-6.94x) - (.00139)exp(-4.06x) for x >.0787 (2)
The distribution used in the BWRVIP-05 report and in the staff's assessment were based on the Marshall flaw size distribution.
The Marshall flaw size distribution (Reference 25), after normalization, is (page 41, Reference 22):
N(x) = (0.0346)exp(-4.06x) + (6.88)exp(-6.94x)............
(3) with ultrasonic inspection, and N(x) - (4.06)exp(-4.06x).......................
(4) without ultrasonic inspection. Although Marshall did not mention (Reference
- 25) whether the inspection was preservice or inservice, the staff interprets
" ultrasonic inspection" as PSI.
The VIPER manual (Reference 5) indicates that the Marshall distribution vith inspectfon (i.e., Equation 3), was used for the as-manufactured (wfthout PSI) flaw size distribution.
If this were the case, the BWRVIP would have double counted the benefit of inspection for the case when inspection was considered and an associated POD curve was utilized.
This is because Equation (3) has already considered the benefit of PSI inspection.
However, the staff found through examining the source coding of VIPER that instead of Equation (3), the VIPER : 1e employed Equation (1) in its PFM calculations for the as-manufacu. red flaw size distribution.
The BWRVIP did not provide a basis for using Equation (1).
However, for large flaw sizes, Equation (1) is a good approximation to Equation (3).
The staff's flaw size distribution was calibrated against the PVRUF data and follows the Marshall distribution in Equation (3) for x > 0.0787 inch.
Section 7.2.4 contains the derivation of the staff's flaw distribution.
Since the staff's flaw size distribution has a probability of 91 percent for flaws under 0.0787 inch, and for larger flaws, the exponent in the major term of the staff's equation is the same as that in the BWRVIP's equation, the staff's l
flaw distribution may give lower vessel conditional P0Fs.
i 6.2.2 Flaw Density The range for flaw density as stated in the Marshall report is 0.4 to 40 flaws per cubic meter, with the most likely value at the low end of this range.
The BWRVIP used a flaw density of 30 flaws per cubic meter.
Recent information (Reference 24) on flaw density indicates that the flaw density in RPV welds may be more than 30 times higher than that used by the BWRVIP.
Therefore, the staff used a flaw density of 995 flaws per cubic meter in its assessment, which was based on detailed examinations on the Midland vessel and the PVRUF vessel and is discussed in Section 7.2.4.
6.2.3 Flaw Initiation and Growth Data The VIPER code assumed that after an SCC initiated in the clad, it would penetrate the clad instantly and line up with underlying fabrication defects.
The time to initiate the crack in the clad was developed by the BWRVIP and is discussed below. The staff assumed that service would induce SCC cracks in the BWR vessel cladding, but did not consider the initiating time of the cladding cracks. The staff then adopted the same assumption as in VIPER code _ _ _ _ __
4 that the SCC would penetrate through the clad and combine with the fabrication flaws.
The VIPER code used a cladding SCC initiation curve that was derived from five
" cast /weldment" data points.
The VIPER code did not consider data pertaining to sensitized 304 stainless steels because BWRVIP believes that cast /weldment materials should have far greater resistance to SCC than sensitized 304 stainless steels.
However, Figure 8-4 of the BWRVIP-05 report indicates that two of the five data points for cast /weldment material have the same initiation rate as that for sensitized 304 stainless steels in the BWR environment.
Consequently, the BWRVIP's approach of forming a separate group for these five cast /weldment data points and using their best-fit curve in its PFM analyses may not be prudent. The BWRVIP later performed additional PFM analyses in its response (Reference 2) to the staff's RAls (References 7 and
- 8) using the conservative SCC initiation curve for sensitized 304 stainless steel. These results indicated that the P0Fs became an order of magnitude higher because the crack initiated earlier in the cladding.
However, the general trend regarding the effect of percent inspection on P0Fs remained unchanged.
Nevertheless, when absolute values of P0F become important and no additional cast /weldment data are available to show a definitive trend, the more conservative cladding SCC initiation curve should be used.
The BWRVIP developed its LAS SCC growth curve from low and high oxygen environments in Reference 26 and recent General Electric (GE) data from References 27 and 28. The recent GE data shows a lower crack growth in high oxygen, which falls reasonably in line with the old low cxygen data, and virtually no crack growth in a low oxygen environment.
To take advantage of the lower crack growth rate exhibited by the new high oxygen data while discarding the zero crack growth exhibited by the new low oxygen data, the BWRVIP used an engineering approach to derive the growth rate for the low oxygen environment of BWR vessels.
The BWRVIP found the ratio of the low to high oxygen crack growth rates from the old data, then applied this ratio to a fitted curve from the new high oxygen data. The BWRVIP's approach of deriving the LAS SCC curve for low oxygen environment gives heavy weight to the new GE data and to the ratio of the high oxygen growth to low oxygen growth rates from the old data.
The staff considers the proposed LAS SCC growth curve not acceptable because there is no evidence that the new data is more accurate than the old data.
Further, the test time for the GE new low oxygen data is only a few hundred hours, which is relatively short for such tests.
Reference 29 showed that crack growth is virtually zero for hundreds of hours, and yet
-there was a quite significant average crack growth rate over much longer periods of time.
6.2.4 Fracture Toughness The VIPER code has four options for the fracture toughness curve:- ASME K3 ASME K, the mean K * (VISA K,), and the mean K, (VISA K,).
The ASME K,,and i
i 3
i the ASkE K,, curves kn the Viper code are the lower bound curves in Sectio,n XI and the mean K curves are from NRC analyses of the ASME code. The mean K,d in the VISA-/I code.
i (Reference 30) and are also use According to Reference 2, the BWRVIP used the mean K, curve without sampling in all analyses i
The mean K curve is truncated at 200 ksi perfgr,medintheBWRVIP-05 report.Although VIPER treats the upper-shelf' fracture i
(in) truncated K value) as a random variable, this feature was deactivated in the i
BWRVIP's ana* lyses.
Consequently, there was no sampling on the upper-shelf fracture toughness either.
The staff used the FAVOR code to perform the vessel simulations defined in Section 7.
Unlike VIPER, FAVOR employs a mean K, curve and samples about this curve.
The mean K, in FAVOR is different i
i from the mean K in VIPER and VISA-l!.
FAVOR's mean K was derived from the i
i bounding K, curv,e in the ASME Code in which the boundin,g curve was considered i
to be a mean-minus-two standard deviation curve.
The BWRVIP's approach of using a mean K without sampling is not appropriate.
g Since the limiting transient that the BWRVIP evaluated occurs on the upper shelf, the BWRVIP essentially eliminated one of the key random variables from their PFH analyses.
No sampling about the fracture toughness in the transition and lower-shelf regions also limits the application of the VIPER cede.
For instance, for the PFM analysis presented in Section 7, the foreign transient occurs below the upper-shelf region.
Analyses using the mean K i curve without sampling would give lower vessel P0Fs, and therefore, is non,-
conservative.
The staff believes that sampling about the mean K, is i
nccessary to produce accurate P0F, The fracture toughness is calculated as a function of T-RT where T is the temperature of the metal and RT,3, isthereferencetempera1og,ure of the vessel weld material.
The RT, is the sum of the initial RT,i'n term to account for,
, the increase in RT resulting from neutron Trradiation (ART,or)d nickel. contents and calculatio uo
, and a marg uncertainty in the initial RT,Se, presented in Section 6.2.6.
copper an procedures.
The details will 6.2.5 Probability of Detection Function The VIPER manual indicates that four inspection methods can be specified by the user: eddy current (EDDY), full-V ultrasonic angle beam examination (FULLV), multi-mode ultrasonic examination (SLIC), and combined techniques (COMBINED), for both PSI and ISI.
The BWRVIP's corresponding POD curves are shown in Figure 2-11 of Reference 5.
The BWRVIP performed sensitivity studies on the effect of lower POD on the P0F. The results summarized in Figure 8-14 of the BWRVIP-05 report shows that, for high P0D curves, the P0F for 50 percent inspection is about 7 times the P0F for 90 percent inspection.
For a.
low POD curve, the P0F for 50 percent inspection is only 2 times the P0F for 90 percent inspection. This is the basis that the BWRVIP used to apply only high POD curve to real vessel cases.
Because if results from a P00 curve that reflects high benefit of ISI can justify a relaxation of ISI requirements, the results from a POD curve that reflects less benefit of ISI can certainly justify the relaxation.
The staff agrees that a lower P00 would show a reduced benefit of ISI.
However, as discussed previously, a realistic P00 should be used when performing a probabilistic risk assessment.
Assumptions regarding more or less optimistic PODS are useful for performing uncertainty or sensitivity studies.
6.2.6 Embrittlement Function and Reference Temperature Embrittlement is measured as an increase in the reference temperature resulting from neutron radiation, ART and is a function of copper and nickel content and fluence.
IntheVNERandFAVORcodesthecopperand
nickel contents and the neutron fluence are random variables, and the mean MT ' ion 2 (Reference 31).is calculated according to the procedure of Regulatory Guid 1.99, RevIs The RT of the embrittled vessel material sum of the initial RT themeanMTuoi, and a term to account for uncertaintyintheinIt'ialRT copper and nickel contents, and calculational procedures. A1DIo, ugh in theory the uncertainty in MTwo, has artially accounted for by treating the initial RT and copper and been nicke contents as random variables, VIPER conservative 9 adopted the same approach as FAVOR by specifying a standard deviation for MT,of. The square of this standard deviation for MT is added to the square o the standard deviation for the initial RT,o,.
TE term is calculated as the square root of this sum.
6.2.7 Other Input Parameters The BWRVIP reported its summary of results for real vessel cases in Table 8-9 of the BWRVIP-05 report. The results for axial welds are based on a " typical" BWR vessel with an initial RT, of 10*F, a copper content of 0.30 percent, a nickel content of 0.3g percenI, a room temperature clad stress of 32 ksi, and a fluence of 4.9 x 10 n/cmt. The corresponding values for circumferential welds are -32*F for the initial RT,or, 0.06 percent for the copper content, and 0.97 percent for the nickel content. The mean fluence and the mean room l
temperature clad stress are the same for both welds. The standard deviation values for both welds are 17'F for the initial RT,oy, 0.042 percent for the i
copper content, 0.05 percent for the nickel content, 5 ksi for the room temperature clad stress, and 20 percent for the fluence.
l Although the effect due to high variability in copper and nickel contents has been explored by the BWRVIP through making additional calculations using standard deviations of copper (0.045 percent. 0.075 percent) and nickel (0.0165 percent 0.075 percent), the effort was not extensive enough to determine the uncertainty associated with the reported P0F results.
The staff surveyed the reactor vessel integrity database (RVID) and found that there are apppximately 258 BWR RPV welds that have fluence 10 times larger than 4.9 X 10 n/cmt, and have a copper content and an initial RT,okaff does for the circumferential weld much higher than those of the BWRVIP's.
The s not believe that one reference case for axial welds and one reference case for circumferential welds can be used to represent all BWR reactor vessel welds.
The staff surveyed the RVID database, and grouped the vessels according to fabrication processes for different vendors.
Based on the difference in materials used in the fabrication of the CE, B&W, and CB&I vessels, the staff determined that three reference cases are necessary to comprehensively evaluate all BWR reactor vessels.
The reference cases and the results from the PFM analysis of the reference cases and the associated uncertainty analyses are reported in Section 7.
Other inputs, such as the mean (32 ksi) and standard deviation (5 ksi) of the clad stress, are considered adequate by the staff.
6.3 Importance Sampling Importance sampling was used to shift the random variables of the Monte-Carlo simulation in the unfavorable direction to reach a convergent P0F solution l 9
with many fewer simulations.
Each vessel failure was adjusted by multiplying the ratio of the probability density function for that random variable in the shifted distribution versus that in the unshifted distribution.
For importance sampling with multiple variables such as the current case, the probability ratios are multiplied or summed, depending on how the associated variables occur in the 3roblem. The vessel P0F was obtained by first summing the final ratio for eac1 failed vessel until the specified number of simulations has been performed, then dividing this sum by the specified number of sinlations.
Examples showing the convergence of P0F with respect to the number of iterations were provided in the BWRVIP-05 report for sample single-variable problems.
In Reference 2, the BWRVIP provided similar convergence information for sample multi-variable problems, which are more pertinent to the current application.
A detailed parametric evaluation of sampling strategies and shifting parameters was also provided there.
However, the independent simulations performed by the staff using VIPER indicated the contrary.
Some other concerns with the importance sampling scheme used in VIPER have also been identified. When the distributions for the number of manufacturing flaws and SCC flaws have importance sampling performed, the ratios of the density functions corresponding to manufacturing flaws and SCC flaws are multiplied.
Since either an SCC flaw or a manufacturing flaw may fail in a given vessel simulation, it appears inappropriate to multiply these two ratios.
It would appear that it is more appropriate to use only the ratio for the type of flaw which fails (i.e., manufacturing or SCC).
Another concern involves the importance sampling weights when a very low P0F is being estimated.
In such a case, the observed value of the shifted distribution will tend to lie very far out in the tail of the unshifted distribution, so that the ratio of these density functions will tend to be very large. Unless there are a large number of simulated failures, there is a danger that the sum of the ratios (i.e., the importance sampling weights) will be dominated by a small number of very large ratios.
In such a case, the sum of the ratios and hence the estimated P0F would have a large variance.
- Thus, if the P0F being estimated is very low, the importance sampling estimate may have a large variance.
Hence, the staff concludes that the accuracy and repeatability of the VIPER code is yet to be demonstrated. The staff plans to conduct additional reviews in this area.
6.4 Evaluation of the BWRVIP's Vessel P0F Results The BWRVIP benchmarked the VIPER code against the VISA-II code using some test problems, and summarized its findings in Figure 8-12 of the BWRVIP-05 report.
The staff concludes that the results from using VIPER are in reasonable agreement with those from using the VISA-II code when the same inputs are used. However, as discussed in Sections 6.2.1 to 6.2.7, the staff does not believe the inputs used by the BWRVIP will produce proper results.
As described in Section 6.2.1 to 6.2.7, some of the BWRVIP's assumptions in its VIPER simulations were inappropriate.
Further, the BWRVIP's PFM analysis identified only two reference cases -- one for axial welds and one for circumferential welds.
It did not consider the variability between vessels.
To determine the failure probability for all the BWR RPVs, the PFM analysis must consider the variability between vessels or present the failure probabilities and associated uncertainties as a function of adjusted RT or other appropriate indexing factor over the range aaplicable to BWRs.
Thxe variability between vessels will be addressed in tie staff's uncertainty analysis in Section 7.
The BWRVIP used the pressure test transient in all its analyses.
To provide a more risk-informed perspective, the staff examined extensively BWR transients and selected a foreign transient as the input transient for all the staff's analyses reported in the assessment.
The staff's discussion on the limiting transients is provided in Sections 5 and 7 and Appendix C.
7.0 INTEGRATED PROBABILISTIC ASSESSMENT The BWRVIP-05 report identified loss of feedwater or single safety / relief valve blowdown events as the most limiting operational transients which could challenge RPV integrity. The most limiting transients were identified as those which result in a rapid cooldown and rapid depressurization of the vessel shell welds.
Since the BWRVIP-05 report only considered risk from design basis events, the staff performed an independent review of LERs and ens to search for events outside of design basis that could be potentially more limiting to the RPV.
This review is discussed in Section 5.0 and in Appendix C.
Events that occur at cold shutdown, for example, were not identified as significant in the BWRVIP-05 report.
However, the staff review identified 35 potential precursors to cold overpressure events at U.S. BWRs and one actual cold overpressure event which occurred at a foreign BWR. Accounting for these precursor and actual events, the staff performed a preliminary probabilistic assessment of events that could challenge the RPV integrity at cold shutdown.
7.1 Frequency Estimation of Potential Initiating Events and Vessel Integrity Challenges Based on the staff findings of past shutdown events as discussed in Appendix C.2, four types of events have been identified that could potentially lead to RPV cold overpressurization conditions during shutdown.
These include (1) inadvertent injection (e.g., via feedwater, HPCS) as a result of spurious actuation, or failures of operator to follow procedures (act of commission),
(2) excessive cold water injection into the RPV prior to or during hydrostatic testing or vessel leak testing, (3) loss of reactor water cleanup (RWCU) system while the control rod drive pumps continue to operate, and (4) intentional injection as an act of commission without procedures or management or regulatory oversight. The likelihood of these events leading to a reactor pressure vessel low temperature overpressurization challenge, based on a limited set of historical data, is discussed below.
7.1.1 Inadvertent Injection During Shutdown As discussed in Appendix C.1 and shown in Table C-1, several inadvertent injections, including both high pressure and low pressura inj.ctions, have "f
0 e
m j
occurred during cold shutdown in U.S. BWRs.- An uncontrolled injection of feedwater or High Pressure Core Spray (HPCS) during shutdown could potentially result in a water-solid condition in the reactor and inadvertent repressurization of the RPV could occur with the potential to exceed the P-T curve limits.
One feedwater injection event and one HPCS injection event were identified in the staff LER search.
In addition, one HPCI injection event
_during a HPCI automatic initiation surveillance test at cold shutdown was reported.
It should be noted that none of these events resulted in a violation of TS P-T limits.
1 The LER search revealed a more frequent occurrence of inadvertent low pressure injections during shutdown via low Pressure Coolant Injection (LPCI) or Core i
Spray.
According to Appendix C.1, these injection modes can also repressurize the RPV while maintaining or lowering reactor vessel temperature.
As shown in j
Table C-1, a total of 14 low pressure injections were identified in the LER j
search.
Assuming that each one of the above injections (both high pressure and low' i
pressure) were precursors, the staff estimated the probability of an inadvertent injection to be 0.023.
This estimate accounted for 17 reported inadvertent injection events during an estimated 755 reactor years of U.S. BWR operation.
To arrive at a probability of 0.023, 755 reactor years of BWR operation was interpreted as 755 " plant states" during which a cold overpressurization event could occur.
The ratio of time at shutdown to total reactor years was accounted for by the initiating condition frequency estimated in the previous section. However, the staff notes that this probability estimate of inadvertent injection has large uncertainty and that a thorough review of industry data would be needed to obtain a more accurate
- estimate, i
Once an inadvertent injection occurs during a cold shutdown, the last barrier i
to prevent such an event from leading to cold overpressurization and thus challenging the vessel integrity is the operator / system recovery action. To estimate the probability of non-recovery, the above 17 precursor events were reviewed to note reactor conditions (e.g., high moderator temperature),
operator actions (e.g., operator stopping injection), and/or system recovery i
actions which prevented these events from propagating to a cold overpressurization condition.
Due to the complexity of incorporating these and other factors such as instrumentation failure into estimating a non-recovery probability, a bounding probability of 1/18, or 0.055, was assumed. This non-recovery probability value assumes that the next inadvertent actuation (18th of known/ reported event) would result in no recovery action. This non-recovery value, in comparison, is similar to a value of 0.05 suggested in NUREG/CR-1278 to depict a human error probability reflecting moderately high stress condition for dynamic interplay between the operator and system indications.(')
(')
NUREG/CR-1278, " Handbook of Human Reliability Analysis With Emphasis on Nuclear Power Plant Applications," assigned "... moderately high level of stress in a reliability analysis to transients that involve shutdown of 4
(continued...)
i
Based on the estimation of above parameters, the accident sequence 1,eading to challenging the vessel integrity was estimated to be about 1.0 x 10' / Year for a U.S. BWR. This sequence invo withafrequencyof0.075/ Year){vesaninitiatingcondition(coldshutdown
)
inadvertent injection (0.023), and operator / system failure to recover or restore the reactor to a stable condition (0.055).
7.1.2 Hydrostatic Tests or Vessel Leak Tests A vessel hydrostatic test is performed about once every ten years and a vessel
-leak test is performed once every refueling cycle.
As discussed in Appendix C, during a water solid leak test or a hydrostatic test, the RPV temperature and pressure are maintained at approximately 65.5'C to 93.3*C (150'F to 200*F) and 6.9 MPa (1000 psig), respectively, and are within the P-T curve limits. These temperature and pressure conditions are required by the plant TS; the specific limits are dependent upon the level of neutron embrittlement of the RPV.
The estimated frequency of hydrostatic injections is simply related to the test frequency.
However, because of the strict temperature conditions and oversight required for conducting hydrostatic tests, it is assumed that the likelihood that these tests could lead to an undesirable temperature and cold overpressurization of the RPV is negligible. As shown in Table C-1, none of the reported precursor events related to )erforming a hydrostatic or vessel leak test lead to a condition exceeding tie P-T limits.
The difficulty in quantifying the low temperature overpressurization challenge to a BWR vessel lies in estimating a probability of operator failing to heat the vessel in preparation for its leak or hydro test.
gowever,if})hishuman error is assumed to have a mean probability of 1.0 x 10' por event (, then
")(... continued)_
the reactor and turbine, to certain tasks during startup and shutdown which must be performed within time constraints..." The human error probability of 0.05, suggested as a high end value for dynamic tasks involving interplay between the operator and system indications, was assumed for this analysis.
")
The initiating condition reflects the plant state at the beginning of the event tree.
Thus, the initiating event frequency was estimated to be the percentage of time that the reactor would be considered in a cold shutdown condition with the vessel head tensioned. The staff assumed for this study an average plant capacity factor of 0.7'with 25 percent of the shutdown period being in a cold shutdown condition with the reactor vessel head tensioned. Thus, the frequency of plant state during which a cold overpressurization event could occur was estimated to be about 0.075/ Year. The staff notes that this frequency is plant specific and could vary substantially across industry.
(a)
Human error probability of 1.0 X 10'3 was estimated from NUREG/CR-4550, Vol. 6, " Analysis of Core Damage Frequency: Grand Gulf, Unit 1 Internal Events."
~ -.
1 estimating a frequency of RPV low temperature overpressurization challenge due to these tests can be derived by multiplying this value with the frequency of t
leak (andgydrostatic) tests. Assuming that a leak test is performed every in about 7.0 x 10',verall frequency from both leak and hydro tests results 1.5 years', the o
/ Year. The estimated uncertainty range of RPV low temperature overpressurization challenge frequency associated with the RPV hydrostatic or leak test is as follows: Assuming a log normal distribution of 4
estimatedat2.5x10',errorfactorof10,themediansequencefreg/ Year,and the frequency with an ency was
/ Year, the 5th ercentile value at 2.5 x 10 the 95"' percentile value at 2.5 x 10'p/ Year.
7.1.3 Loss of Reactor Water Cleanup System Another way that a vessel could experience cold overpressurization involves a loss of RWCU system which is used for vessel water letdown operation, while the control rod drive (CRD) pumps continue to inject water into the vessel.
Using a loss of RWCU frequency of 1.5 x 10'3/ Year (assumed in NUREG/CR-6143,
" Grand Gulf low Power and Shutdown Study"), the loss of RWCU during cold shutdown was estimated by multiplying this value by the fraction of the time the plant is in cold shutdown, 0.075. The probability of an operator failing to control the level via CRD was assumed to be 0.011 (also assumed in NUREG/CR-6143 to depict same accident scenario).
The point estimate of this 4
sequence frequency was calculated to be 1.2 X 10 / Year.
7.1.4 -Intentional Injection to Support Maintenance, Activities 3
A foreign BWR experienced a cold overpressurization event during shutdown. As described in Appendix C.2.1, a series of operator errors of commission resulted in the CRD pump continuing to run until the vessel went water-solid with no outflow from the reactor.
Unlike the above precursor events, this event challenged the reactor vessel, which resulted in a conditional vessel challenge probability of 1.0.
If a total of about 1700 reactor years of world-wide BWR operation is assumed to be from the same population (i.e., no differences between operator training, plant management, maintenance, operations, and hardware) it can be assumed that the frequency of cold overpressurizing the reactor from this class of event is about 6 x 10/ Year. The staff notes that a thorough review of industry (world-wide) data and information is needed to confirm the actual number of world-wide years of BWR operation and to determine if other low temperature overpressurization events have occurred.
]
7.1.5 Summary of Transient Frequency Estimation Based on the information and data collected via the staff LER search of past U.S. BWR events and staff assumptions, a frequency of cold overpressurization challenging vessel integrity for U.S. BWRs due to inadvertent injection was A refueling outage is assumed to occur once per 18 months. A hydrostatic test of RPV is performed once every 10-years. However, for the purpose of this analysis, it is assumed to replace the RPV leak test for the year that hydrostatic test is performed.
e estimated to be about 1.0 x 10/ Year. The sequence frequency leak and hydrostatic tests was estimated to be about 7.0 x 10',from the RPV
/ Year.
The sequence,frequencyfromtheRWCUisolationwasestimatedtobeag/ Year, out 1.0 x 10' / Year.
The total of these frequencies, about 8.0 x 10' accounts for the theoretical estimation of cold overpressurization frequency which is based on limited industry data and assumptions which include leak / hydrostatic test frequency, past inadvertent injection events, fraction of time vulnerable to cold overpressurization, and human error probabilities.
In addition to the above estimates, the staff estimated a cold overpressurization frequency based on one reported cold overpressurization event occurred at a fore,ign reactor.
The estimated sequence frequency from this event was 6.0 x 10 / Year.
Although the sequence frequency (6 x 10/ Year) estimated from an actual cold overpressurization event is in close p/ Year), it is noted that the two results roximity to the theoretical cold overpressurization frequency (8 x 10' represent different types of scenarios leading to cold overpressurization.
Because of the characteristics of the single foreign event (as a potential outlier), it would be premature to conclude that the theoretical estimation of.
cold overpressurization frequency realistically represents frequency estimate from world-wide operating data.
However, if the comparison is made within the general context or the general class of cold overpressurization events, the theoretical estimate appears to be consistent with an estimate derived from the world-wide operating data.
As mentioned previously, to better estimate the frequency of cold sverpressurization events, a thorough review of industry data and information would be needed to reduce the uncertainties associated with the selected parameter values and to improve the modelling capability.
7.2 Background for Probability of Failure and Uncertainty Analyses Integrated PFM analysis was performed by the staff to determine the P0F for potential BWR transients that may challenge irradiation embrittled BWR reactor vessel integrity.- This involves identifying the BWR transients that have non-negligible event frequencies, P(E), that may cause appreciable conditional probability of vessel failure, P(FjE), given that the transient under consideration occurs. The frequency of vessel failure, P(F), is determined by multiplying the estimated P(E) by the estimated P(FjE).
To determine the P0F of all BWR RPVs, the analysis must consider the variability between reactor vessels.
An unce6 Finty analysis was performed to determine the effect of the variability of the input parameters (e.g., flaw size distribution, neutron fluence, unirradiated reference temperature, weight percent copper, weight percent nickel, and K ) in the PFM analysis on the ie vessel conditional P0F.
The uncertainty analyses are performed by shifting the mean value of a given random variable by 2 sigma (a) in the adverse direction, while holding the other random variables at the reference case level, and determining the effect on P0F.
7.2.1 Reference Case Models Since there are significant differences in material properties for welds fabricated by the three vendors, each vendor (B&W, CE, and CB&I) was considered a reference case to determine the conditional P0F for BWR reactor vessels. The material properties for the welds fabricated by each vendor are discussed in Section 7.2.2.
7.2.2 Mean and Standard Deviation of Chemistry and Material Data The probability of vessel failure and the 95 percent confidence uncertainty bound between vessels are dependent upon the mean value and the standard deviation within a vessel and the standard deviation between vessels of the random variables in the analysis.
The mean value and standard deviation within a vessel are used in determining the reference case failure probability.
The standard deviation between vessels is used in determining the 95 percent confidence uncertainty bound. The method of determining the 95 percent confidence bound is discussed in Section 7.2.4.
Table 7-1 shows the three reference cases for the staff's fracture analysis of the BWR reactor vessels that were fabricated by CE, B&W, and CB&I.
The mean values, the standard deviation within a vessel and the standard deviation between vessels for the.unirradiated reference temperature, percent copper and percent nickel for CE and B&W fabricated welds were determined from the data in an SER of the impact of increased variability in chemistry on the RT"f'o'r the value of PWR reactor vessels (Reference 32).
The corresponding values CB&I vessels were derived by analysis of the data contained in the RVID.
Based on the similarity of materials, the analysis results for the NYS-fabricated vessel should be similar to the results from the B&W fabricated vessels, and the analysis results for the Hitachi-fabricated vessel should be similar to the results from the CB&I fabricated vessels.
BWR reactor vessels beltline welds that were fabricated by CE used the SAW process. The mean copper content from the RVID for SAW welds is virtually the same as the generic copper value for CE fabricated welds. The mean nickel content from the RVID for the SAW welds is lower than the generic nickel value for CE fabricated welds. Hence, utilizing the generic chemistry data for CE fabricated SAW welds is reasonable.
BWR reactor vessels beltline welds that were fabricated by B&W used the SAW and ESW processes. The ESW weld chemistries for each BWR beltline weld are contained in the RVID.
The mean copper content from the RVID for ESW welds is virtually the same as the generic copper value for B&W fabricated SAW welds.
The mean nickel content from the RVID for the ESW welds is lower than the generic nickel value for B&W fabricated SAW welds. Hence, combining data for these two processes appears reasonable.
BWR reactor vessels beltline welds that were fabricated by CB&I used the SAW and SMAW processes.
CB&I generic chemistry values were derived from data contained in the RVID. The mean values for copper and nickel contents were the average of all data from both SAW and SMAW welds.
The data from both of these welds are low in copper and high in nickel, and appear to overlap.
Hence, combining data for these two processes seems reasonable.
Table 7-1.
shows the various input parameters for the three reference cases, and provides a list of six uncertainty cases that are expected to provide information on uncertainties in the analysis model.
7.2.3 Mean and Standard Deviation of Neutron Fluence The mean values for the neutron fluence, the standard deviation of the neutron fluence within a weld, and the standard deviation of the neutron fluence between vessels were determined from data reported in the RVID and from the Cooper surveillance program (Reference 33). The value of the peak neutron fluence are 0.174 x 10}VID indicates that the m g
n/cmt, 0.073 x 10 n/cm,
a 0.263 x 10 g/cmt, and the stanp'ard deviation of tQ neutron fluence are 0.0929 x 10' n/cm2, 0.0327 x 10 n/cm2, 0.2224 x 10 n/cmt, for CE, B&W and CB&I fabricated vessels, respectively. The Cooper surveillance report indicates that the minimum value in a BWR is 45 percent of the peak value.
Using this data, the staff determined the mean value for the neutron fluence i
and the standard deviation within a weld for the generic CE, B&W and CB&I vessels.
The standard deviation of the analysis method was assumed to be 20 percent of the peak neutron fluence. The standard deviation of the values of neutron fluence between vessels was assumed to be the square root of the sum of the squares of the standard deviation ::f the analysis method and standard deviation of the vessel fluence for the CE, B&W and CB&I vessels.
7.2.4 Flaw Density and the Flaw Size Distribution The staff determined the best-estimate flaw size distribution from the inspection of the PVRUF vessel welds (Reference 24). The inspections were performed using the synthetic aperture focusing technique for ultrasonic testing (SAFT-UT).
The PVRUF inspection data was from welds that Iad not been placed in service, but were subject to PSI. Hence, the defects reported were those resulting from manufacturing.
From Table 7-2, there are a total of 148 indications at weld and fusion locations in the near surface zone.
Given that a flaw exists, the distribution of crack size is determined in two stages.
From Table 7-2,134 of the 148 near surface cracks were less than 0.0787 inch (2 mm) in length. Accordingly, the probability of a crack size that is equal j
to 0.0787 inch is assumed to be the ratio of 134 to 148, or 0.9054.
For cracks larger than 0.0787 inch, data for deeper indications was combined with the near surface data, because there was relatively little near surface-data (14 near surface and 42 deep cracks) and because there is no reason to believe that the crack size distributions for cracks larger than 0.0787 inch are significantly different for near surface and deep cracks.
The staff plotted the frequencies of crack sizes from Table 7-2 on a semi-log paper and found that the best fit line has a slope approximately equal to
-7.62.
For the values of crack size considered, the Marshall distribution with PSI is essentially a single exponential with a slope of -6.94, and this also fits the data quite well when plotted. Accordingly, it was assumed that the crack size distribution for near surface flaws greater than 0.0787 inch is given by a Marshall distribution with PSI.
The flaw size distribution for "x" greater than 0.0787 inch (0.2 cm) is tabulated in Table 7-3.
Equation (2) in Section 6.2.1 represents this flaw size distribution in the cumulative form.
o The staff's best-estimate flaw density is based on the observed number of indications in the near surface zone from the PVRUF data. Here, only weld and fusion locations are counted.
Table 7-2 indicates that there are a total of 148 such indications.
Because 0.14875 cubic meter of weld volume was inspected, the flaw density in the near surface zone is calculated to be 995 flaws per cubic meter.
Table 7-4 provides the upper 95 percent confidence bound flaw size distribution derived from analysis of the PVRUF inspection data.
The uncertainty of the assumed flaw density and flaw size distribution is determined as follows:
Because the assumed flaw density and flaw size distribution are based on counts, which have a Poisson distribution, confidence bounds for them can be derived from confidence bounds on a Poisson mean. The best-estimate flaw density is based on 148 indications that were observed in the near by replacing 148 by one-half of the 95'" percent confidence bound is found surface of the PVRUF data.
An upper 95 percentile of a chi-square distribution with 2(148) + 2 - 298 degrees of freedom.
This quantity is equal to 170. Therefore, an upper 95 percent confidence bound for the flaw density is 170 divided by the volume of the PVRUF welds (170/0.14875 - 1143 3
flaws /m ).
This is a 14.9 percent increase over the best-estimate value.
l The uncertainty in the crack size distribution is driven by the uncertainty in the probability of a crack size less than 2 mm (0.0787 inches) in length. Because this probability is close to one, its uncertainty is driven by its complement, the probability of a crack size greater than 2 mm.
From Table 7-2, this probability is estimated at 14/148 - 0.0946.
From a table for Poisson confidence bounds, an upper 95 percent confidence bound for this probability is 21.89/148 - 0.1479. A lower 95 percent confidence bound for the probability of a crack size less than 2 mm is therefore, 1 - 0.1479 - 0.8521.
This value is used for F(.0787) in the formula for F(a) to calculate the upper 95 percent confidence bound on the flaw size distribution, which is given by Table 7-4.
Since BWRs operate in an environment that is susceptible to SCC, the staff assumed that service would induce SCC in the cladding. The staff assumed that the SCC would penetrate through the clad and combine with the manufacturing fl aws. As a result of this assumption, the flaw sizes in Table 7-3 and 7-4 were increased by 0.20 inch (0.508 cm, " typical" thickness of BWR clad) to account for service induced stress corrosion cracking.
Generally the assumption that clad cracks would line up with manufacturing defects would be considered a conservative assumption. However, in reactor vessels that have been subject to repair on the inside of the weld or repairs to the cladding, it is possible that the repairs could result in cracks in the weld and increase the susceptibility for SCC.
In this case, it would be likely that the SCC would line up with the cracks initiated by the repair.
Cracks penetrating a repair in the clad and progressing into the adjacent base metal were observed in the reactor vessel head of Quad Cities Unit 2, (see NRC Information Notice 90-29, " Cracking of Cladding and Its Heat-Affected Zone in the Base Metal of a Reactor Vessel Head," dated April 30,1990).
The number of flaws assumed in the PFH analysis was a product of the flaw density and the volume of a generic BWR weld.
The volume of a generic BWR circumferential weld was calculated to be 0.109 cubic meter (for a generic vessel of internal radius 112.6 inches (286 cm), vessel beltline region thickness of 5.45 inches (13.8 cm) and weld thickness of 1.75 inches or 4.44 cm).
For the reference case,108 flaws per weld (994 X 0.109) was assumed in each analysis.
In the uncertainty study for the effect of the change in flaw size and density, the upper 95 percent confidence bound for the number of flaws was increased by 14.9 percent.
7.2.5 Residual Stresses and Crack Driving Forces for a Limiting Transient in BWR Vessels A limiting cold overpressure transient described in Appendix C (Section C.2.1), is considered for analyses.
This event consists of an internal pressure of 1150 psi at an isothermal temperature of 88'F.
" Typical" BWR vtssel geometric dimensions are assumed, i.e., an internal radius of 112.5 inches, vessel wall base-metal thickness of 5.25 inches, and a clad thickness of 0.2 inch.
Thermo-mechanical properties of the vessel are assumed as following:
Thermal conductivity (BTU /hr.ft.*F) = 24 (base metal),10 (cladding)
Specific heat f BTV/lb.*F) = 0.12 (base-metal, cladding)
Density (lb/ft ) = 489 (base metal, cladding)
Elastic Modulus (ksi) = 28,000 (base-metal), 22 Thermal expansion coefficient (/*F) = 6.9 X 10',,000 (cladding)(base-metal), 9 X 10' (cladding)
Poisson's ratio = 0.3 (base metal, cladding)
Cladding stress-free reference temperature (*F) = 515 Heat-transfer coefficient (BTU /hr.ft.2*F) = 1000 Due to isothermal internal pressure conditions in the vessel, hoop and axial stresses are due to internal pressure, clad base-metal differential thermal expansion, and the residual stresses due to the weld in process for the axial and circumferential welds.
Residual stress distribution through vessel thickness obtained from measurements in an actual PWR vessel axial welds are assumed to be applicable for the BWR vessels. The weld residual stress has a maximum amplitude of 6.5 ksi at the inner surface of the vessel and can be approximated with a third degree polynomial, for 5.45 inches 2 x 2 0 inch, as 3
o,, u,u,,, = [8.437 - 11.699x + 3.565xt - 0.292x ] ksi.
r The crack driving force, K is calculated for axial and circumferential semi-elliptical surface cracTuw,f aspect ratios (total crack length / crack u so depth) of 2, 6, and 10.
The K values are determined in FAVOR code at 10 locationsalongthecrackperiMuwry at 10 degrees elliptical angle apart between the free-surface point of the crack and the maximum depth.
The 2K is checked at each of the condition for cleavage crack growth (K 10 locations in the base-metal between NeNree a) face and the maximum depth i
sur points.
In FAVOR computations, a statistical sampling is performed on the crack aspect-ratios of 2, 6, and 10, based on the assumption of a uniform distribution for the aspect ratios. This will make the conditional P0F, P(FlE), to be smaller for the FAVOR code as compared to VIPER which assumes that the flaw is infinitely long..
7.2.6 Uncertainty Distribution for the Conditional Probability of Vessel failure The between-vessel uncertainty of the conditional probability of vessel failure, P(FlE), is best described by an uncertainty distribution (explained below in Section 7.4) because the mean values of the parameters which determine P(FIE) all vary from vessel to vessel.
There are five such parameters: neutron fluence, unirradiated reference temperature, percent copper, percent nickel and K. These parameters are assumed to be independentlyandnormallydYstributedwithmeansandstandarddeviations between vessels given by Table 7-5.
(The standard deviations within welds given in Table 7-5 are used in the Monte Carlo calculation of P(FlE)).
7.3 Conditional Failure Probabilities Tables 7-6, 7-7 and 7-8 present the conditional P0F, P(FIE) for axial and circumferential welds in CE, B&W and CB&I fabr!cated vessels for the reference cases as well as for the uncertainty cases with shifting of mean values one at a time by 20 for the fluence, copper, nickel, initial RT,Yr,om,hVRUF vessel K
fracture toughness, and new flaw density and depth distributions data for cleavage crack growth.
For the CB&I fabricated vessels with circumferential cracks, no failure was predicted for 10 million simulations performed for the various cases, and as s ch the conditional failure probability P(FIE) is less than 1.0 x 10' 7.4 Uncertainty Analysis A response surface approach is used to obtain an uncertainty distribution for P(FIE).
As an approximation, in(P(FlE)} is assumed to be a linear function of the first five parameters listed in Section 7.3 (this is the response surface).
Therefore, in(P(FlE)) has a normal distribution or, equivalently, P(F!E) has a log-normal distribution. The coefficients in the response surface are estimated by performing sensitivity runs one at a time for each of these five parameters and comparing the resultant failure probabilities with the reference case failure probability.
Based on these coefficients and the between-vessel standard deviations from Table 7-5, the standard deviation of in{P(FjE)) is calculated.
Between-vessel uncertainty bounds for P(FjE) are then calculated from this standard deviation and normal distribution tables.
For' example, a multiplier of 1.645 for the standard deviation is used to calculate a 95 percent uncertainty bound.
The 95 percent confidence bounds between vessels that results from the uncertainty analysis are reported in Table 7-9.
The best-estimate failure probabilities and the 95 percent confidence bounds for the reference case circumferential welds could not be determined accurately because insufficient failures resulted from these analyses. These analyses were based on 10 million vessel simulations.
The failure frequency due to cold overpressure events is denoted by Fr(E,F),
where E denotes the class of overpressure events and F denotes vessel failure.
This is decomposed by writing Fr(E,F) - Fr(E)P(FjE), where Fr(E) is the frequency of the class E and P(ElF) is the conditional probability of vessel failure given that an initiating event from E has occurred.
Based on the foreig,n cold overpressure transient, a point estimate of Fr(E) is 1/1700 = 5.88 x 10, where the denominator is the total number of BWR reactor years of world-wide experience.
From tables for confidence limits for the i
expectation of a Poisson variable, a 95% upper confidence bound for Fr(E) is 4.74/1700 = 2.79 x 10'3 This uncertainty bound can be combined with an uncertainty bound for P(FlE) to obtain an overall uncertainty bound for i
Fr(E,F).
7.5 Sensitivity to Flaw Density and the flaw Size Distribution In contrast to the uncertainty in P(F'E) due to the five parameters listed in Section 7.2.6, the uncertainty in P(F E) due to flaw density and the flaw size 4
i distribution is best described by a sensitivity study. While the five parameters vary from vessel to vessel, the flaw density and flaw size distribution are assumed to be constant from vessel to vessel.
They are uncertain because they are estimated from data as described in Section 7.2.4.
4 Accordingly, the uncertainty in P(FjE) due to flaw density and the flaw size distribution is described by the ratio of P(FjE) based on 95 percent upper confidence bounds for flaw density and the flaw size distribution to the reference case value of P(FlE). The pertinent P0F results have been reported in Tables 7-6 to 7-8.
They have been reproduced in Table 7-10 along with the results of the sensitivity study. However, since insufficient failures were reported for the reference cases for the circumferential welds, their sensitivity to flaw size and density could not be determined.
7.6 Sensitivity to Inservice Inspection The staff performed sensitivity studies to determine the effect that inservice i
inspection would have on the P0F. Two P0D of flaws were evaluated.
They are the POD of flaws resulting from the PISC studies and the P0D of flaws defined j
by Method C in the VIPER Code.
The POD from the PISC studies was developed by Pacific Northwest Laboratories (PNL).
Figure 7-1 compares the POD from the PISC study and Method C.
The staff believes that the PISC based POD should be considered a lower bound for ultrasonic inspection methods that have been qualified to Appendix VIII of ASME Code Section XI.
The conclusion is based on discussions with technical specialist at Pacific Northwest' Laboratories.
The VIPER Method C would be considered an upper bound P0D for ultrasonic inspection methods.
PNL developed flaw size distributions based on the PISC and Method C P00s.
The flaw size distributions included flaw sizing error of 10.2589 inch (10.658 cm) and flaw acceptance of 0.2071 inch (0.526 cm).
Flaws less than 0.2071 inch (0.526 cm) would be acceptable to ASME Code inspection standards and would not be required to be removed following inspection.
The best-estimate flaw size distributions resulting from PISC and Method C and the 95%
confidence bound flaw size distribution resulting from PISC are shown in Table 7-11.
The best-estimate flaw distributions were developed from the staff's flaw size distribution that is shown in Table 7-3, the 95 percent confidence flaw distribution was developed from Table 7-4.
The P0F for the flaw size distributions from PISC and Method C are compared to the P0F resulting from the reference cases in Table 7-12. The reference cases used the flaw size distributions in Table 7-3.
The sensitivity to inservice.
inspection is the ratio of the P0F from the PISC or Method C PODS to the P0F from the reference case.
Lower ratios would indicate that the inservice inspection method would have greater impact on reducing the P0F of the reactor vessel. Table 7-12 also contains the results from additional sensitivity studies performed to determine the effect of using the POD from the PISC studies on the staff results which were based on a 95 percent confidence bound flaw size distribution.
Table 7-1 Three Reference Cases for Staff's BWRVIP-05 Fracture Analysis of RPVs Random Variable CE Vessels B&W Vessels CB&I Vessels (p=Mean), (amStd. Dev.)
(Ref. Case 1)
(Ref Case 2)
(Ref. Case 3) p (Initial RTuo7)
- 56*F
- 5'F
- 56*F o (Initial RTuor) 16.7'F 19.C'F 16.7'F y (Fluence x10")
0.126 n/cm2 0.053 n/cm2 0.19 n/cm2 o (Fluence x10")
0.024 n/cm2 0,01 n/cm2 0.036 n/cmt p (Copper) 0.226 wt %
0.287 wt %
0.04 wt %
o (Copper) 0.062 wt %
0.06 wt %
0.019 wt %
y (Nickel) 0.76 wt %
0.6 wt %
0.93 wt %
o (Nickel) 0.032 wt %
0.0155 wt %
0.079 wt %
l o (6RTuoy) 24*F 24*F 24*F y (K,)
y from ASME y from ASME y from ASME i
a (K,)
0.147xp(K,)
0.147xp(K,)
0.147xp(K,)
i i
i i
y (K.)
y from ASME y froa ASME y from ASME i
o (K.)
0.1xp(K,)
0.ixy(K,)
0.1xp(K,)
i i
i Flaw Size & Density y and a from y and a from g and a from PVRUF data PVRUF data PVRUF data Sensitivity Studies by Shifting to Following Hean Values "One at a Time" p (Fluence x E19) 0.326 n/cm2 0.125 n/cmr 0.651 n/cm2 y (initial RT,or)
-50*F l'F
-50*F l
y (Copper) 0.256 wt %
0.317 wt %
0.05 wt %
y (Nickel) 0.776 wt %
0.608 wt %
0.96 wt %
y (K )
0.94xASME y 0.94xASME y 0.94xASME y ie Flaw Size & Density (p + 20) from (y + 20) from
(# + 20) from PVRUF data PVRUF data PVRUF data
s Table 7-2 Sumary of Data from SAFT-UT Inspection of PVRUF Vessel Welds PVRUF: Number of Near Surface Zone Indications by Category Through-Wall Extent of the Indications (DZ)-
09/15/95
<2m 2-3 m 3-4 m 4-5 m 5-6 m 6-7 m 7-8 m TOTAL
>2m Location V
P V
P V
P V
P V
P V
P V
P i
Clad 978 2
0 2
0 0
0 0
0 0
0 0
-0 4
0 Weld 87 0
5 0
2 0
0 0
0 2
1 0
1 2
9 Fusion (HAZ) 47 0
0 1
1 0
0 0
0 0
0 1
0 2
1 l
Base 392 4
14 1
4 0
2 0
0 1
0 0
1 6
21 Total 1504 6
19 4
7 0
2 0
0 3-1 1
2 14 31 Total Number Characterized > 2 m..... 45 l
Total Number < 2 m.....
1504 Total Number 1549 l
4
'I a
3 u
2,
l Table 7-3 TABLE 7-4 NRC Staff's Flaw Size Distribution Upper 95 percent Confidence Bound Flaw Size (Best Estimate)
Distribution from PVRUF Data FLAW SIZE CDF Probability FLAW SIZE CDF Probability a (inches)
Distribution a (inches)
Distribution Pr{a}
Pr{a}
0.0787 0.905405 0.905405 0.0787 0.852095 0.852095 0.2159 0.963304 0.057899 0.2159 0.942628 0.090533 0.3532 0.985740 0.022436 0.3532 0.977703 0.035075 0.4904 0.994435 0.008695 0.4904 0.991299 0.013596 0.6276 0.997817 0.003382 0.6276 0.996587 0.005288 0.7649 0.999138 0.001321 0.7649 0.998652 0.002065 0.9021 0.999656 0.000518 0.9021 0.999462 0.000810 1.0394 0.999861 0.000205 1.0394 0.999783 0.000321 1.1766 0.999942 0.000081 1.1766 0.999909 0.000127 1.3138 0.999976 0.000034 1.3138 0.999962 0.000053 1.4511 0.999989 0.000013 l.4511 0.999983 0.000021 1.5883 0.999995 0.000006 1.5883 0.999992 0.000009 1.7225 0.999998 0.000003 1.7225 0.999997 0.000005 1.8628 0.999999 0.000001 1.8628 0.999998 0.000001 2.0000 1.000000 0.000001 2.0000 1.000000 0.000002 - - -
TABLE 7-5 Values of Material Property Parameters Used in the BWR RPV Uncertainty Analysis Combustion Engineering Babcock & Wilcox RPV's Chicago Bridge & Iron RPV's RPV's Variables Mean Standard Standard Mean Standard Standard Mean Standard Standard Value Deviation Daviation Value Deviation Deviation Value Deviation Deviation Within A Of Means Within A Of Means Within A Of Means Vessel Between Vessel Between Vessel Between Vessels Vessels Vessels NeutrgnFluence 0.126 0.024 0.10 0.053 0.010 0.036 0.191 0.036 0.23 (X 10 n/cm2)
(E>1MeV)
Unirradiated
-56 16.7 3
-5 19.8 3
-56 16.7 3
Reference Temperature
(*F)
Weight Percent 0.226 0.062 0.015 0.287 0.060 0.015 0.04 0.019 0.005 Copper Weight Percent 0.76 0.032 0.008 0.60 0.0155 0.004 0.93 0.079 0.015 Nickel Static Fracture ASME 14.7% of 3% of ASME 14.7% of 3% of ASME 14.7% of 3% of Toughness (K,,.)
Mean ASME Mean ASME Mean Mean ASME Mean ASME Mean Mean ASME Mean ASME Mean
, l L_____
/~
4
..?-
-: +
1 t.
3 V
- {
TABLE 7-6 i
Conditional Probability of Failure P(FlE) for CE Fabricated BWR Vessels 2
Inner Surface Conditional PDF Conditional'PDF.
SN.
Case Description (RT., + 20)*F-Circumferential Flaws Axial Flaws l
1 Reference Case 1 93.2-
< l.0 x 10
l.5 x'10'3
}
r 137.5:
2.1 x 10'3 3.7 x 10'2 2
Mean Fluence - 0.326 x 10" n/cu 3-Mean Copper'.= 0.256 wt.%
97.4:
< 1.0 x 10
2.9'x 10-3 i
sI 4
Mean' Nickel - 0.776 wt %
94.4
< l.0'x 10
2.5 x 10'3 i
5 Mean Initial RT., - -50*F 99.2
< l.0 x 10
3.8 x 10'3
'l l
6 (p+20) Flaw Density & Flaw Depth 93.2 6.0 x 10
3.1 x 10'3 l
Distribution from PVRUF Vessel Data
.7 Mean Kie - 0.94 x ASME.Mean K 93.2 1.0 x 10
-4.3 x 10~2 ic t
Reference' Case definition.
l i
Mean Fluence
- 0.126 x 10" n/car Mean Copper
= 0.226 wt %
Mean Nickel
- 0.76 wt %
Mean Initial RT,
- -56*F i
t I
i I
f I
j.
l f
- 'l e
i I
TABLE 7-7 Conditional Probability of Failure P(FjE) for B&W Fabricated BWR Vessels Inner Surface Conditional P0F Conditional P0F SN Case Description (RTor + 20)*F Circumferential Flaws Axial Flaws 1
Reference Case 1 114.5 1.0 x 10
9.5 x 10-3 r
145.1 2.5 x 10'S 5.6 x 10~2 2
Mean Fluence - 0.125 x 10" n/ca 3
Mean Copper - 0.317 wt %
135.1 1.0 x 10
l.1 x 10-z 4
Mean Nickel - 0.608 wt %
114.9
< l.0 x 10
9.7 x 10'3 5
Mean Initial RT
- 58.3*F 120.5
< 1.0 x 10
9.8 x 10'3 or 6
(g+2a) Flaw Density & Flaw Depth 114.5 4.0 x 10
l.5 x 10'2 Distribution from PVRUF Vessel Data
~
7 Mean K,c = 0.9A x ASME Mean K,c 114.5 6.0 x 10
2.0 x 10'2 Reference Case definition:
Mean Fluence
- 0.053 x 10" n/cmz Mean Copper
= 0.287 wt %
Mean Nickel
- 0.6 wt %
Mean Initial RT.,
- -5'F 2
l t
r 1
~
TABLE 7 ~
Conditional Probability of Failure P(I.
..-C8&I Fabricated BWR Ve>sels Inner Surface Conditional P0F Conditional PDF i
SM Case Description (RT.,+ 20)*F Circumferential Flaws Axial Flaws 4
I Reference Case 1 32.7 N/A 4.6 x 10 f
4 2
Mean Fluence - 0.651 x E19 n/ car 50.0 N/A 1.7 x 10 3
Mean Copper - 0.05 wt %
40.5 N/A 4
Mean Nickel - 0.96 wt 7 32.7 N/A 1
5 Mean Initial RT., - -50*F 38.7 N/A l
6 (in20) Flaw Density & Flaw Depth 32.7' N/A 3.8 x 10*
)
, Distribution from PVRUF Vessel Data i
7
' Mean K,e - 0.94 x ASME Mean K,c 32.7 N/A Reference Case definition:
i r
l Mean Fluence
- 0.191 x 10" n/ca i
Mean Copper
- 0.04 wt %
Mean Nickel
- 0.93 wt %
Mean Initial RT.,
- -56*F i
t P
I l
8 t
I i
h t
i
! 1
.: 3 t
Table 7-9 Results of Uncertainty Analyses Fabricator Weld Orientation Reference Case Failure 95% Conf. Uncertainty Probability Bound Between Vessels Babcock & Wilcox Axial 9.5 x 10-3 4.6 x 10-2 (B&W)
Circumferential 1.0 x 10
> 1.4 x 10-5*
4*
Combustion Engineering Axial 1.5 x 10-3 2.9 x 10-2 (CE)
Circumferential
<l.0 x 10-r.
Chicago Bridge & Iron Axial 4.3 x 10
> 8.9 x 10-5
- 4 (CB&I)
Circumferential
<l.0 x 10-7
- l Notes:
Insufficient or no failures to accurately determine reference case failure probability and sensitivity to flaw size, flaw density and inservice inspection.
Failure Probability based on analyses of the neutron fluence only. Other variables were not analyzed for this case. However, the other variables in the other cases only made a small impact in the increare in the 95 percent confidence uncertainty bound.
. 4-+--w+-l- -~-
?
J
Table 7-10 Results of Sensitivity Analyses Fabricator Weld Orientation Reference Case Failure Sensitivity to Flaw Size Probability and Density" Babcock & Wilcox Axial 9.5 x 10~5 1.6 (B&W)
Circumferential 1.0 x 10 '
Combustion Engineering Axial 1.5 x 10~3 2.1 (CE)
Circumferential
<l.0 x 10-I
- Chicago Bridge & Iron Axial 4.3 x 10
>1.0 *
(CB&I) 1.0 x 10'#*
Circumferential Insufficient or no failures to accurately determine reference case failure probability and sensitivity to flaw size, flaw density and inservice inspection.
The sensitivity to flaw size and density is the ratio of the probability of failure using the 95 percent b
confidence bound for flaw size and density to the probability of failure using the best estimate flaw size and density. (Conditional PDF values taken from Tables 7-7, 7-8, and 7-9).
Observed ratio was less than 1.
Insufficient failures to accurately determine ratio.
________-__________________________________________l
~~
Table 7-11 NRC Staff's Flaw Size Distribution With Adjustment for Inservice Inspection Using POD from PISC Using POD from VIPER Method C Using POD from PISC FLAW SIZE CDF Probability CDF Probability QF Probability a (inches)
(best-Distribution (best-Distribution (95%
Distributica estimate)
Pr(a) estimate)
Pr(a}
confidence)
Pr(a}
(best-(best-(95%
estimate) estimate) confidence) 0.0787 0.950135460 0.950135460 0.964561300 0.964561300 0.922032500 0.922032500 0.2159 0.538873790 0.038739190 0.993327320 0.028766020 0.982606520 0.060574020 0.3532 0.998149148 0.009275358 0.999282796 0.005955474 0.997107020 0.014500500 0.4904 0.999548076 0.001398928 0.999721720 0.000438924 0.999294464 0.002187444 0.6276 0.999881868 0.000333792 0.999890820 0.000169100 0.999816371 0.000521907 0.7649 0.999956837 0.0000746 %
0.999956870 0.000066050 0.999933563 0.000117192 0.9021 0.999982737 0.000025900 0.999982770 0.000025900 0.999974063 0.000040500 1.0394 0.999992987 0.000010250 0.999993020 0.000010250 0.999990113 0.000016050 1.1766 0.999997037 0.000004050 0.999997070 0.000004050 0.999996463 0.000006350 1.3138 0.999998737 0.000001700 0.999998770 0.000001700 0.999999113 0.000002650 1.4511 0.999999387 0.000000650 0.999999420 0.000000650 1.000000163 0.000001050 1.5883 0.999999687 0.000000300 0.999999720 0.000000304 1.000000613 0.000000450 1.7255 0.999999837 0.000000150 0.999999870 0.000000150 1.000000863 0.000000250 1.2628 0.999999887 0.000000050 0.999999920 0.000000050 1.000000913 0.000000050 2.0000 0.999999937 0.000000050 0.999999970 0.000000050 1.000001063 0.000000107 m
Table 7-12 Results of Sensitivity Analyses due to ISI Fabricator Weld Sensitivity to Sensitivity to Sensitivity to Orientation Inservice Inservice Inspection Inservice Inspection
estimate) estimate) confidence)
Babcock & Wilcox Axial 0.33 0.26 0.31 (B&W)
Combustion Engineering Axial 0.29 0.20 0.24 (CE)
Chicago Bridge & Iron Axial 0.11 b
(CB&I)
Notes:The sensitivity to inservice inspection is the ratio of the probability of failure using the flaw distribution derived from the probability of detecting flaw during inservice inspection to the probability of failure using the best estimate flaw size and density.
6 No failure in 12 million simulations.
PISC II RRT Raculta fcr Selected Procedurea for Plata N2. 3 N
1.0 f-3 Special
'/
~
cn Procedures
_ (European)
/
c
/
/ MarsheN Committee
/
E f
/
/
2
/
8
/
r y
E o.s
/
/
l (New Proposal) 1 1
4
?
I iC I
(Industriel)
/
u I
a.
o.0 =-i i
i i
i H
0 10 20 30 40 so so 70 Defect Through wen Size (mm)
Figure Probabliky of Detection and Correct Sentencing Curves from PISC-H Studies (Nichols and Crutzen 1993) v e
-m-am-
t l
1.E+00 how Acceptance Standard = 0.2071 1.E r No inservice inspection POD = PISC Special Procedures A
.E.c I
POD = YlPER Method C i
o.
O 3: 1.E43 - t
.u.
%o g 1.E44 - r EE
.D I
=
$ 1.E45-
[
r u
a.
j 1.E4s -
I r
I 1.E47 l
l c.0 e.2 e.4 e.s e.s 1.9 1.2 1.4 1.s i.
Flaw Depth, Inch j
t i
t Figure Effect of Inservice Inspections with Dinering POD Curves on Maw Depth Distribution 1.
I,
l
\\
APPENDIX A REFERENCES 1
September 28, 1995, Letter from J. T. Beckham, Jr. to USNRC Document Control Desk,
Subject:
BWR Vessel and Internals Project, BWR Reactor Pressure Vessel Shell Weld inspection Recomendations (BWRVIP-05), EPRI TR-105697"
~
2 BWRVIP Response to the two NRC RAls on EPRI TR-105697, BWRVIP, June 24, 1996 3
October 29, 1996, letter to the staff modifying proposed axial weld inspection criteria 4
" Transmittal of BWRVIP VIPER Computer Code," dated May 16, 1997, in response to Reference 9 5
" Detailed Programming Information for VIPER," dated June 4, 1997, which provides information on the VIPER program including identification of the functions in each source file, a description of each function, the function call hierarchy structure, and a description of the variables for the VIPER program 6
" Responses to Requests for Additional Information [RAI) Regarding BWRVIP Recommendations for BWR Reactor Pressure Vessel Shell Weld Inspections,"
dated June 13, 1997 7
April 2,1996, Staff's Request for Additional Information (RAI) 8 May 20, 1996, Second RAI 9
May 20, 1997, Third RAI 10 Summary of July 18, 1995, meeting dated July 25, 1995 11 Summary of March 19, 1996, meeting dated March 26, 1996 12 Summary of October 15, 1996, meeting dated December 10, 1996 13 Summary of January 16, 1997, meeting dated February 13, 1997 14
. April 18, 1997 letter from the BWRVIP requesting a meeting with the Commission 1
15 Transcript of May 12, 1997, Commission meeting with the BWRVIP 16 May 30, 1997, Staff Requirements Memorandum (SRM M9705128) 17 Oak Ridge National Laboratory Report ORNL-NSIC-21, " Technology of Steel Pressure Vessels for Water-Cooled Nuclear Reactors, A Review of Current Practice in Design. Analysis, Materials, Fabrication, Inspection, and Test", December 1967, A-1 we -- --
-e
- rme e -+ * -e
--w--,
e
.w-
,-,*n,.re,vew,-
.r o w
---m m %w-,
,--,,,r--.
,,r,e
-m-.w.-
m
-s--r-.----
, --4.,
e em--
~*,w-,
,. - -w,,-,w-mr-v
18 March 6,1995, Letter from Pedro Salas to USNRC Document Control Desk,
Subject:
Browns Ferry Nuclear Plant-Unit 3-Reactor Pressure Vessel Shell Welds Augmented Examination and Inservice Inspection Relief Request 3-151-17."
J 19 Chockie, L.J. 1981. PVRC Round Robin Ultrasonic Program, Results and Assessment of Reliability, in Nondestructive Evaluation in the Nuclear Industry - 1980. American Society of Metals, Metals Park, Ohio.
20 PISC, 1979. Plate Inspection Program. Nuclear Energy Agency, Comm. On the Safety of Nuclear Installations, Org. For Economic Co-operation and Development. Paris France.
l 21 Ed. by Nichols, R. W., and S. Crutzen.1988. Ultrasonic Inspection of Heavy Section Steel Components, The PISC !! Final Report. Elsevier Applied Science. New York, New York.
NUREG/CR-4486," VISA-Il-AComputerCodeforPredictin$monen,robability the P 22 of Reactor Pressure Vessel failure", March 1986 F.A. S K.I.
Johnson, A.M. Liebetrau, D.W. Engel, E.P. Simonen, Battle 11e Pacific Northwest Laboratory, for U.S. Nuclear Regulatory Commission.
j 23 ORNL/NRC/LTR/94/1, " FAVOR - Fracture Analysis of Vessels: Oak Ridge,
=
Release 9401, A Fracture Analysis Code for Nuclear Reactor Pressure Vessels, Preliminary Users Guide", February 1994, T.L. Dickson, Oak Ridge National Laboratory, for U.S. Nuclear Regulatory Commission.
24 Doctor, S.R., G.J. Schuster, and F.A. Simonen, " Experience in Determining Fabrication Flaw Distributions in a Reactor Pressure Vessel," Proceedings of the Fourth International Conference on Nuclear Engineering (ICONE-4), Am(rican Society of Mechanical Engineers, Japan Society of Mechanical Engineers, March 10-14, 1996, New Orleans, j
25 Marshall Committee, "An Assessment of the Integrity of PWR Pressure Vessels," United Kingdom Atomic Energy Authority, October 1, 1976.
26 Ford, F.P., " Environmentally Assisted Cracking of Low Alloy Steels,"
Electric Power Research Institute (EPRI) NP-7473-L, Palo Alto, California, January 1992, 1
27 GE Nucinr Energy, Interim Technical Report on EPRI RPC 102-4, " Stress l
Corrosion Cracking in Low Alloy Steels," San Jose, California, February 1994.
28 Van Der Sluys, W.A. and Pathania, R., " Studies of Stress Corrosion Cracking in Steels Used for Reactor Pressure Vessels," Fifth International Symposium on Environmental Degradation of Materials in 4
Nuclear Power Systems - Water Reactors, Monterey, California, August 25-29, 1991.
A-2 4
--,--<ar m
n
>,,,n,-,
,,,~w-e-,,-w--.-
e,-e w
,,--w,~,,
r
.n-v e
,v s,
ea-+--,---,--r es-
{
29 Ruther, W.E., Kassner, T.F., and Soppet, W.K., " Environmentally Assisted l
Cracking in Light Water Reactors: Semlannual Report October-March 1990,"
NUREG/CR-4667 Vol. 12, ANL-91/5, January 1991.
30 Policy issue from J.W. Dirks (USNRC) to NRC Commissioners, " Enclosure A:
NRC Staff Evaluation of Pressurized Thermal Shock, November 1982." SECY-82-465, U.S. Nuclear Regulatory Commission, November 23, 1982.
31 Regulatory Guide 1.99, Revision 2 32 Memorandum from J. R. Strosnider to A. C. Thadani, " Assessment of the Impact of increased Variability in Chemistry on the RT Value of PWR Reactor Vessels," U.S. Nuclear Regulatory Commission, Na'y 5,1995, 33 Report GE-NC 523-159-1292, " Cooper Nuclear Station Vessel Surveillance Materials Testing and Fracture Toughness Analysis," February 1993.
A-3
-w--6=
.-----v~
APPENDIX B ACRONYMS AND INITIALISMS
'C degrees Celsius
'F degrees Fahrenheit mean a
standard deviation ACRS Advisory Committee on Reactor Safeguards ARI Alternate Rod injection ASME American Society of Mechanical Engineers ATWS Anticipated Transient Without Scram B&W Babcock & Wilcox BWR boiling water reactor DWRVIP Boiling Water Reactor Vessel and Internals Project CB&l Chicago Bridge and Iron CE Combustion Engineering CFR Code of Federal Regulations em centimeters CRD control rod drive CS core spray CST condensate storage tank d
expected number of defective circumferential welds DAC distance-amplitude-correction E
class of overpressure events EDDY eddy current inspection EN event notifications ESW electroslag welding F
vessel failure Fr(E) frequency of the class E Fr(E.F) failure frequency due to cold overpressure events FULLV full-V ultrasonic angle beam examination GE General Electric GMAW gas metal arc welding gpm gallons per minute HPCS high pressure core spray HPCI high pressure core injection ID inner diameter ISI in-service inspection K,
static fracture toughness i
K reference stress intensity LES low alloy steel LER licensee event reports LOCA loss-of-coolant accident LPCI low pressure coolant injection L/s liters per second a
number of circumferential welds nn millimeters MPa megaPascals n
number of axial welds n/cmt fluence, neutrons per square centimeter NDE non-destructive examination NRC Nuclear Regulatory Commission B-1
.... m-
-~
r.
i NYS New York Shipbuilding N(x) function to describe flaw size distribution OD outer diameter p
proportion of defective welds in the parent population P(E) non-negligible event frequencies l
P(ElF) conditional probability of vessel failure given that an initiating event from E has occurred P(F) frecuency of vessel failure P(F PFMlE) concitional probability of vessel failure probabilistic fracture mechanics PISC Plate Inspection Steering Committee P00 probability of detection P0F probability of failure PSI preservice inspection psig pounds per square inch gage P-T pressure-temperature curve PVRC Pressure Vessel Research Committee PVRUF pressure vessel research user facility PWR pressurized water reactor RAI request for additional information RHR residual heat removal RPV reactor pressure vessel RT radiographic inspections RT reference temperature of vessel weld material RWN reactor water cleanup RVID reactor vessel integrity database SAFT-VT synthetic aperture focusing technique for ultrasonic testing SAW submerged arc welding SCC stress corrosion cracking SER safety evaluation report SI Surveillance Instruction SLic multi-mode ultrasonic examination SMAW shielded metal arc welding SRM Staff Requirements Memorandum SRV safety relief valve TS technical specification UT ultrasonic testing x
crack length in inches B-2
~.
APPENDIX C LIMITING TRANSIENTS The BWRVIP-05 report discusses the operating characteristics of a BWR with respect to design transients and their effects on the vessel. These f
transients generally occur when a large steam region exists (i.e., a steam bubble is present). According to BWRVIP-05, the most limiting operational transients with respect to the vessel are loss of feedwater or single safety relief valve (SRV) blowdown events for normal and upset conditions.
Normal operating temperature and pressure for a BWR RPV is 260'C (500*F) and 6.9 MPa (1000psig).
During the hydrostatic test, RPV temperature and pressure is approximately 65.5'C to 93.3'C (150'F to 200'F) and 6.9 HPa (1000 psig) and is maintained on the pressure-temperature (P-T) curve for thtt particular vessel.
The most limiting transients for emergency and faulted operating conditions are any transient which causes or results in a rapid.ooldown, and rapid depressurization, of the vessel shell welds.
These transients are limiting for pre-existing cracks in the vessel shell welds.
In comparison, the BWRVIP-05 report states that the water-solid leak test condition, or hydrostatic test, is limiting for small flaws in the inside diameter of the vessel.
Based on the types of events discussed above, the staff performed a sampling review of approximately 20 years of licensee event reports (LERs) and event notifications (ens) to determine if other events (i.e., shutdown events) coJld be potentially more limiting to the vessel. The staff's sampling review does not encompass all LERs and ens due to the time restraints. The staff also notes that some overpressure events during shutdown, as described in Appendix C.1, may not be reportable under 10 CFR 50.72.
The preliminary staff sampilng review is discussed in Appendix C.2.
Table C-1 includes the results of the staff's sampling review of LERs and ens.
The table provides the maximum pressure and temperature reached at the end of the transient. Where the information was not provided, an NP is listed.
It should be noted that, although the following are examples of the types of events that could result in cold overpressure events, these events occurred at temperatures that are high relative to the reference temperature of the vessel weld material, as required by TS.
Therefore, these events did not, in end of themselves, represent significant challenges to the RPV.
C.1 Description of Rapid Depressurization and Shutdown Events The staff evaluated the following classes of transients and shutdown events with the potential to be more limiting than the operational transients discussed in BWRVIP-05.
C.l.1 Rapid Depressurization Rapid depressurization of the RPV during operation can occur as a result of a large break loss-of-coolant accident (LOCA), or lifting of a SRV.
Most facilities have technical specifications (TSs) for the rate of cooldown (i.e.,
cooldown limit of 37.8'C (100*F) per hour).
These TSs are in place to protect C-1
the vessel from pressurized thermal shock.
Examples of known rapid depressurization events at U.S. BWRs are as follows:
- 1. Hope Creek I had an event which cause reactor pressure to decrease from i
5.3 MPa to 689.5 kPa (765 psig to 100 psig), 266.7'C to 164.4'C (512*F to 328'F) respectively, in 51 minutes on October 10, 1987.
- 2. Perry had an event which exceeded the cooldown rate of 37.8'C (100*F) per hour on October 27, 1987.
- 3. Quad Cities I had a stuck open relief valve at 10 percent power which caused a cooldown rate of 93.3*C (200*F) per hour for the first hour and 46.l'C (115'F) for the second hour on April 17, 1989.
To assess the P0F for rapid depressurization, the staff assumed a hypothetical transient and performed a parametric study using the VISA-!! code.
The transient was characterized by a rapid depressurizatior, to 10 psi (69 kPa) with the fluid temperature decreasing from 550*F to 206'F (288'C to 97'C) in 12 seconds (i.e., equivalent to a cooling rate of 1720*F w 956*C per minute).
Use of this essentially instantaneous cooldown transient was meant to bound all real transients.
The staff examined both axial and circumferential welds inthisgiudy.
In one case, the staff intentionally increased the fluence to 5.0 X 10 n/cmt and the initial RT for the axial weld to 40'F to increase theprobabilityofproducingvesseffailure.
However, the staff's results indicated no failure for all cases subjected to rapid depressurization transient. As a result, the severe foreign transient discussed in Appendix C.2.1 was used in the staff's independent simulations reported in Section 7.
C.1.2 Hydrostatic Testing and/or Vessel Leak Test As discussed in Section 5.0, the hydrostatic test (or vessel leak test), is performed in the water solid condition with RPV pressure and temperature maintained on the veael P-T curve, as is generally specified in the licensee's TSs.
With the vessel in the water solid condition, any addition of water to the vessel increases RPV pressure while maintaining temperature. An example of potential water addition to the vessel is control rod drive (CRD)
) umps running with no reactor water cleanup (RWCU) letdown.
This condition las the potential to exceed the P-T curve limits.
C.I.3 Feedwater injection and/or High Pressure Core Spray injection An uncontrolled injection of feedwater or high pressure core spray (HPCS) during shutdown could potentially result in a water solid condition in the reactor. The staff notes that HPCS is applicable only to BWR 5 and 6.
The HPCS pump is capable of delivering at least 97.8 L/s (1550 gpm) at 7.9 MPa (1147 psig) reactor pressure, 385.4 L/s (6110 gpm) at 1.4 MPa (200 psig), and a maximum of 492 L/s (7800 gpm) at run-out conditions.
HPCS normally is aligned to the condensate storage tank (CST) for suction.
Depending on the temperature and pressure at the time of the injection, normally 48.8'C to 93.3'C (120*F to 200*F) and 0 HPa (0 psig), an inadvertent repressurization of the RPV could occur with the potential to exceed the P-T curve limits.
The staff notes that HPCS injection during shutdown is usually avoided due to precautionary tagouts during maintenance of the systems.
C-2
The rest of the BWR fleet have a high pressure core injection (HPCI) System.
HPCI is turbine driven and is not designed to operate below 689.5 kPa (100 psig).
Since the reactor is subcooled at atmospheric pressure during shutdown, inadvertent actuation of HPCI is generally not possible.
- However, the staff did identify one LER which describes an inadvertent HPCI injection during cold shutdown.
This LER is discussed in Section 5.2.
C.1.4 Low Pressure Core Injection and/or Core Spray Injection A more common event during shutdown is a LPCI or core s) ray (CS) injection.
Injections of this type can also repressurize the RPV wille mLintaining reactor vessel tem)erature.
If the injection is for a long duration, the potential exists t1at the injection may cause a reactor temperature decrease.
Both LPCI and core spray take suction from the suppression pool as the primary source of water.
An alternate alignment could be the CST.
The suppression pool water temperature ranges from 18.3'C to 35'C (65'F to 95'F), depending on the plant-specific TSs and the time of the year.
In general, each LPCI pump (RHR pump) is designed to produce a flow rate of 458 L/s (7260 gpm) at 1.2 MPa (172 psig) and a maximum flow rate of 537.5 L/s (8520 gpm) for run-out 1
conditions.
For this pump, shutoff head is less than 2.3 MPa (329 psig).
For some facilities, ccre spray is capable of delivering 394.3 L/s (6250 gpm) at 2.3 MPa (329 psig) and 482.9 L/s (7655 gpm) at run-out.
C.I.5 Unvented Condition During Shutdown An unvented reactor VM sel during shutdown with reactor temperature below a certain value is a violation of some plants' TSs.
In this condition, any addition of water to the vessel, from any source, without letdown, has the potential to repressurize the vessel.
However, this generally is a violation of plant TSs and, therefore, it is assumed that the licensees make every effort to avoid violation of their TSs.
C.2 Discussion of staff LER and Event Notification Review The staff performed a sampling review of LERs and ens.
The LER search involved reviewing potential BWR overcooling or overpressurization events since '.980.
A total of 229 LERs met the initial search criteria.
The EN search consisted of the same review of notifications since 1985.
Approximately 81 ens were identified which met the initial search criteria.
Of the 310 events identified, none of which were overlapping, 35 events were singled out as potential precursors to cold overpressure events at U.S. BWRr.
These events are identified in Table 5-1.
In achiition, the staff is aware of-a cold overpressure event which occurre:t at foreign BWR.
The following paragraphs highlight the circumstances of a of the more significant events and what factors prevented a complete overpreasurization and,otential failure of the vessel.
C.2.1 Cold Overpressure Event at a Foreign Reactor A cold overpressure event occurred at a foreign reactor during shutdown which appears to be more limiting than the leak test condition.
The foreign reactor had nearly completed its sixth refueling outage with the reactor vessel head tensioned and RHR-A running in shutdown cooling mode. At the beginning of the C-3
~
.~
i..
event, reactor water level was approximately 15 cm (5.9 inches) below the
'1 vessel flange and the main steam line isolation valves were in the shutoff position.
in addition, the SRV, the vessel head vent pipe, and RWCU were administrative 1y blocked.
A weld overlay on a recirculation loop was in progress and required the "B" recirculation pump to be aligned to enhance heat removal capability. The CRD pump was started to establish sufficient seal 3
injection flow for the recirculation pump. CRD flow was approximately 1.9 L/s (30.1 gpm).
As reactor water level gradually increased, 0)erators aligned a RHR pump and 1
related valves to letdown the excess water.
.ater on that day, maintenance personnel were informed of inaccurate level indication due to reference leg i
leakage. Almost concurrently, a reactor o>erator shift turnover occurred, shift 2 to shift 3.
Records showed that tie information regarding the inaccurate level-indication was passed to shift 3.
At shift turnover, the level indication was reading 700 cm (275.6 inches) based on the instrument zero line with a system pressure approximately 0 MPa (0 psig).
Actual reactor water level was approximately 500 cm (196.8 inches).
This condition continued until reactor system pressure reached 917 kPa (133 psig) and, on high pressure, tripped the RHR pump. The operators were not aware that the RdR pump had tripped.
The CRD ) ump continued to run until the vessel went water-solid with no outflow from tie reactor.
Pressure continued to increase at a higher rate until reactor system pressure reached 7.9 MPa (1150 psig) and tripped the CRD pump due to low suction head. The "B" recirculation pump also tripped at 7.9 MPa (1150 psig). During this event, reactor coolant temperature was maintained around 26*C to 31*C (79'F to 88'F).
C.2.2 Hydrostatic Tests and Vessel Leakage Tests at U.S. Plants On March 19, 1989, the Clinton RPV was water solid for the leakage test and the trode switch was set to REFUEL. Concurrently, scram time testing was being At the completion of the scram time testing, the reactor mode performed.
switch was repositioned to SHUTDOWN which initiated an expected scram signal.
This signal caused the scram inlet.and exhaust valves to open which resulted in the control rod drive bypass flow being directed to the water-solid This in turn caused the reactor pressure to increase at a rate of reactor.
approximately 48.2 kPa (7 psig) per second and resulted in 4 SRVs lifting per Operators attempted to reduce pressure, prior to the SRVs lifting, by 4
design.
using the discharge paths established for the RPV pressure test but were unsuccessful.
The maximum pressure recorded during the event was 7.8 MPa (1130 psig).
Temperature at the beginning of the event was 71.l*C (160*F).
No change in temps.rature was reported.
On May ll, 1993, a system leakage test of the RPV was being performed at J
Browns Ferry Unit 2.
Pressure was being maintain using a CRD pump and RWCU reject flow control valve.
In parallel, a " Functional Line Flow Check Valve" was being performed.
Due to isolations with Surveillance Instruction (SI)d indications that reactor pressure was both tests, operators receive The unit operator attempted to maintain pressure in the band decreasing.
The prescribed by the ASME Section XI test by lowering RWCU reject flow.
resultant pressure increase resulted in an Anticipated Transient Without Scram (ATWS) scram signal that tripped the reactor recirculation system pumps and L
C-4
initiated an Alternate Rod injection (ARI) signal.
The highest pressure encountered during the event was 7.7 MPa (1120 psig).
Moderator temperature at the beginning of the event was 87.8'c (190*F). No change in temperature was reported.
i C.2.3 feedwater Injection and High Pressure injection Events at U.S. Plants On January 25, 1984 Peach Bottom Unit 3 was set up for long path rocirculation (feedwater system flush to the condenser) after completion of maintenance on the reactor feed pump bypass valve.
The operator failed to close the feedwater inlet valves to the RPV as required in the system
)rocedure. With a condensate pump in service, the operator opened the 5th 1 eater outlet valve and inadvertently injected condensate into the reactor i
vessel.
Reactor water level increased approximately six feet and a minimal 4
pressure increase was noted on the wide range reactor pressure stri) chart recorder.
This pressure increase was estimated to be less than 69 (Pa (10 psig). Minimum reactor vessel temperatures recorded were 42.2*C (108'F) in the A recirculation suction line and 46.l*C (115'F) at the feedwater nozzle.
This transient violated the TS limits of 48.9'C (120'F)h Bottom Unit 3 TSs.and 0 kPa (0 referenced in Paragraph 3.6.A. Figure 3.6.2 of the Peac On July 27, 1990, Nine Mile Point Unit 1 in an extended refuel outage with the core loaded and with the mode switch in the REFUEL position.
The HPCI System i
initiated while the licensee was preparing to perform a HPCI Automatic Initiation Surveillance Test.
The feedwater isolation valves were opened by procedure. Due to a worn cam in the feedwater flow control valve positioner, feedwater leaked into the reactor vessel. A high reactor water level, 241.3 cm (95 inches), signal was received which initiated a turbine trip signal.
HPCI initiated and injected into the reactor vessel for approximately 3 seconds. At the time of the event, tha reactor vessel water temperature was about 65.6*C (150*F).
1 On May 3,1995, during the performance of the LaSalle Unit 2 Reactor Vessel Water level Roference leg Continuous Backfill Panel 2Cll-P002 Operation, instrument maintenance personnel returned a continuous backfill panel to service.
The backfill-panel provides a low continuous flow of high pressure water to the reactor vessel instrument reference leg piping to ensure that the reference column is filled to its proper height and that any non-condensable gases are flushed from the reference leg piping. During the valving sequence specified in the procedure, a pressure line spike caused the HPCS diesel generator and HPCS pump to initiate.
The HPCS pump started and injected into the vessel.
Reactor water level rose approximately 20.3 cm (8 inches),
if the reactor operator had not secured the HPCS pump, the HPCS system would have stopped injecting due to high level closure of the HPCS injection valve. The i
addition of suppression pool water did not reduce the reactor water temperature to below the minimum bolt-up temperature, 30'C (86*F),
or the minimum temperature used in shutdown margin calculations, 20'C (68'F).
Reactor water temperature remained above 48.9'C (120*F) during the event.
C.2.4 Low Pressure Core injection and/or Core Spray injection at U.S. Plants l
On May 15, 1985, instrument and control personnel were in the )rocess of j
performing the " Channel Logic Response Time" procedure at Hatc1 Unit 2.
The C-5
---.-,,,~-m.,,,-m or m.--
.._-,.-.r---..
.---,--.-,.--r-
.,-.--,, =,e
procedure mistakenly called for a jumper to be placed between AA3 and AA4 in panel 2Hll-P617; proper placement was between AA3 and AA4 in panel 2Hll-P627.
RHR pump C was running in shutdown cooling mode.
The improperly installed jumper initiated a loop "A" LOCA signal which caused RHR pumps "B" and "D" to start in the LPCI mode and inject water from the torus into the RPV.
RPV level increased from approximately 91.4 to 254 cm (36 to 100 inches),
referenced to instrument zero, as a result of the injection.
The RPV pressure and temperature at the time of the event was not reported in the LER.
On June 9,1988, during performance of the Cooper Surveillance Procedure 6.3.4.3 on the emergency diesel generators, approximately 56,775 liters (15,000 gallons) of suppression pool water were injection into the reactor vessel.
The surveillance procedure is performed once per cycle to functionally test the emergency start of each emergency diesel generator, RHR pump and core spray pump, as well as the loading sequence of safety-related equipment on the associated diesel generator. Upon simulating the low reactor vessel water level, the DC powered RHR loop "B" inboard injection valve began to open.
When the critical bus was re-energized, the RHR loop "B" outboard injection valve began to open and the RHR loop "B" outboard to suppression pool valve began to close.
These valve operations, along with RHR pump "D" starting when the bus was re-energized, created an injection path into the reactor vessel via RHR loop "B."
Additionally, the core spray loop "B" outboard injection valve and the core spray loop "B" inboard injection valve began to open.
When the core spray pump "B" started ten seconds later, an additional flow path via core spray loop "B" was created.
The total injection time for RHR loop "B" was 97 seconds, and 53 seconds for core spray.
Approximately 41,635 to 45,420 liters (11,000 to 12,000 gallons) of water was injected by RHR and 11,355 to 15,140 liters (3,000 to 4,000 gallons) of water was injected by core spray to the RPV.
Had the injection continued with no operator action, the reactor vessel would have completely filled and pressure would have stabilized at approximately 2.3 HPa (340 psig), shutoff head for core spray.
Actual reactor vessel metal temperatures ranged from 76.7'C to 87.8'C (170*F to 190"F) in the flange area, with no vessel temperatures below 51.7'C (125'F).
On July 6,1994, instrument and control personnel were back filling instrument lines to support excess flow check valve testing at Washington Nuclear Plant Unit 2.
Due to a lineup error, the differential pressure sensed by the inservice reactor vessel level detectors was increased creating an invalid low level indication which caused several automatic actions including a low pressure core spray system actuation and injection.
The residual heat removal (RHR) system was in a shutdown cooling lineup, and therefore, no LPCI system injection to the RPV occurred via the RHR system.
Reactor water level increased approximately 50.8 cm (20 inches) due to the core spray injection.
The injection water temperature was 21.l*C (70'F) with the RPV at 54.4'C (130'F) and 730.9 kPa (106 psig).
C-6
TABLE C Sampling Review of LERs and ens Plant Date Type of Event P (psig)
T ('F)
Nine Mile Point 1 06/21/84 Hydrostatic test 1068 NP Peach Bottom 2 05/30/85 Hydrostatic test 1030 NP Shoreham 06/06/85 Hydrostatic test 1047 NP Clinton 03/19/89 Hydrostatic test 1130 160 Browns Ferry 2 05/11/93 Hydrostatic test 1120 190 Peach Bottom 3 01/25/84 feedwater inj.
10 108 l
Nine Mile Point 1 07/27/90 HPCI inj.
NP 150 l
LaSalle 2 05/03/95 HPCS inj.
0 120 Limerick 1 03/26/85 LPCI inj.
NP NP Hatch 2 05/15/85 LPCIinj.
NP NP Shoreham 05/05/87 LPCI inj.
125 160 Nine Mile Point 2 02/02/89 LPCI inj.
O Amb.
Brunswick 2 10/21/86 Core spray inj.
NP NP FitzPatrick 04/07/87 Core spray inj.
NP NP Cooper 05/26/88 Core spray inj.
NP NP Vermont Yankee 04/12/92 Core spray inj.
NP NP WNP2 07/06/94 Core spray inj.
106 130 Brunswick 2 05/30/86 Core spray & LPCI inj.
NP 150 Dresden 2 02/05/89 Core spray & LPCI inj.
NP
> 100 Hope Creek 03/08/92 Core spray & LPCI inj.
O NP Cooper 06/09/88 RHR & Core spray inj.
< 340 170/190 Hatch 2-11/14/92 RHR & Core spray inj.
NP NP Peach Bottom 3 09/28/81 Unvented condition 78
<120 Dresden 3 07/11/84 Unvented condition NP 130/140 Dresden 2 10/08/85 Unvented condition NP 133 Dresden 2 05/24/86 Unvented condition NP 130/140 Peach Bottom 3 10/26/89 Unvented condition NP 155 Nine Mile Point 1 08/03/86 Pot. HPCI inj.
NP NP C-7
.. o
-~
i TABLE C Sampling Review of LERs and ens Plant Date Type of Event P (psig)
T (*F)
Perry 07/11/86 Pot. HPCS inj.
NP NP WNP2 06/18/89 Pot. HPCS inj.
122 NP WNP2 06/18/89 Pot.HPCSinj.
NP NP WNP2 06/18/89 Pot. HPCS inj.
NP NP WNP2 07/04/94 Pot. HPCS inj.
NP NP Nine Mile Point 2 10/08/88 HPCS init. - no inj.
35 104 Nine Mile Point 2 02/19/89 HPCS init. - no inj.
2 129 Pot.
Potential inj. = injection init. = initiated i
t C-8
RE0VESTFORADDITIONALINFORMATIONREGARDiNGREPORTEPRITR-105697 "BWR VESSEL AND 1NTERNALS PROJECT - BWR REACTOR PRESSURE VESSEL SHELL WELD INSPECTION RECOMMENDAT'ONS (BWRVIP-05).*
SEPTEMBER 1995 A.
Identification of Types and frequencies of Beyond Design-Basis-Accident (DBA) Events 1.
Provide an evaluation of the potential for Beyond-DBA events, including low-temperature over-pressure (LTOP) events.
2.
Describe the types of Beyond-DBA events identified in your assessment.
3.
Provide an estimate of the frequency of LTOP events for different BWR designs, as appropriate.
4.
Provide an estimate of the expected maximum pressures and temperatures of occurrence for LTOP events.
Provide the following information supporting the above evaluations:
5.
The average fraction of a given year that a typical BWR is in cold
- shutdown, 6.
The fraction of cold shutdown period that a typical BWR is susceptible to cold overpressure conditions.
7.
What are the possible ways of injecting water into the vessel during cold shutdown? Describe how these injections could result in a cold overpressurization condition.
8.
Describe the recovery actions that can be taken in response to inadvertent injections. What plant features can help mitigate such accidents? What is the estimated recovery time from an inadvertent injection? What indications are available to the operators to diagnose the inadvertent injection event? Estimate the non-recovery probability.
9.
What training and/or procedures are provided to prevent inadvertent injections from occurring? What training and/or procedures are provided to prevent cold overpressurizations from occurring given an inadvertent injection?
- 10. Estimate the probability of operator failing to heat the vessel in preparation for its leak or hydro test.
What training and/or procedures are provided to prevent this operator error from occurring? (Non-Proprietary)
.4
- 11. Identify initiating events (precursors) and their frequencies that could lead to cold overaressurization events -- considering all modes of operation.
Provide listorical industry data; otherwise, provide bases for assumed parameter values.
(Note: the staff review / analysis only contained Licensee Event Report data dating back to 1900.)
12.
Identify miti conditions - gating features for preventing cold overpressurization both at system / equipment and operator response levels and the associated failure probabilities.
Provide historical industry data; otherwise, provide bases for assumed parameter values.
B. Conditional RPV Failure Probabilities 1.
Provide an assessment of the conditional probability of RPV failure for the Beyond-DBA identified in response to Section A, above.
2.
Presentthe({.onditionalfailureprobabilitiesforthe-threeBWRRPV fabricators ) (e.g., Babcock & Wilcox
- B&W), Chicago Bridge and Iron (CB&l), and Combustion Engineering [CE;) as a function of adjustad RT,o, or T-RT,,,d cond' tional failure probabilities., as a)propriate, to support integratio frecuencies an If the concitional failure probabilities are presented as a function of RT,,,
alone, 3rovide an evaluation of the sensitivity to the temperature at which tie event occurs.
3.
Provide an uncertainty analysis accounting for the variability of parameters within an RPV.
4.
Provide the basis for the conclusion that large, deep circumferential1y-oriented flaws have a low probability of existing in BWR RPV circumferential shell welds,
- a. From the analysis of circumferential welds, identify the location, size and frequency of flaws that could lead to vessel failure,
- b. For the flaws in a), what frequency were they assigned in the flaw distribution used in the fracture analysis? Provide the basis for the flaw frequency, include BWR, PWR and research ex:.mination results.
c.
Identify the BWR RPV shell welds that have been inspected using an ultrasonic inspection procedure that is capable of detecting flaws of the size and location identified in a).
Provide the procedure used in the ultrasonic examination and identify whether the procedure has been qualified to detect flaws of the size and location identified in a).
") Note that, based on the staff's evaluation, the New York Shipbuilding (NYS) fabricated vessel is comparable to the B&W vessels and the Hitachi fabricated vessel is comparable to the CB&l vessels.
2-
1
,o o
- d. Based on the BWR RPV shell welds ultrasonic inspection results what can be inferred as to the frequency of flaws of the size and location identified in a) in BWR RPV shell welds?
Address the following regarding the VIPER Code:
NOTE:
The following questions are Proprietary and are not included in this Non-Proprietary version C. Estimated frequency of RPV Failure 1.
Provide an assessment of the frequency of RPV failure based on the evaluations performed in Sections A and B, above.
2.
Provide an uncertainty analysis for the frequency of failure assessment.
- D. Consequences of RPV Failure 1.
Provide an assessment of the consequences of RPV failure which-includes the mode of RPV failure (e.g., brittle vs. ductilti and the configuration of the containment during periods of RPV failure vulnerability.
2.
Provide an assessment of the relationship of the frequencies of RPV failure addressed in Section C and whether they correspond to core damage frequencies (CDFs) or large early release frequencies (LERFs).
E. Risk-Informed Assessment 1,
Provide a risk-informed assessment of the proposed change in RPV inspections based on the results from above and the guidance in Draft Regulatory Guide (DG) 1061.
Considering the guidance in DG 1061, identify how defense-in-depth and safety margins will be maintained.
i
,e
)
- 4 l
Plant-Specific Relief Requests 4
l 1.
Provide a plant-specific (or bounding) assessment of the estimated j
frequency of RPV Failure due to Beyond-DBA events.
Include an assessment of the plant-specific potential for L10P events and any 4
)
special precautions (e.g., operator training, procedures, etc.) in place j
to avoid such events.
2.
For any plants that will be requesting schedule relief, regarding the inspection of RPV circumferential welds, from the 10 CFR 50.55a(g)(6)(ii)(A) requirement that all licensees perform an augmented i
inspection of RPV welds, provide:
-a.
The mean and upper bound (per RG 1.99) adjusted reference temperature (ART)hich the relief is being requested.at the outage corre time period for w Include:
j 1.
the initial (unirradiated) reference temperature 11.
the chemistry factor (CF)-
1
Neutron Fluence c.
The increase in reference temperature due to irradiation (ARTm) d.
The Margin term
. _