ML20024E661
| ML20024E661 | |
| Person / Time | |
|---|---|
| Site: | Limerick  |
| Issue date: | 08/15/1983 |
| From: | Azarm M BROOKHAVEN NATIONAL LABORATORY |
| To: | NRC |
| Shared Package | |
| ML20024E662 | List: |
| References | |
| CON-FIN-A-3393 NUDOCS 8308160343 | |
| Download: ML20024E661 (188) | |
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1 NUREG/CR-BNL-NUREG-A PRELIMINARY REVIEW 0F THE LIMERICK GENERATING STATION SEVERE ACCIDENT RISK ASSESSMENT VOLUME I: Core Melt Frequency M. A. Azarm, R. A. Bari, J. L. Boccio, N. Hanan, I. A. Papazoglou, C. Ruger, K. Shiu, J. Reed *, M. McCann*, A. Xa f ka **
Engineering and Risk Assessment Division Department of Nuclear Energy Brookhaven National Laboratory Upton, New York 11973 August 15, 1983 i
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-111-ABSTRACT A preliminary review is performed of the Severe Accident Risk Assessment for the Limerick Generating Station. The review considers the impact on the core melt frequency of seismic and fire initiating events. An evaluation is performed of methodologies used for determining the event frequencies and their impact on the plant components and structures. Particular attention is given to uncertainties and critical assumptions. Limited requantification is per-formed for selected core melt accident sequences in order to illustrate sensi-tivities of the results to the underlying assumptions.
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! ABSTRACT ................................................................ iii LIST OF FIGURES .........................................................
4 LIST OF TABLES ..........................................................
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SUMMARY
1.0 INTRODUCTION
....................................................... 1-1 1.1 Background .................................................... 1-1 1.2 Objecti ve , Scope , and Approach to Revi ew . . . . . . . . . . . . . . . . . . . . . . 1-1 1
1.3 Organization of Report ........................................ 1-2 2.0 EXTERNAL INITIATING EVENT CONTRIBUTORS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-1 2.1 Review of the Seismic Hazard and Fragility Analyses ........... 2-1 i 2.1.1 Introduction ........................................... i-1 2.1.1.1 Sensitivity Analysis for Seismic Effects ...... 2 -4 2.1.1.2 Sei smi c Secti on Orga ni zati on . . . . . . . . . . . . . . . . . . 2-6 1
2.1.2 Se i smi c Ha z a rd . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-6 2.1.2.1 Review Approach ............................... 2-6
, 2.1.2.2 Sei smi c Haza rd Methodol ogy . . . . . . . . . . . . . . . . . . . . 2-6 2.1.2.3 Seismogenic Zones ............................. 2-9 2.1.2.4 Sei smi ci ty Parameters . . . . . . . . . . . . . . . . . . . . . . . . . 2-14 2.1.2.5 Ground Motion Attenuation ..................... 2-17 2.1.2.6 Comparison of the LGS Hazard A~ n alysis with Hi st o ri c Sei smi ci ty . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-18 2.1.2.7 Su mma ry . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-21 2.1.3 Se i smi c Frag i l i ty . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-23 2.1.3.1 Dama ge Fa ct o r . . . . . . . . . . . . . . . . . . . . . . . . . . r. . . . . . 2-24 2.1.3.2 Upper Bound Accelerations ..................... 2-30 2.1.3.3 Reactor Enclosure and Control Structure ....... 2-31 2.1.3.4 Reactor Pressure Vessel Capacities ............ 3-32 2.1.3.5 Potential Impact Between Reactor Building and Containment ................................... 2-33 2.1.3.6 El ectri c and Cont rol Equi pment . . . . . . . . . . . . . . . . 2-36 2.1.3.7 Review of Significant Components .............. 2-37 J
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-v-TABLE OF CONTENTS (Cont. )
Pace 2.1.3.8 General Fragi l i ty-Rel ated Comments . . . . . . . . . . . . 2-45 2.1.3.9 Closure ....................................... 2-48 2.1.4 References to Section 2.1 .............................. 2-49 2.2 Fire .......................................................... 2-58 2.2.1 Determi ni sti c Fi re Growtn Mocel i ng . . . . . . . . . . . . . . . . . . . . . 2-58 2.2.1.1 Introduction .................................. 2-58 2.2.1.2 Summary Evaluation of Deterministic Fire Growtn Moceling ............................... 2-60 2.2.1.3 Detailed Evaluation of Cetermit istic Fire Growen Modeling ............................... 2-62 2.2.1.3.1 Fuel Burning Rate .................. 2-62 2.2.1.3.2 Fu el El ement Igni ti on . . . . . . . . . . . . . . 2-64 2.2.1.3.3 Fire Near Enclosure Walls or C o rn e rs . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-68 2.2.1.3.4 Stratified Ceiling Layer ........... 2-68 2.2.1.4 Recommendations for Improving Fire Growth Modeling ...................................... 2-69 2.2.2 Probabi l i sti c Fi re Analysi s Revi ew . . . . . . . . . . . . . . . . . . . . . 2-70 2.2.2.1 Evaluation of Significant Fire Frequencies in Gene ra l Locat i on s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-71 2.2.2.1~1 . Sel f-Ignited Cabl e Fi res . . . . . . . . . . . 2-72 4
2.2.2.1.2 Transient Combustible Fi res .. .. .. .. 2-73 2.2.2.1.3 Power Distribution Panel Fires ..... 2-73 2.2.2.2 Screening Analysis ............................ 2-73 2.2.2.3 Probabilistic Modeling of Detection and Suppression ................................... 2-74 2.2.2.4 Probabilistic Modeling of Plant Damage State .. 2-75 2.2.2.4.1 Zone Speci fi c Comments . . . . . . . . . . . . . 2-78 2.2.3 References to Section 2.2 .............................. 2-80 3.0 ACCIDENT SEQUENCE ANALYSIS ......................................... 3-1 3 .1 - Se i s mi c . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-1
, 3.1.1 - Pl ant Frontl i ne Syst ems . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-1
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Page 3.1.1.1 Overview of the LGS-SARA Approacn in Frontline System Modeling ............................... 3-1 '
3.1.1.2 BNL Revision and Review of Frontline System Fault Trees ..............J.................... 3-2 3.1.2 Acci dent Sequence Analysi s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-13 3.1.2.1 Overview of L35-SARA Accident Se Analysis .......................quence ............... 3-13 3.1.2.2 BNL Review of Accident Sequence Quantifica-
-l tion .......................................... 3-14 3.2 Fire.......................................................... 3-32 3.2.1 Overview of tne LGS-SARA Accident Se ti on . . . . . . . . . . . . . . . . . . . . . . . . . . . . .....................
. .quence Quanti fi ca- 3-32 3.2.2 UNL Revisions in Quantification of Accident Sequences .. 3-33 3.2.2.1 Fire Zone 2: 13 kV Switchgear Room .......... 3-34 3.2.2.2 Fire Zone 25: Auxiliary Equipment Room ....... 3.-37 3.2.2.3 Fire Growth Event Trees for Fire Zones 20, 22, 24, 44, 45, and 47 ............................ 3-39 3.2 .3 R e vi ew R e s u l t s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-40 3.3 References to Section 3 ....................................... 3-47 i
4.0 SOME GENERAL ISSUES AND SPECIFIC RECOMMENDATIONS ................... 4-1 4.1 Sei smic Hazard and Fragility Recommendations . ... .. .. .... .. .... 4-1 4.1.1 Introduction ........................................... 4-1 4.1.2 Se i s mi c Ha z a rd . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-1 4.1.3 Se i smi c Fra gi l i ty . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-2 4.2 Fire .......................................................... 4-9 l
APPENDIX A: Detailed Review of the Quantification of the Fire-Growth i
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' Event Trees ................................................ A-1 APPENDIX B: Report of Professor Alan L. Kafka: A Critique of " Seismic Ground Motion at Limerick Generatin Rocky Mountai n , Inc. . . . . . . . . . . . . . .g Stati on , " by ERTEC
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SUMMARY
Overall, the Severe Accident Risk Assessment (SARA) for the Limerick Gen-erating Station appears to use state-of-the-art methodologies for evaluation of the core melt frequency due to seismic and fire initiating events. These re-suits are useful in a relative sense and should not be viewed as absolute numbers. The authors of SARA are well aware of the uncertainties associat&d with analyses of these events and provice discussions of the major contr1butors to uncertainties. However, it snould be noted that the prescriptions used for quantifying uncertainties in the SARA are themselves somewnat arbitrary.
The procedure used to quantify seismic risk is based on simple probabil-t istic models wnica use some data, but currently rely heavily on engineering judgement. The analysis does not include a comprehensive consideration of design and construction errors and, hence, may be (conservatively or non-i conservatively) biased.
The method used for estimating the probability distribution on frequency of exceedance for the seismic hazard is a well-established, straightforward, approach and is considered appropriate. With regard to the application of this method, it is not well defined by the coarse sampling of parameter hypotheses used in SARA. In addition, specific concerns are raised with regard to the definition and sele ^c tion of seismogenic zones and to the assignment of Seis-micity parameters. It was judged that the various issues raised with regard to the seismic hazard analysis would individually have a small impact (less than a factor of two) on the mean value of the seismic-induced core melt frequency, but that the total impact could be moderate (less than a factor of ten).
The seismic fragility analysis also was found to be reasonably within the state-of-the-art, but specific questions are raised with regard to the justi-i fication for the fragility values of various components and structures.
Simple audit calculations were performed'in an attempt to replicate the i
results given in the SARA for the mean frequency of seismic-induced core melt
. from dominant accident sequences. The simple calculations were generally in
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good agreement with tne SARA results. However, a closer examination of the event trees suggests that the Boolean algebra was not executed properly and that the seismic SARA quantification procedure is not correct. While the effect of this apparent error is small (about a factor of two reduction) for
. the total core melt frequency due to seismic events, it does alter the relative contributions from the seismic failure of components and structures. Namely, SARA determined that tne mean frequency of core melt is dominated by five electrical components in series,' wnich nave nearly the same median capacities.
On the other hand, if corrections are made to tne Boolean equations, tnen it is found that the mean frequency of core melt is dominated by contributions from structural failures ratner tnan electrical component ftilures. Full resolution of these issues would require further analyses.
In the analysis of the core melt frequency due to plant fires, tne SARA employs sta'te-of-the-art technology for the determination of fire growtn, de-tection, and suppression. In addition, tne impact of fires on plant systems is
- within the current state-of-the-art. It was found that the analysis was con-servative in many aspects, but that this is in keeping with current methodol-ogies in Onis difficult area wnica is fraugnt witn large uncertainties. Addi-tionally, it was found enat some of tne analysis, particularly the determin-istic fire growth modeling, has unphysical aspects which may be either conser-vative or nonconservative. Based on the foregoing, the reviewers believe that it would be difficult to quantify tne effect of these uncertainties, particu-larly as tney relate to probabilistic analyses.
The approach taken on tne fire analysis to the identification of critical plant areas is sound and the analysis appears to have identified all of these areas. However, in some cases, critical components, cabling, and layout of panels were not properly identified. The data base adopted for estimating the fire frequency is appropriate, but in some cases the specific estimates appear to te incorrect. The cumulative fire suppression distribution function N
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SUMMARY
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t generated in the SARA does not seem to agree with available data. 8NL obtained a distribution fit (Weibull) to the appropriate data base and thereby generated a cumulative distribution which, for any given time, yields a lower probability of fire suppression than the corresponding SARA results.
Based on tne review of probac;11stic aspects of fire initiation, growtn, '
and suppression, a limited requantification was performed of tne fire-induced core melt frequency. An overall factor of (approximately) two increase in tne fire-induced core melt frequency was estimated and this is attributed to dif-ferences in 1) the procability of fire suppression at any given time and 2) in the frequency of sel f-ignited cabl e-raceway fires. A. major contribution to the core melt frequency comes fecm the stage of fire growth in wnica all safe-snutdown systems are assumed to be damaged and faulted by the fire. Sensitiv-f ity studies were performed to examine each -of these contributors separately and they were found to be equally impo. tant. Sensitivity studies were performed
- with regard to operator error and it was found that the fire-induced core melt frequency was not very sensitive ta (one order of magnitude) changes in 1) the failure of the operator to depressurize the reactor in a required, timely fasnion or 2) in the failure of the operator to initiate required systems from a remote shutdown panel.
In the main text, this report contains recommendations for further work and information requirements in the seismic and fire areas which would be help-
! ful in assessing these risks at the Limerick plant.
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1.0 INTRODUCTION
1.1 Background
J In February 1983, Brookhaven National. Laboratory (BNL) issued a re-port (1) (NUREG/CR-3028) on its review of the probabilistic risk assess-ment (2) for tne Limerick Generating Station (LGS-pRA). The LGS-pRA excluded seismic events, fires, tornadoes, nurricanes, floods, and sabotage from tne
, set of initiating events (internal events) that it considered. In April 1983,
, Philadelphia F.lectric Company (PEco) completed a study which included tne l evaluation of risk due to seismic initiating events and to fires that mignt be initiated witnin the plant. This study, tne Severe Actident Risk Assessment for the Limerick Generating Station (LGS-SARA), also included a revised an-I alysis of the offsite consequence analysis with the CRAC2 computer code.
1 In June 1983, IRC requested that BNL undertake a preliminary, snort term review of the LGS-SARA. The review will be contained in a two-volume report.
The present d',cument is a draft of Volume 1, which reports the review of sels-mic and fire methodologies as they relate to the determination of the core melt
- frequency. Volume II will report the review of the analysis of the core melt phenomenology, fission product behavior, and offsite consequences and will be issued at a later date.
1.2" Objective, Scope, and Approach to Review The objective of this work, as, reported in Volume I, is to perform a pre-liminary review of the LGS-SARA including consideration of the core melt fre-quency. This includes an evaluation of the appropriateness of the overall
- methodology used to identify structures and components damaged and faulted due j to seismic events and fires and a comparison of PEco's methodology with current state-of-the-art approaches. In particular, this work reviews PEco's estimates of
the occurrence frequency of ground motion acceleration and the fragility analysis of structures and components damaged during seismic events; and the frequency of significant fires and the conditional failure probabilities of
- The concept of core melt frequency used here and in the LGS-SARA is equivalent to the concept of core damage frequency used in NUREG/CR-3028 (and in some t
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mitigating systems damaged and faulted during the fire. Finally, a deter-mination is made of tne influence of the findings of this review on the i
prediction of the core melt frequency as calculated in the LGS-SARA.
' It is noted at tnis point, that the determination of the impact of the finaings on the core melt frequency is qualitative in some places and, at best,
- semi-qualitative in otner places. In general, major uncertainties in ene analysis are nighlighted, subjective notions are identified, and limited re-calculations are done to focus concerns and indicat t sensitivities. A more i
detailed, quantitative reevaluation of the core melt frequency due to seismic Qvents and to fires would be a more time extensive, resource intensive enter-prise.
l This preliminary review was conducted over a two month period by BNL with j
the assistance of Jack R. Benjamin Associates, Inc. (JBA) for the seismic par-tions of tne review. The 8NL reviewers ' included J. L. Boccio (overall fire
- nazard and vulnerability review), M. A. Azarm (probacilistic fire modeling)g.
C. Ruger (deterministic fire modeling), I. A. Papazoglou (overall systems / core melt review), N. Hanan (fire / core melt review), and K. Shiu (seismic / core melt review). The JBA reviewers included J. Reed (overall seismic hazard and fragility revie'w) and M. McCann (seismic hazard review). Finally, dBA subcon-tracted with. Professor A. Kafka of Boston College for a review of the seismic hazard analysis from a seismologists viewpoint. The overall review contained in Volumes I and II was coordinated by R. A. Bari of BNL.
The review process was facilitated by several discussions and meetings
, held between BNL, NRC and PECo and its consultants (notably NUS Corporation and Structural Mechanics Associates). BNL and JBA reviewers visited the Limerick site on July 15, 1983 in order to obtain direct plant configuration information for the seismic and fire reviews.
1.3 Organization of Report Section 2.1 contains a review of the seismic hazard and fragility analy-ses. Section 2.2 contains a review of both the deterministic and probabilistic aspects of fire growth and suppression analyses. Section 3.1 contains a review
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Note that all references are provided locally in the corresponding fj sections or subsections.
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k 2.1 REVIEW OF THE SEISMIC HAZARD AND FRAGILITY ANALYSES 2.1.1 Introduction, Jack R. Benjamin and Associates, Inc. (JBA) was retained by BNL to perfonn a preliminary review of the !.GS-SARA for the effects of seismic events. The following sections of the LGS-SARA were the principal focus of the review by JBA:
Appendix A: Seismic Ground flotion Hazard at Limerick Generating Station Appendix B: Conditional Probabilities of Seismic-Induced Failure for Structures and Components for the Limerick Gener-ating Station.
Also included in JBA's review was applicable information in Chapter 3 and Appendix C.
Jack R. Benjamin and Associates, Inc., has performed similar reviews of the Indian Point Probabilistic Safety Study (IPPSS)III ,and the Zion Probabilistic Safety Study (ZPSS).I2I (See Reference 3 for the Indian Point review. The Zion review has not been published.) The review of the LGS-SARA focused on the critical issues which may signifi-cantly impact the results. Based on the experience gained from the IPPSS and ZPSS reviews, a preliminary review 'of the LGS-SARA was :;on-ducted in a short time period in order to discover the critical issues .
. and to make recommendations to address those issues which remain unre-solved. In contrast to the previous reviews which consisted of an in-depth evaluation of each section and subsection of the PRA reports, this 2-1 .
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resul ts. In order to help the reader, an effort is made to indicate, where possible, the ultimate impact of the issues which have been raised. Comments are primarily directed to the mean frecuency of core melt or to the individual secuence's which contribute significantly to core melt. Where possible, the impact of the issues raised on the median frequency of core melt is indicated. The following scale has l been adopted to quantify comments made in the review of the LGS-SARA:
Effect on Mean Fr eouency Comment of Core Melt Small Factor < 2 Moderate 2 < Factor < 10 Large Factor > 10 The methodology used in the LGS-SARA for seismic effects is appro-priate and adequate to obtain a rational measure of the probability distribution of the frequency of core melt. The results from the LGS-SARA are useful in a relative sense and should not be viewed as absolute j numbers. The procedure used to quantify seismic risk is based on simple
- probabilistic models which use some data, but currently rely heavily on engineering judgment. The analysis does not include a comprehensive consideration of design and construction errors and, hence, may be biased (note that errors may be either conservative or uncons~erva-tive). Because of the rewness of these types of analyses and the Ifmi-tations pointed out above, the results are useful only in making rela-tive comparisons. Although more sophisticated analytical models exist, the limitation of available data dictate that the simple models used in the LGS-SARA are in a practical sense at the level of the state-of-the-art.
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l 2.1.1.1 Sensitivity Analysis for Seismic Effects The approach used by NUS to combine the hazard and fragility curves is different from the method used by Pichard, Lowe, and Garrick (PLG) for the IPPSS and IPSS. In the PLG method a discrete probability distribution (DPD) approach was used to systematically account for the variability (i.e., randomness and uncertainty) in the hazard and fragil-ity parameters. Sequences were combined to form the final Boolean equations for core melt and the various release categories. System fragility data for core melt or the release categories were obtained and provided in the PLG reports for Zion and Indian Point. The combination of the system fragility curves and the hazard curves were performed l directly using numerical integration.
In contrast, the NUS approach differs frcm the PLG methodology in two respects. First, NUS included the potential for random equipment i failures and operator errors in the seismic event fault trees. Second, they used Monte Carlo simulation instead of the DPD approach adopted by PLG. It appears, based on a preliminary review, that random equipment failure and operator errors have a small effect on the mean frequency of cora melt, but may have a moderate effect on the median frequency of core melt relative to the case where only seismic contributions are incl uded.
As part of the preliminary review, an attempt was made to replicate the results given by NUS for the mean frequency of core melt as contri-buted by the significant sequences. This exercise also provided a basis for detamining the possible changes which differences of opinion could produce on the mean frequency of core nelt. The procedure used was based on the component fragility curves represented by their median values and combined variabilities (i.e., the randomness and uncertainty i logarithmic standard deviations were combined). In a.ddition, mean values for the random equipment failure and operator error events were assimed. This approach is approximate, but gives reasonable results for
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i The fragilities for the components in each of the sequences which were considered to contribute significantly to the mean frequency of core melt, were combined according to the Boolean equations and inte-grated with the hazard curves. Table 2.1 gives the comparison between the approximate values calculated as described above and the values i
reported in the LGS-SARA. In general, the approximate results compare reasonably well with the valu'es given in the LGS-SARA. The calculated mean frequency of cose melt is 6.5-6* and is within 15 percent of the LGS-SARA vaTTae of 5.7-6. The maximum difference for individual sequences is 150 percent, which is a small effect. However, the differ-ence for sequence sT RPV, which consists of a single component (i.e., the RPV), is approximately 50 percent. It was surp-ising that the calculated value was relatively different as compared to the LGS-SARA reported value (i .e. , 4.4-7 compared to 8.0-7).
Table 2.2 gives the breakdown of the mean frequency of core melt contributed by the various hazard curves. Over 86 percent is contri-i buted by the Decollement and the Piedmont, Mmax = 6.3 hazard curves, with the Decollement contributing slightly more. The Northeast Tectonic hypotheses, which is weighted by a probability of 0.3 in the LGS-SARA, contributes only about 4 percent.
Table 2.3 considers the hypothetical case that only one hazard curve exists and gives the value for mean frequency of core melt assum-ing that only one hazard curve is possible (i.e., probability weight is 1.0). This assumption is made independently for each of the six hazard hypotheses and the corresponding mean frequency of core melt values are
, given in Table 2.3 along with the ratios of values compared to the case l
l where the curves are weighted as assumed in the LGS-SARA. It is inter-l esting to note that if the Deco 11ement is the only hazard curve, the j mean frequency of core melt will only increase by a factor of 4.6, which
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I j The comparisons given in Tables 2.1, 2.2 and 2.3 give an indication of the potential sensitivity of the mean frequency of core melt to changes in the contributions from the dif'erent sequences and hazard curves.
2.1.1.2 Seismic Section Orcanization Section 2.1.2 presents the results of the review of the seismic hazard analysis, while Section 2.1.3 gives the review of the fragility analysis. Reconenendations for actions to address the significant unre-solved issues are presented in Section 4.1 of Chapter 4.
2.1.2 Seismic Hazard 2.1.2.1 Review Anproach A critical review of Appendix A of the LGS-SARA, which describes the methodology and analysis of the earthquake ground motion hazard at the Limerick site, was conducted. Section 3.3.1 of the LGS-SARA summar-izes the methodology and the results of the probabilistic seismic hazard analysis which is provided in Appendix A. To assist in the review, the services of a consultant, Professor Alan L. Kafka, were retained by JBA
,_ to review Appendix A from the seismologist's viewpoint. Professor Kafka's report is provided in Appendix B to this review, while important points are' incorporated in this body of this report.
The review of the seismic hazard analysis in the LGS-SARA has concentrated on a number of issues. To begin, the adequacy and appro-priateness of the overall probabilistic methodology to estimate the frequency of ground motion is considered in Section 2.1.2.2. Individual elements of the seismic hazard. analysis: seismogenic zones, seismicity-parameters, and the ground motion attenuation are reviewed in Sections 2.1.2.3-5, respectively.
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In Section 2.1.2.6, a preliminary assessment of overall reasonable-ness and accuracy of the LGS-SARA hazard curves is made through a
.. comparison with results derived from the historic site intensity data. s A qualitative summary of the preliminary review of the seismic hazard analysis is given in Section 2.1.2.7.
As discussed previously in Section 2.1.1, the impact of comments on the mean frequency of core melt is assessed in a qualitative manner. In addition the influence they have on the estimate of the site seismicity curves is also indicated where possible.
2.1.2.2 Seismic Hazard Methocolooy The approach used in the LGS-SARA seismic hazard analysis is well established and considered appropriate to estimate the frequency of ground shaking. M , 5) The analysis consists of two basic elements. The first step involves establishing hypotheses to model the seismicity'in the tectonic vicinity of the site and the ground motion associated with seismic events. Hypotheses are established to consider reasonable models of seismogenic zones, estimates of seismicity parameters (f.e.,
maxima magnitudes, b-values, etc.) and ground motion attenuation. For the most part, expert opinion is the principal basis for establishing the hypotheses used in the LGS-SARA. Associated with each hypothesis is a probability value that expresses the degree-of-balief that a given set of parameters is the "true" representation of the site seismicity.
The second step in the analysis involves the calculation of the annual frequency that levels of. ground motion will be exceeded at the site. This step is performed for each seismogenic zone hypothesis and the suite of likely parameter values (i.e., activity rates, b-values, maxim a magnitudes, etc.). The final product of this analysis is a family of seismicity curves, each having a discrete probability value associated with it. The discrete p'robability values s a to one, imply-ing that a complete probability distribution on the annual frequency of exceedance'has been derived.
2-7
~;_, ._ _ i- - ,_ ___ . ._. - _ _ .
l The application of this approach in the LGS-SARA is appropriate to estimate the seismic hazard at the plant site. With regard to adequacy, the application does not insure that the probability distribution on frequency has been completely defined. In the LGS-SARA study, an impli-cit decision was made only to consider those hypotheses for seismogenic zones, sources parameters, etc., that had a major influence on the estimate of the frequency of occurrence. That is, of the many reason-able hypotheses that could be considered to estimate the grounc motion hazard at Limerick, a relatively small sample was selected. In a sense, a filtering of the various parameter sets that could be included in the analysis was made. The consequences of this approach depend on the random pricess being considered. However, the consequences include the possibility that the probability distribution on frequency is defined by a coarse set of discrete probability values. Further, depending on the manner in which the hypotheses are selected, the tails of the prob-ability distribution on the annual frequency of exceedance may be poorly defined.
The approach used in the LGS-SARA presupposes that the analyst, in consultation with a seismologist, can adequately sample the space of alternate hypotheses, such that the probability distribution on frequen-cy is adaquately defined. Although the influence of individual para-meters can be reasonably estimated prior to perfoming the analysis, it is generally not true that the analyst can select a set of hypotheses that will adequately define the probability distribution on frequency l over its entire range.
In the LGS-SARA, six discrete probability values are used to define the distribution on frequency, which generally ranges over one or more orders of magnitude. This is not to suggest that a discrete representa-tion of such a wide distribution by 6-10 points Is not adequate.
Certainly, if the entire distribution were known and the points were
, selected in a prudent manner, this may be reasonable. However, in the
. 2-8
- LGS-SARA, six hypotheses and their discrete probability values were selected beforehand wi.thout knowledge of their counterpart result.on the 4
, . probability distribution on frequency. The solution to this issue is simple; a more complete sampling of the possible model hypotheses and distributions of individual parameters is needed. Specific examples where this could be achieved in the LGS-SARA are discussed in the sec-tions which follow.
In regards to the importance of having an adeyrate representation of the probability distribution on the frequency of exceedance, one point should be considered. A reliable representation of the proba-bility distribution on seismic risk (i.e., care melt), is determined fer the most part by the hazard analysis. That is, both the order of magnitude of the results and the uncertainty are dominated by the probability distribution on the frequer , of ground motion. For new plants such as Limerick, this issue becomes more important because the tails of the seismic hazard curves, which are even more uncertain, determine the estimate of seismic risk. If the seismic hazard analys.is does not adequately represent the probability distribution on frequency, results based on it may be jeopardized.
It should also be pointed out that in terms of estimating the mean frequency of core melt, the LGS-SARA results may not be influenced by the above comments. However, if the entire distribution on the frequen-cy of core melt is of concern, then these comments are more important. ;
2.1.2.3 Seismogenic Zones To model the seismic hazard at the LGS site, four hypotheses on the tectonic origin of earthquakes in the plant vicinity were defined. The definition of the different seisaogenic zones is based in part on geo-logic, geophysical, and seismic data and expert opinion. Seismicity parameters are then estimated for each zone. On the basis of expert opinion, the Piedmont, Northeast Tectonic, and Crustal Block zones were F
2.g '
)
I l
. n., ,
m e...e --+am,me-~~. ~~~ -*-me.m + ~ ~ *' *m-*" ~ ~ ' " "' ' "
, assigned probability weights of 0.30, and the Deco 11ement hypothesis was,
( assigned a 0.10 weight. Major co'ncerns with the zonation used in the LGS-SARA are discussed.
As described in the LGS-SARA, the Crustal Block hypothesis attempts to account for the occurrence of earthquakes in the northeast by the movemenc along the boundaries of large blocks of the earth's crust. It is assumed that earthquakes occur along block boundaries while the I
interior areas are relatively quiet. In the LGS-SARA, eight zones make l
up the Crustal Block hypothesis (see Figure 2.1). Of these, Zone 8 is l
the dominant contributor to the hazard at the site. This hypothesis 1:;
questioned on two accounts. First, while the principle that 1.arge blocks of the earth's crust may control the seismicity in the region along their boundaries is reasonable, such a theory should correlate reasortably well with historic and instrucentally located seismicity. In general, this is not the case (see Figure 2.1).
As stated previously, Zone 8 is reported to have the greatest contribution to the site hazard. A review of Figure 2.1 indicates that the closest proximity of Zone 8 to the LGS-SARA site is approximately' 30-40 miles. This fact alone explains to a large extent why the hazard curves derived for the Crustal Block hypothesis produced the lowest frequencies. It is further noted in Figure 2.1, which also shows the distr ibution of seismicity to 1980, that the northwest boundary of Zone 8 anpears to be inconsistent with the pattern of earthquake occurrences in southern New York, New Jersey, and eastern Pennsylvania. At the meeting at SMA, it was learned that Zone 8 was modeled to represent the Triassic Basin. The inconsistent delineation of Zone 8, with respect to local seismicity patterns, may be attributed to two factors. The LGS-FSAR (Ref. 6) reports that Limerick is in the Triassic Lowlands, suggesting that the northwest boundary of Zone 8 should be moved toward the plant. This would also be consistent with the distribution of seismic events in the region (see Figure 2.1).
._ 2-10 n
- ~ _ .
l
i i
l l
Secondly. it is not apparent that the boundaries of seismogenic
' i zones should be coincident with the perimeter of a large geologic struc-y ture. If in fact these boundaries generate seismic events, it may not be realistic to restrict their occurrence to the boundary itself.
- Instead, events should be modeled as occurring in a volume of crust, defining a zone of weakness., In one sense, this has been doce for Zone 8 towards the southeast.
~
t A redefinition of Zone 8 in the Crustal-Block hypothesis that 4
places the LGS site within its boundaries is judged to have a moderate impact on the estimated hazard curves (i.e., at least a factor of 2).
The consequences of this change on the mean frequency of core melt is estimated to be small (i.e., a factor of 2 or less).- However, a moder-ate increase in the median ccre melt frequency is considered possible.
To consider the possibility that large magnitude events could occur in the northeast, the Deco 11ement source zone was defined. A maximum magnitude of 6.8 was assumed, and a probability weight of 0.10 was assigned to this hypothesis. The selection of maximum event size is discussed in Section 2.1.2.4. The Deco 11ement hypothesis is one of a number of theories being considered by seismological experts to explain t
the possible occurrence of large magnitude events in the eastern U.S.
The physical basis of this hypothesis is the identification of a shal-low-dipping reflector beneath and along the east coast that has been 4
interpreted as a seismically distinct block of the earth's crust. p, 8)
A major concern with t1e Deco 11ement hypothesis is the fact that patterns of instrumentally located seismicity do not correlate well with j .i t. That is, fault plane solutions and source depths do 'not suggest i that earthquakes in the region of Charleston, South Carolina, or any-i
'where else along the eastern seaboard occur on a decollement surface.
In addition, since the evidence that a major decollement may exist 4
generally appifes to the southern Appalacians, it is not clear that a decollement seismogenic zone should extend to the northeast in the vicinity of the Limerick. site.
~
2-11 i
l
-. .:_ L v7 : - ~~T - _ . - . - J. :T :X T
i e
At the SMA meeting, discussions with Dr. McGuire revealed that the Decollement hypothesis was not selected solely on the basis of physical I arguments that expl' a ins the seismicity in the east. A principal motiva-tion was its use as an all-inclusive hypothesis, in a probabilistic sense, in that it allows the possible occurrence of events as large as M6.8. That is, an assumption is made in the LGS-SARA that all reason-able hypotheses watch would consider the possibility that large-magni-tude events could occur in the vicinity of the plant site are fully represented by the Decollement hypothesis. Although such an approach may provide a t,est estimate of the ground shaking hazard at the LGS site, it is not clear that it is appropriate or adequate for use in the LGS-SARA. No basis is provided to support the belief that the l Decollement hypothesis in fact adequately represents, even in a best
! estimate sense, the hypothesis that large events can occur. Also, the variability in key parameters was not considered in the Deco 11ement hypothesis (i.e., b-values and Mmax). Neither is it clear that the Decollement source zone is the most apprcpriate way to model the occur-rence of large magnitude events in the eastern seaboard.
The use of decollement tectonics to explain the occurrence of large l
i magnitude events in the east is one of many theories based in part on r.
scientific evidence and expert speculation. Although experts differ as to the validity of any theory to explain the 1886 Charleston, South Carolina, earthquake or the occurrence of future large events, the Deco 11ement source zone is certainly one that could be used. However, in the LGS-SARA the Decollement zone serves as a single physical charac-terization of the process that generates large-magnitude events as well as a summary of a multitude of hypotheses that define other physical processes. It is with this expanded role that a concern is raised.
A number of alternatives exist to model the occurrence of large--
magnitude events in the east. Among the possibilities is to allow the occurrence of M6.8 events in the other source zones defined in the 2-17.
4 p%
s
,,,g s
(...
--*=g*^=9" " " * ' " * * * * * * ' " " ^ *
- _ _ _ ~_9*= . _ . _ '____ . _ _ - - - - - - - - -- -
O 4-1 LGS-SARA. That is, an Mmax = 6.8 would be considered as one hypothesis on maximur. magnitude for each source zone. The basis for this approach is straignt-forward. The occurrence of large-magnitude events in the east is considered possible on pre-existing zones of weakness in the
, earth's crust. What defines these zones as earthquake generators va ry. In part a variety of such theories are the basis of the seismo-genic zone and hypotheses in the LGS-SARA (i.e., Piedmont, Northeast Tectonic zones, and Crustal Bloch). The concept of pre-existing zones of weakness is consistent with the thinking expressed by the four experts in Appendix B to the LGS seismic hazard analysis. Furthe rmore, a preference was given in the hazard analysis to the Piedmont, Northeast Tactonic zones, and Crustal Block hypotheses. A combined probability j weight of 0.90 was assigned to them. A 0.10 probability was given to >
l the Decollement hypothesis. Consistent with this degree-of-belief and the consensus in Appendix B that large earthquakes can be expected on pre-existing zones of weakness, the possibility of large-magnitude events in source hypotheses that define such zones, should be consid-ered. This approach was discussed at the SMA meeting with Dr. McGuire, and recognized by him to be a reasonable alternative to model the occur-rence of large magnitude earthquakes. However, it is the opinion of Dr.
McGui'e r (and but not necessarily the consensus of all the consultants) that the total probability weight assigned to any and all hypotheses is 0.10.
The question 'as to whether 0.10 probability is a reasonable value to be assigned to the hypothesis that large magnitude events (i.e.,
M6.8) can occur in the vicinity of the Limerick site is a difficult' question and one that must be answered on the basis of expert opinion.
In Appendix B to the LGS seismic hazard analysis, the four experts interviewed ag:eed universally that such events could occur at the LGS. The degree-of-belief assigned to such a hypothesis varied from zero to twenty-five or thirty percent. Presumably the value of zero is actually a very small number, otherwise there could not have been the 2-13
~-
7_....... -.
7, , :-.:..----.- -,,.-w.- ,; w ~, -- ,mm - -+e==~~~~~----
l
- 9 aforementioned universal agreement. At this point in the preliminary
{ review of the seismic hazard analysis, the value of 0.10 is not accepted i
l by JBA nor all the experts retained in the LGS-SARA. Qualitatively, this value should be considered a lower bound.
I The alternative approach suggested to model large-magnitude events would produce at least one additional hazard curve for each source zone. By virtue of the arguments on maximum acceleration, these addf-tional hazard , curves would be unbounded as is the Deco 11ement zone.
Depending on the source considered, the impact on the frequency of ground motion varies. However, it is felt that in most cases the hazard curve associated with a large-magnitude event will be higher by a factor j of 2 or less, ccmpared to the existing hazard curves. At higher accel-l erations, thase new curves will be unbounded and thus have nonzero i
occurrence frequencies, unlike the previous hypotheses.
i
[
With respect to their impact, the fact that these additional curves l are unbounded means that they will make a greater contribution to the i mean frequency of core melt than their counterparts for each source .
zone. Previously, the Piedmont, Nax = 6.3 and Deco 11ement hypotheses contributed 86 percent of the mean core melt frequency, since they only allowed accelerations greater than 0.80g to occur. All zones will have
- scoe contribution to the mean frequency of core melt. The overall influence of these additional curves is judged to result in a small t
increase in the mean cort melt frequency.
2.1.2.4 Seismicity Parameters For a prescribed zone of seismicity, the random occurrence of earthquakes is defineit by the seismic activity rate, the Richter b-value, and the maximum magnitude that can be generated by the sour,:e. Estimates of seismic activity are based on the historic record. However, the statement that seismic activity rates are well determined in the eastern U.S. is in some ways an overstatement or at 2-14
, _ , ._m,, ,. ..m.-- --
- -*====- ~ ' ~
%e =.e---r -mw - w -- d- - v- - m w e =- -%we wvs a w
- s l
l least easily misinterpreted. For a prescribed area in the east, the catalog of earthquake occurrences is generally believed to be long j enough and sufficiently complete that estimates of activity rates are reasonably well determined. That is, their uncertainty is low enough that its impact on the frequency of exceedance of ground motion can be ignored. However, from the point of view of the rate of seismic activ-ity per unit area (i.e., say 104 km2) the variation can be large.- From Table 2 in the LGS-SARA hazard analysis, the rate of seismic for the four source hypotheses varies from 4.33 to 38.0 x 10-3 events per-year, per-104 km2. This effect is taken into account in the LGS-SARA, however this variation per se is not recognized as such, In the LGS-SARA, the estimate of Richter b-values was based solely on expert opinion as reported in Reference 9. A best estimate of 0.90 was used for all source zones, and no uncertainty was considered. In Reference 9, the experts came to a consensus that 0.90 was a realistic, albeit default value that can be used for all seismogenic zones in the eastern U.S. However, it was further stated by many of the experts that it is believed that b values for different seismogenic zones may vary-from 0.90 as a best estimate. This notion suggests that variability in '
the mean value of b exists. That is, a difference exists between the 0.90 global estimate, and the true best estimate for a given source zone. In fact, some experts indicated a preference for a regional
, dependence for b-values. Furthermore, there is the contribution of statistical variability in b-value estimates derived from the data, which depends on the number of data points. Thus, as a minimum, two sources of variability exist in the estimata of Richter b-values:(1) a possible bias in the use of 0.90 best estimate value recommended by experts for all source zones, regardless of the actual distribution of the data and (2) the statistical variability due to limited sample size. The failure to account for the variability in b-values is an example of the inadequate degree to which parameter hypotheses have been sampled in the LGS-SARA. It should be noted that the LGS-SARA did not 2-15 l
l .
l l
_~ _ _
~ ' ~~
j' .. . l_ ..[ _
1 .
directly estimate Richter b-values from the catalog of earthquake occur-rences. In considering the estimate of b-values, PECo should consider the results obtained using the historic data.
I l
The impact of a complete characterization of the variability in b-values on the mean core melt frequency is judged to be small. ,
The final seismicity parameter defined for a seismogenic zone is the maximum magnitude. In the previous section, the manner in which
, large magnitude events were modeled in the LGS-SARA was considered.
! Here, the matter of what the size of the largest events should be assumed is addressed.
The estimate of maximum magnitudes for the Piedmont source zone reflected the issue of the 1982 New Brunswick, Canada event and the Cape Ann earthquakes. The magnitude 5.7 New Brunswick event is used as the basis for establishing the distribution on M,,x, while it was stated that the Cape Ann earthquakes do not belong in the Piedmont zone. The basis for limiting the occurrence of the Cape Ann events to New England is presumably related to the theory that a Boston-Ottawa seismic belt '
exists as discussed in References 10, .11 and 12. Hcwever, the existence of such a trend does not correlate very well with results of recent studies questioning the existence of a such a trend.(12) Thus, no definitive basis exists to support the hypothesis of a Boston-Ottawa seismic belt and therefore no reason exists to exclude earthquakes near Cape Ann, from the Piedmont region. This is further supported by the arg ments provided in the LGS-SARA that suggest the 1982 New Brunswick, Canada, earthquake belongs in this seismic province.
If the 1755 Cape Ann earthquake is considered to be a 6.0 event,I14I the distribution on M would be modified to reflect the aax fact that the largest observed event had a magnitude of 6.0 as opposed l
to 5.7. If it is assmed that the two point distribution on M,,x was changed from S.8 and 6.3 to 6.0.and 6.5, it is estimated that the effect
- on the frequency of exceedance curves and the mean core melt frequency l would be small.
l
. 2-16 ~
- - , ,-n- - - ,_ -
l l
The hypothesis that a large-magnitude event, the . size of the 1886 1
Charleston, South Carolina, event could occur on the eastern seaboard was considered in the Decollement source zone. A magnitude of 6.8 was assigned to this Modified Mercalli Intensity (MMI) X event. No basis is provided in the LGS-SARA to support the implicit assumption that the observed magnitude of the Charleston event is the maximum event that could occur. Should it for example be considered a lower bound on
'%ax? This question and the uncertainty in 'hax should be addressed _
by PEco.
2.1.2.5 Ground Motion Attanuation To describe the attenuation of ground motion with magnitude and distance, Nuttlf's relationship for sustained acceleration was used.(15) The uncertainty in ground motion predictions is described by a lognormal distribution with a standard deviation of 0.60. This stand-ard deviation value corresponds to a factor of 1.8 times the median val ue.
The attenuation relationship was modified in the hazard analysis to predict sustained-based peak acceleration and to account for the random orientation of grcund motion. This factor is magnitude dependent.
Above magnitude 6.0, sustained-based peak acceleration is 1.23 times sustained acceleration. The attenuation model used in the LGS-SARA is appropriate and adequate to describe the ground motion at the plant site. -
The prediction of ground motion in the eastern U.S. is a difficult task due to the Ifmited strong motion data available for that region.
However, a number of relationships have been developed and used in probabilistic hazard analyses.Il* 9) Results of sensitivity studies are available to compare the impact of various functions on the estimated hazard curves. A preliminary review of these studies suggests the attenuation for sustain-based peak accelerations used in the LGS-SARA is 2-17 -
1 t
. . . . . . . - , - - - - -- - - - ~ - ' - '
_ _ _ . _ _ _ . _- . - - ~.~ ~ ~ ~ ~ .
i s .
generally on the conservative side (i.e., it g1ves higher accelerations at a given-frequency of exceedance level).(9) It is noted however that there can be considerable variation in the hazard analysis results for various attenuation relationships. This suggests that a more compre-i hensive sampling of attenuation functions is appropriate, since it is generally believed that the capability to predict ground motion in the eastern U.S. is not well established. The impact of including alterna-tive attenuation hypotheses on the mean core melt frequency is consid-ered to be small.
2.1.2.6 Comparison of the LGS Hazard Analysis with the Historic Seismicity The accuracy of the LGS-SARA seismic hazard analysis might be compared with the historic distribution of earthquake ground motion l experienced at the plant site. However, since a record of the ground shaking intensity at the LGS site is not available, another approach must be taken. In the Limerick FSAR(6) the earthquakes that have occur-red since 1737 within 200 miles of the site (Table 2.5-2, Reference 6) are reported. These data provide a basis to estimate the distribution of historic ground motion. The approach used to do this is summarized below.
The catalog of earthquake occurrences provided fit the FSAR describes event size in terms of Modified Mercalli Intensity. To estab-lish a distribution of the m intensities experienced at the LGS, the *
,, reported epicentral intensities are attenuated to the plant. This is done using the intensity attenuation relation in Reference 16 for rock sites given by the following equation.
l Is*Io + 2.6 - 1.39 inR (2.1) 2-18
~
~
7 --- - ..t.. ~~ ~ ~ ~
d l
where: Is = site intensity
-In = epicentral intensity R = distance (miles)
For each event and distance reported in the FSAR, a site intensity was estimated using Equation 2.1. In establishing a record of the MMI level experienced at the LGS site, no attemot was made to verify the cataloo reported in the FSAR or to correct the record for inconsistencies.
Also, no uncertainty in the estimate of site intensities was consid-ered. Intensities above MMI ecual to IV are considered.
To define the distribution of seismic intansities at the site, the Gutenburg-Richter relation that describes the number of events versus intensity is given as follows:
log 10N (I ) = a + bI S 3 (2.2) where a and b are parameters fit to the data. The b term is known as the Richter b-value. The b-value on intensity is estimated to be
-0.72. The seismic activity rate for events of MMI > IV is 0.0266 events per year based on a 226 year record.
An estimate of the historic ground motion in terms of ground accel-eration can be obtained by a transformation of intensity to peak ground acceleration using an appropriate relation. To do this, the following equation was used:(17) log 10A = 0.014 + 0.30Is (2 3) where A is peak ground acceleration in cm/sec 2 . To account for the '
uncertainty in estimating A in Equation 2.3 and the uncertainty in
, attenuating intensity in Equation 2.1, a lognormal distribution on peak acceleration is assumed, with a logarithmic standard deviation of 0.28 .
(base 10), which corresponds to a factor of 1.9.
2-19
..,4-,e,,yeg<*h- * ' " ' '
- The distribution on acceleration at the LGS is estimated according to v(A >a) = v f(I) al P(A > a l I) (2.4) where v(A > a) = annual frequency of peak acceleration A, greater than a value a.
v =
seismic activity rate for intensities greater than or equal to IV.
f(I)*a I = doubly truncated exponential distribution on intensity I with parameter b in 10 where AI is the increment on intensity.
P(A>alI) = probability of peak acceleration A greater than a, given an intensity I. This is described by a legnormal distribution whose median is defined by equation 2.3 with a logarithmic standard deviation, of 0.28 (base 10).
The result of this computation, using Imax of VI, is shown in Figure 2.2 with selected curves from the LCS hazard analysis. The historic seis-micity curve ranges at accelerations around 0.10g from the results ,
obtained for the Decollement and Piedment zones to the lower frequencies estimated by the Crustal Block zone. These observations suggest that the overall frequency of events producing accelerations of 0.10g is reasonably well described by the Decollement and Piedmont zones and Crustal Block zone, M = 6.0, to within a factor of 2. Since the maximum intensity falt at the site is MMI VI, the historic frequency curve falls off sharply.
l Equation 2.4 can also be used as a prediction tool by allowing the possibility of site intensities greater than VI to occur. To do this, an estimate of the maximum site intensity that can occur must be made.
2-20 m -
_ l
)
l This is the same step that is taken in the probabilistic seismic hazard analysis. A maximum intensity of X is assumed, which corresponds to a large-magnitude (=M7.0) event occurring very near the site. The result of estimating f(I) in equation 2.4 and calculating v(A > a) for a maximum intensity of X, is also shown in Figure 2.2. This assumation a? lows the possibility of high accelerations associated with large events to occur. In general, the site intensity curve tracks the trend of the Piedmont and Decollement seismicity curves quite well.
As a final estimate based on the historic distribution of ground motion at the LGS, a seismicity curve is estimated assumino a Richter b-value of 0.45 which corresponds to the 0.90 value used for the magnitude scale in the LGS-SARA. Also, a maximum intensity MMI X is assumed. The hazard curve for this case is also shown in Figure 2.2. The effect of assuming a b-value of 0.45 (equivalent to 0.90 for the magnitude scale) is a factor of four increase in the hazard.
The results based on the historic-site intensity distribution agree reasonably well with the seismicity curves derived in the LGS-SARA.
From the point of view of prediction, if a maximum site intensity of X is postulated, the Piedmont. and Deco 11ement zones agree most closely with the historically derived curve. The same could be said for the Northeast Tectonic zone, expect. that the truncation on peak acceleration produces ~a sharp fall-off at 0.30g. ~
2.1.2.7 Summary .
The previous sections provide the results of a preliminary review of the LGS-SARA seismic hazard analysis. The adecuacy and appropriate-ness of the analysis approach were considered. The appropriateness of individual technical aspects of the analysis were also reviewed.
The methodology used to estimate the probability distribution on frequency of exceedance is considered appropriate to estimate the seismic risk due to nuclear facilities. The method used in the LGS-SARA 2-21 S
i
_ , _ . , _ . . . - , ~ e ---- -~*-- ' ~ ~
~~~~ ~ ^
is a well established straightforward approach to estimate the ground shaking hazard. With regard to the adequacy of the way the method was.
applied, it is felt that in principle the estimation of the probability distribution on frequency is not necessarily well defined by the coarse sampling of parameter hypotheses used in the LGS-SARA. The approach used in the LGS-SARA was to select six hypotheses, each with an assigned probability weight. It was then assumed that the six hazard curves generated, fully define the probability distribution on frequency.
Although a best estimate can be obtained in such a manner, this approach
-~
does not insure that the probability distribution on frequency will be adequately represented.
With regard to seismogenic zones, two major concerns were raised.
First, delineation of the boundaries of the Crustal Block hypothesis was questioned. In particular, Zone 8 in this model was considered inappro-priately defined to be approximately 30 miles from the LGS at its closest point. The impact of redefining Zone 8 on the mean frequency of core melt was considered to be small. Secondly, the Deco 11ement source was used as an all-inciasive model to consider the general hypothesis that large-magnitude events can occur in the east. This approach was not considered to be the most reasonable means of evaluating the hazard due to such hypotheses. An alternative was recommended that allows the possible occurrence of large-magnitude events to occur on the other i
source zones as well. The impact of this alternative on the mean core melt frequency was considered to be small.
With regard to seismicity parameters, two issues were raised. The first deals with the assignment of Richter b-values. The LGS-SARA uses a single b-value for all source zones. The basis for this was expert opinion. No uncertainty in b-values was considered. This approach was not considered appropriate, rather, a distribution on b-values should be used since there exists a source of bias in the best estimate of the b-value for each source zone, as well as statistical uncertainty. The 1
2-22 e
4-
l impact of not considering the uncertainty in this parameter is consid-
,, ered to be small.
Particular concern was expressed with regard to the estimate of maximum magnitudes. For the Pfedmont source, evidence was presented 4
that questioned the basis for establishing the distribution on maximum cagnitude. Specifically, the Cape Ann events should be included in the Piedmont province and considered in the estimate of Mmax. D e overall impact on the mean core melt frequency is considered to be small.
The possible occurrence of large-magnitude events (=M7.0) was considered in the Decollement source hypothesis. The 1886 Charleston, South Carolina event was estimated to have a magnitude of M6.8 in the LGS-SARA and was used as the basis to estimate the largest event that could occur. No uncertainty in thf s estimate was considered, neither was there any basis for this hypothesis.
In a preliminary assessment of the hazard analysis results, the frequency distribution of ground motion due to historic earthquakes.was computed. Generally, the results from the analysis of the historical data suggest that LGS-SARA study results are reasonable. Hazard curves that include the possibilf ty of an intensity X event are consistent with the hazard curves estimated for the Piedmont, Deco 11ement, and Northeast Tectonic zones at low accelerations.
The recommendations given in Section 4.1.2 c.e directed towards resolving the issues summarized above. Although the effect of the individual issues on the mean frequency of core melt is judged to"be small, the total effect could be moderate.
2.1.3 Seismic Fragility The preifminary review of the sef smic fragility parameter values focused on Appendix B of the LGS-SARA and included a review of those
~
portions of Chapter 3 and Appendix C pertinent to the se'smic risk 2-23
.J l
t l
i 4 -, . - y-- .--- - -
- ~ - - ~ ~ ~ ~ - ' '
analysis. As described in Section 2.1.1, the results of the meeting with SMA and the plant tour helped direct the review effort to the critical components and issues. In addition, the calculations for the significant contributors in Table 3-1 of the LGS-SARA were obtained and studied. The fragilities for other components were considered in rela-tionship to their potential impact on the mean frequency of core melt.
For example, the median capacity of the batteries and racks is reported to be 2.56g and, thus, was not included in the sequences. This compo-nent was inspected during the plant tour, and its capacity value is judged to be reasonable.
The coments concerning the seismic fragility analysis are organ-ized in a manner to highlight the concerns, which were either most potentially critical or which were the most controversial during the review. Sections 2.1.3.1 through 2.1.3.6 discuss this category of concer ns. Section 2.1.3.7 presents the results of the review of the calculations for the significant components. Many of the concerns found during the review of the calculations are also discussed in detail in .
Sections 2.1.3.1 through 2.1.3.6. Section 2.1.3.8 addresses general fragility-related issues which should not be overicoked, but which are philosophical in nature (i.e., do not have an imediate resolution) or
_ whicn are unlikely to have a major impact on the results. Finally, Section 2.1.3.9 gives final closing comments on the preliminary review of the seismic fragility analysis in the LGS-SARA.
Throughout the discussion recommenda'tions are made for additional information. Section 4.1.3 summarizes the recomendations for addi-tional actions required to resolve the fragility-related issues which have been raised but not answered.
2.1.3.1 D mage Factor Three adjustment factors are used in the LGS-SARA to estimate capacity to resist earthquakes. The hazard analysis documented in 2-24
_ _ _ _ _ _ _ _ , . .- - - - - -- * ~ ~ ~ ~ - - * * ' ~~ ' ~ ~~~~
g S m , ---g- , -- _ y ,,- -
. 4 I
Appendix A of the LGS-SARA presents the frequency of exceedance for seismic hazard in terms of a sustained-b3 sed peak acceleration para-meter. As explained in Section 3.3.1 of the LGS-SARA, the accelerations from the Appendix A hazard curves were scaled by a factor of 0.81 (i.e.,
1/1.23) to convert the sustained-based peak accelerations to effective peak accelerations .to reflect the less damaging characteristics of low magnitude earthquakes. This adjustment is identical to the adjustments 1 made in the IPPSS and the ZPSS. As explained in Reference 18 (Reference 18 was provided to the reviewers by PECo to support the LGS-SARA), this factor was conservatively selected to account for smaller nonlinear res- i pense and, hence, damage caused by lower magnitude events. It is im-plied in Reference 18 that the adjustment factor should be 0.5 for mag-nitudes less than MS and distances less than 20 km. For magnitudes greater than M7 and distances greater than 40 km, the adjustment factor is unity.
A second factor was introduced in the LGS-SARA which is discussed in Section 4.1.3 of Appendix B of that report. This factor is called an earthquake duration factor, which is used to increase the median capac-ity of structures by a factor of 1.4. The justification for this factor as discussed in Section 4.1.3 is very similar ta the justification for the hazard reduction factor (i.e.,1/1.23) described above; thus, it is concluded that these factors account for the same phemonena and only one factor should be used. Note that the duration factor of 1.4 was not included in the IPPSS and the ZPSS.
This apparent discrepancy was discussed at the meeting held at SMA,
~
and it was explained by SMA that for future PRAs only the 1.4 factor will be used and no adjustment will be made to the seismic hazard l
curvesi In defense of the LGS-SARA analysis, SMA explained that very low ductility values had been used in the development of the ductility factors for Limerick (f.e.,'2.0 for shear and 2.5 for flexural failure of concrete walls). The ductility factor is the third adjustment factor-used in the LGS-3 ARA. More realistic values of 3 to 4 for the ductility
, 2-25
- - ~ ~ - - - - - - - ~ ~ - - - - -
e ratio should have been used. The use of low ductility values compen-sated for the extra 1/1.23 factor used to adjust the hazard curves for structures. The 1.4 factor was not used for equipment which generally had realistic ductility values. In conclusion, if only the 1.4 duration factor and realistic concrete ductility values had been used for the structures, the results would have been essentially the same. The reviewers concur with this explanation.
The justification for the duration factor of 1.4 was also reviewed. The underlying basis for the duration factor is recent work reported in Referenc1t 19. As documented in this report, a series of analyses were conducted to investigate the response of' single-degree-of-freedom (SDOF) nonlinear oscillators to real earthquake motions. Earth-quakes which varied in magnitude from M4.3 to M7.7 were used. It was explained at the meeting at SMA that a ' duration factor is required to correct the capacity of SDOF systems when subjected to earthquakes less than M6 to obtain the same level of damage.
The ductility factor based on the approach developed by Riddell and i
Newmark,(20) which was used in the LGS-SARA, assumes earthquakes larger than M6. Since this method is used to develop the ductility factors for structures, a duration factor (really a magnitude factor) was applied for events with magnitudes less than M6. An analysis was conducted by SMA using the data from Reference 19, where the response of the non-linear SDOF oscillators to earthquakes less than M6 to events greater than or equal to M6, were compared. By fitting a lognormal distribution to the ratios of the response factors for these two groups of events, the median adjustment factor of 1.4 was determined. In the LGS-SARA this factor was applied for all hazard curves, which implicitly assumes all earthquakes have magnitude less than M6.
- l In an effort to verify the earthquake duration factor used in the l
LGS-SARA fragility analysis, the data contained in Reference 19 was reviewed. As described above, argtments which support the use of an 2-26 I
- - - - r-- -n--
~
1 earthquake duration factor are based on the assumption that sef smic events of magnitude less than 6 contribute to the likelf hood of failure. It was on this basis that the median value of 1.4 was derived for use in the LGS-SARA. As a check, the data in reported in Table 4-1(a) for u= 4.27 in Reference 19 were considered in two groups: M<6 and M>6. The artificial time history was included in the M>6 grouo.
From the histogram for each group the median response factor and .
logarithmic standard deviation were cerived. Then, the ratio of the response factors was determined and compared to the LGS-SARA values. A summary of the estimates made are given below.
Resconse Factor Data Grouc F, 8 M<6 2.65 0.25 M>6 2.15 0.26 ,
FED
- FM<6 /FM)6 1.23 0.36 LGS-SARA . 1.40 0.20 Sc 0.12 S p 0.08 Su From this comparison, it appears that the median factor used in the LGS-SARA is over estimated by 14 percent (f.e.,1.23 compared to 1.40). It should be noted that including the artificially generated time history in the M)6 group has a negifgible effect on the median.
A second look at the scale factor data was taken by dividing the data in short and long duration (TO) groups. The data were divided according to whether durations were less or greater than 2.5 seconds, as defined in Reference 19.
2-27 s-
.%_ .. ._. . .- w--- == *= - * = * * * " - * - '*** ' " ' * ' " * * * * " ' - '
y - - - - - -
g --y +
In this case the artificial time history is in the T >2.5 D second group. Basically all the records in the M)6 group were in the T D>2.5 second data set with one exception. The UCSB Goleta recording of the MS.1 (M3 5.6) 1978 Santa Barbara earthquake had a duration of 3.0 seconds, and thus was included in the long duration subgroup. The results for these data sets is given below.
Resoonse Factor Data Group S F_
7042.5 sec. 2.85 0.51 TO )2.5 sec. 2.05 0.26 FED = FTD'2.5/FTD>2.5 1.39 0.57 LGS-SARA 1.40 0.20 Sc 0.12 S r 0.08 Su From this comparison, it would seem that in deriving the duration factor, that a duration, rather than a magnitude criteria was used.
This is inconsistent with the application in the LGS-SARA. Possibly of greater significance is the fact that a single earthquake record produced a variation in the estimated median duration factor from 1.4 to 1.23. This muld seem to point out that, although Reference 19 provides a clear indication of the duration effect of strong motion on structural damage, it is a study limited in its appifcation because of the rela-tively small data base. As discussed at the meeting with SMA, the use of the M6 cutoff to establish the duration factor is a gross character-ization of a process that is continuous over magnitude and/or duration. Thus, a median duration factor should preferably be a func-tion of magnitude. Data to establish such a function are not avail-abl e. Furthermore, Reference 17 also suggests that the duration factor m
2-28
]
M<*w ""* * ' * ' * ' *
. .o .
4 has a frequency dependence. This is not taken into account in the
~
LGS-SARA.
l The estimate of the logarithmic standard deviation of the duration factor in the LGS-SARA appears low. In particular, due to the uncer-tainty in estimating FED and the limited data base, s u = 0.08 is low,
, and in any case should not be lower then the randomness component.
Direct estimates of the variability in FED ranged from 0.36 to 0.57.
Values of S c of this size are considered more appropriate.
In principle, incorporating the effects of duration in the estimate of seismic capacities is appropriate. And although the results reported in Reference 17 are consistent with engineering judgment and observed earthquake damage, the approach used in the LGS-SARA is a simplification of a ecmplicated issue.
The arguments leading to the 1.4 duration factor, when included with the ductility adjustment factor based on Reference 19, are gener-
~
ally reasonable for earthquakes with magnitudes less than M6; however, as discussed above, the 1.4 factor may be slightly high and the '
uncertainty estimate low. For events greater than M7 it was agreed by SMA that the duration factor should be unity. Between magnitude M6 and M7 events the data in Reference 17 do not support a duration factor of 1.4 in the opinion of the reviewers. If the duration factor of 1.4 is changed to 1.0 for structures and equivalently the hazard curve adjustment of 1/1.23 for equipment is also changed to 1.0, for the region of peak-sustained accelerations corresponding to average magnitudes greater than M6.0, the frequency of core melt distribution will be affected. Note that the ductility values used for equipment are generally realistic, hence the 1/1.23 hazard curve factor is analagous to the 1.4 duration factor used for structures.
Based on Reference 21, the hazard curve for the Decollement seismo-genic zone is the only curve which has average magnitudes equal to or greater than M6.0. For sustained-based peak accelerations equal to or
~
greater than 0.40g, the average magnitudes equal or exceed M6.0. It is I-
... 2-29
~.
_ . . . .-~.4 -
-- -. --- v +~ ** * - '~ ' ~ ' ' * - * -
- a e estimated that if the duration factor is changed to unity for this region of the Deco 11ement hazard curve the mean frequency of core melt will increase by a factor of approximately 1.4. The effect of this adjustment will not significantly affect the median core melt frequency.
2.1.3.2 Upper Bound Accelerations All the hazard curves, except the Decollement case, are truncated l to reflect the belief that maximum accelerations are associated with each seismic hazard hypothesis. The argument leading to the limiting acceleration values is documented in Reference 18, which was provided to i the reviewers by PECo to support the LGS-SARA. This is the same argu-ment which is given in the IPPSS and ZPSS reports (1,2) for limiting j accelerations. The explanation for limiting upper-bound accelerations
{ consists of two steps. The first step is the assumption that there is a l maximum intensity associated with each source zone corresponding to the maximum magnitude for that zone. This is assumed to be true by sef s-mologists. The second step relates the predicted accelerations for masonry structures with the qualitative descriptions of the MMI scale.
The basis for the argument leading to maximum acceleration values in the second step is as follows. Masonry structures are selected since they are the only engineered components for which damage is systemati-cally described in the MMI scale. If the accelerations are higher than predicted, then a higher MI value (corresponding to more damage) would occur. 4towever, since the maximus MMI values are Ifmited by the seis-l mologist, a higher acceleration is not possible. The problem with limiting accelerations for the Deco 11ement hazard curve is the assigned maximum magnitude value of M6.8 which corresponds to a maximum intensity of approximately MMI X. This intensity is associated with failure of most masonry structures; thus, the argument cannot be used since all higher MI values also include failure of most (if not all) masonry structures. As explained at the meeting at SMA, it was conservatively decided not to truncate the Deco 11ement hazard curve.
2-30 -
s
_, , ,_ _ _ _ . . _. - . - . * . * **~
-v
i t
It also follows directly that if upper bounds on intensity exist then upper bounds on damage exist since intensity is a scale which measures damage. Although it is believed by the reviewers that it is -
more appropriate not to truncate the hazard curves but to reflect a limit on damageability in the fragility curves, the effect of modifying i
the hazard curves produces the same result. Thus if upper bounds exist for lower intensity values, similar limits should apply for higher
!' intensity values for engineereo concrete structures. However, it is difficult to quantify this belief at this time. In conclusion, the assumption not to truncate the Deco 11ement hazard curve is on the con-servative side.
Based on the approximate analysis described in Section 2.1.1, the effects of truncating the Decollement hazard curve were investigated.
It was found that when truncating the curve at 1.0g (which represents a reasonable lower bound) the mean frequency of core melt will change by a factor of approximately 0.85. The effect on the median frequency of core melt is expected to be very small. Thus, it is concluded that truncating or not truncating the Deco 11ement hazard curve has a small .
effect on the results of the LGS-SARA.
2.1.3.3 Reactor Enclosure and Control Strxture The median capacity of the reactor enclosure and control structure is reported in the LGS-SARA to be 1.05g (see Table 3-1 in the LGS-SARA). The structural calculations for this component were reviewed.
The reviewers believe that the capacity of the walls is rationally represented by 0.90g, which is based on the total capacity of the walls in the north-south direction between elevation 177 feet and 217 ' feet.
This capacity is based on the capability of the floor diaphragm at elevation 217 feet to redistribute. forces. At the meeting with SMA, it was stated that the diaphragm capacity for the Susquehanna plant was checked in detail and since the Limerick plant is structurally the same, 2-31 s_
., . ww %.4 .e*--=+--** * * * - * ~"
' ~ ~ '
g y * -y g y- - , -,- c- + - - - -y
,"mN^'y- 7-r& -, .e - w-y- g
- ~ -
the diaphragm capacity is adequate to redistribute forces as the various wall sections yield. -
Based on a median capacity of 0.90 9 , it is estimated that the mean frequency of core melt would increase by a factor of approximately 1.2.
2.1.3.4 Reactor Pressure Yessel Capacities Three of the significant earthquake-induced failure components listed in Table 3-1 of the LGS-SARA are associated with the reactor pressure vessel (RPV) which is located in the containment structure. In the development of the median capacity values for the reactor internals, RPV, and the CR0 guida. tubes, it was assumed that the containment struc-ture had an effective damping value of 10 percent. Since the original analysis of the ccmbined containment /NSSS was based on 5 percent damping i for the concrete structure, a 1.3 factor, which increased the capacity i
of the RPV components, was developed from the ground spectral accelera- ,
tions by SMA.
It is not obvious from the LGS-SARA or the calculations that the.
1.3 factor is appropriate since the stresses in the containment struc-ture may not be sufficiently high to warrant the assumed 10 percent damping value. The median capacities of the three RPV components ran'ge between 0.67 and 1.379, while the limiting median capacities of the supporting containment structure components are as follows:
Sacrificial shield wall 1.6g Containment wall (shear failure) 3.4g RPY pedestal (flexural failure) 2.8g i
The upper portion of the RPY is resisted by a ring at the top of the shield wall which, in ' turn, is anchored to the containment wall by steel lateral braces. The relative stiffness of the lateral supports versus the stiffness of the sacrificial shield wall is not known. If a 2-32
, .m._. . .-mu - - > - -
^ ^ - - ' * ^ * ' " " '
3
major portion of the resistance comes from the shield wall, then 10 percent damping is probably appropriate. On the other hand, if the input to the RPV is dominated by the support at the top of the shield wall,10 percent damping may be tco large.
If the 1.3 damping response factor is changed to unity, which is the most conservative assumptifon for this factor, it is estimated that the mean frequency of core melt would increase by a factor of approxi-mately 1.10, which is a small effect.
In the original analysis conducted for the design of the contain-ment and RPV components, a coupled model was used with a single input time history. An additional uncertainty for variation in response due to time history analysis snould be included for the RPY-related im-ponent capacities. Also, the model used to develop the capacity of the RPV lateral support is approximate and, hence, additional uncertainty is present.
l It is believed that due to the SRSS operation for combining uncertai'ities, the effect of these additional uncertainties would have a small effect on the mean frequency of core melt.
2.1.3.5 Potential Impact Between Reactor Building and Containment The reactor building and containment are constructed on different foundations and are separated by a gap filled with crushable material.
The gap reportedly varies between one inch at the foundation level to
, three inches at the top of the structures. It is stated in Appendix B of the LGS-SARA that at 0.1g, the containment begins to uplift, and at 0.459 the two structures begin to impact at elevation 289 feet (it is believed that elevation 283 feet is the correct level). It is also stated that since the' reactor building shear walls are expected to fail j between 0.74g and 1.0g no signficant additional damage due to impact is expected to occur.
This assumption was questioned during the review. Three possible effects were considered. First, the impact between the structures might 2-33 i
- w. w+,=e-l- --a * * " * " * * - ~ *
- 9.m***9**-* " '
l t
cause higri frequency motions'which could affect electrical and control equipment. Based on inspection of the plant, the gap between the reactor butiding and the containment appears to be irregular; thus, the transfer of energy during impact would occur over some finite period of time which would soften the impact. The suddenness of impact would also be cushioned by local crushing of the concrete. Because of the large size of the walls and floor slabs, gross structural failure due to impact is not expected. As a minimum, the chatter and trip of relays would increase; however, NUS states that this is not a problem whether caused by either impact or just due to dynamic motions.
It is not clear whether failure of the electrical equipment located in the reactor building will be increased by impact between the two strv-tures. The capacity of the electrical components located in the reactor building (some of which are located at elevation 283 feet within 30 feet of the seismic joint) range between 1.46g and 1.56g. This is considerably higher than the motion level at which impact may occur; hence, these capacities may, in reality, be less.
The second potential problem is spalling of concrete which could fall and impact safety-related equipment. It was learned during the tour of the plant that all electrical and control equipment are located away from the seismic joint. Thus, these types of components will not
- be affected. Various safety-related pipe lines cross between the two
! buildings. It is expected that the size of any spalled concrete p'1ces will be small since the reinforcing steel will tend to hold any frac-tured " concrete pieces in place. In addition, the slope of the contain-ment wall will break the fall of spalled concrete pieces. The risk of a major rupture of a pipe or valve due to impact from spalled concrete is believed to be relatively small.
The final concern is the relative displacements caused by the movement of the two buildings and their effects on safety-related piping. It was stated at the meeting with SMA that all piping which l
l 2-34 J
,.. -. .------n-- - ~ ~a.--~. .. ~ :-n - - - . . . . -
l .
l l
contains hot water has sufficient flexibility to accommodate temperature changes to resist the potential relative displacements between the two structures. due to earthquakes. Subsequent to the meeting at SMA, the question arose concerning whether piping with lower temperature require-
! ments could resist the potential relative displacements. During the tour of the Limerick plant, an 18-inch diameter line was identified and inspected. The line number was obtained (GBB119) and the locations of l
lateral supports were found on the isometric plans in the plant engin-eering office. It was confirmed that this ifne belongs to the RHR system and is a low temperature line. The first critical support was located approximately 10 feet horizontally and 12 feet vertically from
! the containment wall in the reactor building. The flexibility of this pipe was checked approximately and it appears to have sufficient flext-bility to resist two-to-three inches of relative movement. A stress of approximately 10,000 psi would be caused by a three-inch relative dis-placement which, when adced to other stresses, probably would not signi-ficantly affect the core melt frequency distribution.
Several small ifnes (probably control-related) were attached to a valve close tc the containment wall. These lines were also attached to the reactor building close to the valve. It is possible that these lines might fail during large relative motions; however, it was stated by NUS that small leakage in small lines is acceptable.
The concerns raised regarding impact between the containment and reactor building have not been entirely resolved. The effect of impact on the capacity of electrical and control equipment should be addressed by PECo. In addition, all the safety-related piping which connects both buildings should be systematically reviewed to verify that sufficient' flexibility is provided to accommodate relative displacement between the two structures.
i e
- 2-35 wa l
l l
1
(
u ..
I 2.1.3.6 Electrical and Control Equipment The mean frequency of core melt reported in the LGS-SARA is 5.7x10-6 per year. About 60 percent of this value is contributed by sequencess T E uX, which includes the following five electrical or control components which are in series:
e 440-Y bus /SG breakers e 440-V bus transformer breaker e 125/250-V de bus e 4-KV bus /SG e Diesel-generator circuit breakers These components have median effective peak acceleration capacities which reportedly range from 1.46g to 1.56g (see LGS-SARA Table 3-1), and which contribute most of the mean frequency of core melt value of 3.15x10-6 reported in th e LGS-SARA for sequence T s3 E UX. A concern raised in the review is the actual number of units which exist for each one of these five components. For an increase of one additional inde-pendent unit (e.g., if there are two independent switchgear breakers instead of only one), the mean frequency of core melt will increase by approximately 0,*x10-6 per year.
Several issues should be considered in determining whether addi-tional units should be added in series. First, the fragility values for these components are based primarily on generic data obtained from equipment tests for the Susquehanna nuclear power plant. It is not apparent from the documentation in Appendix B nor the LGS-SARA whether the test specimens used in the Susquehanna tests were for single or multiple units (i.e was one switch gear breaker tested at a time, or were multiple units tested simultaneously?). Also, how similar are the components in the two plants?
2-36
. 'T l s-1
_ _ , , . , , . . . . ,.- . - - - ~ ~ * '%,$ {f *NW' ~ ~ * * * * * ~ ~ * '~~ - " ~
The second consideration is the question of independence between components. It can be argued that identical units have high cacacity dependence (i.e., if two units of the same component are subje ,ted to the same dynamic' motion either they both will survive or they both will fail). If two components are located next to each other and receive the
, same dynamic input, they al:o may have high resconse dependence. This j is true even though they may be different types of components.
If multiple units of a particular component exist in series (e.g.,
440-V bus /SG breakers) but they are identical units located next to each other, they may be in a practical sense perfectly dependent, and the frequency of failure would be equal to tha frequency of failure of one unit. On the other hand, if the units are constructed differently and/or placed at different locations, they may approach being indepen-dent which in the extreme case implies that the frequency of failure is l approximately equal to the stan of the . individual failure frequencies.
In order to evaluate the impact of this concern pECo should deter-mine the nunber, location, .and characteristics of the electrical and control equipment which are part of sequence T E 33UX, and compare the components to the generic test specimens from the Susquehanna tests. As suggested in Section 2.1.3.7, component-specific calculations should be performed to develop the fragility values for these components since they are significant contributors to the frequency of core melt.
2.1.3.7 Review of Significant components A copy of the calculations perfomed by SMA for the signficant components listed in Table 3-1 of the LGS-SARA were obtained and reviewed. Although the capacities of other components were considered in the review, the effort focused on the significant components which affect the dominant sequences leading to core melt. As an aid in this phase of the review, equipment fragility values developed in the Seismic Safety Margins Research Program (SSMRP) were used as a guide. (22, 23) t l
2-37
'9 N
- . , _ . . _ _ .- ,_ , m_ _ . - _ , . . , . . . , . , , , , , , , , ,y-, . ~ , , _g.,-
The following comments are given for the 17 significant components.
Offsite Power (500/230-KV Switchyard)I (S ) - The fragility for offsite power is based on the failure of porcelain ceramic insulators.
No specific calculations were given for this component. The capacity is based on historic data and is reasonable.
, Condensate Storage Tank (S 2 ) - This component is not a major con-tributor to the mean freq'uency of core melt. The capacity of the tank is based on the weakest failure mode which is shell buckling. A small ductility value of 1.3 was assumed. This is probably reasonable but may not be conservative since a buckle could caust a leak in the tank. This assumption is also inconsistent with the analysis performed for the SLC tank where buckling also controlled. For this case, no ductility was
, assumed.
No adjustment for soil structure interaction was made which assumes that the tark is on rock. It was not apparent from the tour of the Limer1ck site that the tank base is founded on rock; however, based on the fundamental frequency of the tank given in the calculations, the ,
effect of fill would increase the capacity. In sunwary, the frigility parameters for the condensate storage tank appear to be reaso'nable.
Reactor ' Internals (53 ) - The capacity of this component is limited by the strength of the shroud support. The exact failure location was notgivenInthecalculations. The capacity factor was derived based on i the calculated stresses obtained from the original design analysis. As
! discussed in Section 2.1.3.4, only one time history was used in the analysis. Although a randomness logarithmic standard deviation of 0.05 was used, this value is low for the amount of variability which could occur, if multiple time history analyses had been used. The total l
effect of increasing the logarithmic standard deviation for time history _
! variability is small. -
x
+
2-38 ~'
s_
. ..} -- . - _- ; - .
3
~
. i f
f
- ,r j '
/,
As discussed in $ectiof 2.1.3.4; the factor of 1.3 which increases
' ~ ~ ~
the capacity of the reac' tor faternals to reflect 10 percent damping expected for the containment (as opposed to 5 percent damping in the original design analysis) may be high. It is estimated that the maximum impact, if this factor wer,e 1.0, would be an increise in the mean frequency of core melt by a factor of approximately 1.10.
Reactor E'nclosure and Control Structure (S 4 ) - The capacity of this
~. >
, component is controlled by the failure of the icwest story shear walls and is based on adjusting the forces obtained frem the original design analysis to median-centered values. As discussed in Section 2.1.3.3, f .
s ,,
! "$ie v Mdian, capacity is better represented by 0.909 (as compared to 1.05g giver in the LGS-SARA). This change wou'Id increase the mean frequency
.chcire melt by approximately 20 percent.
,,>,.V >
Ifmiaf noted that the uncert6aty value ;for modeling was only f 4
~.0.L.} tecau,se of the approximate natuh of f jbb analysis which was con-oucted,_a val'ue of at least 0.20 is moie anbedgefate. Ir. comparison, a a se < -
., modeljag un.;ertainty value of 0.17 7as uses for testing in developing the,frigility for equipment, which gives an in$1 cation of a value for '
this factor that is nore reasonablet As discussed in Section 2.1.3.1, a ductility value of 2.5 assumed
_. for the case of shear wall flexural failu're is low. However, the effect of this value is balanced by the extra factor assumed for earthquake
^
s'ize' effects used to adjust ~'the hazard curves from sustained-based peak p acedi'erationi,0 i an effect3v'e peak acceleration parameter.
-,- s -
N, CR0 Guid4%na (Sg) - The capacity offs CR0 guide tube is con- ,
-trc11M by,.fLncti,onII binding-of the control rod due to bending. The a fr'ajflief,partneters , , . ,n art based on test results' coupled with the response i
of the guide tube calculated during the plant. design. The test capacity was increased about 20 perc'ent based on judgme$t since failure was not observed in the tests. - -
- Qf -
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~, j
, 7' ,
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e
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m_ . . _m _ _ _ _ _ _ _
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, x
, j ' '
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Since the CR0 guide tubes are. attached to the reactor pressure vessel (RPV) the consents above for the reactor-internals, pertaining to '
use of a one-time analysis history and containment' damping, also apply to the CRD guide tube analysis. '
j . Reactor Pressure Vessel (5 6 ) - The capacity of the RPV is due to 4
tha potential
- 4 ,
failure in the weld between the connections of the top 5
- ~ ' 'wm, ,r
- orts for the RPY and the top of the shield wall . An approximate
.n snelysiscwas used to detemine the median capacity factor, wherein the total capacity was at,sumed to be equal to the sum of the capacities frem the support skirt and failure in the weld at the top support. A 0.1 uncertiinty value' was included for modeling, which, in the opinion of the reviewers, is s5all. Similar: t( the comments made for the reactor enclosure and control structure above, a value cf at least 0.20 is appropriate for- thiftipeLof apprcrimate analysis. The effect of this i
size of increase in variability would have a small effect on the mean s frequency of core melt. -
l The ccwasents given for the reactor,intarvals, pertair.ing to one-time history.,_and containment damping, aiso apply to the RPV capacity.
, i s Hyefraulic Control Unit (S7 ) - The ccioponents of the hydraulic co'a'thl unit consist of valves, tanks,i piping, 'and electrical I control s. The fragility parameters are cased on tests and fragility I
cliculations- perfomed for the SusquehaNna nuclear power plant. In essence, the median capacity from Susquehanna was scaled by the ratio of
,e. the twos SSE peak ground acceleration values (i.e., 0.10/0.15). It is
-t
. Q, not appsrent from the documentation fn either the LGS-SARA nor the sup- ~
! '. porting halculations for this component whether the SSE scaling from s
N Susquehanna is appropriate. The. cancerns include possible' differences in the foundation condition andr hence, the response-of the reactor enclosure, loqations of the hydraulic conth1 units in the two plants q .s (i.e.Nis one unit ^ higher, therefore it has-a higher response?) and, l T i finally, constdoction and, hence,' similarity of the two units. These d issues spould be addressed by PECo. .
1 3'
3_ - s i; y 2-40 W ,
^
'k 3 , ,
, ?- ,
3
.,..,,2 r---, : -- -- - -
I*., ; _1_ _ . .t- C.s -, - - - - - - - =-- " "--' ~ '
~
-~
The uncertainty for the spectral shape factor for this component appears to be conservative. The logarithmic standard deviation values are based on the range of ratios between the test response spectrum 3'
(TRS) and the required response spectrum (RRS) at different frequen-cies. The total range of values for different frequencies and for the two horizontal directions were used to calculate the uncertainty val ue. If the components have similar dynamic characteristics and capacities in the two horizontal directions, the range should be based on the minimum of the largest ratio in the two horizontal directions and the maximum of the largest ratio. If tnis approach is used, the uncer-tainty value is approximately one-third (i.e., 0.09 canpared to 0.29).
Even if the revised value is doubled for modeling uncertainty, the value used in the LGS-SARA will still be conservative.
The median capacity value also appears to be conservative, but was developed using considerable judgment. The minimum ratio .of the TRS and RRS values at the frequencies considered in the analysis was used. This value was assumed to represent the 95 percent level of survival (i.e., 5 percent would fail above this level) along with a 0.40 logarithmic -
standard deviation value. These two assumptions lead to doubling the minimum ratio to produce the median value. The final median value is essentially equal to the average of all ratios of t."e TRS to RRS values.
It should be noted that the total uncertainty logarithmic standard deviation value for the hydraulic control unit is 0.52 which is tne
, highest value for any of the significant components. Although the uncertainty value for the spectral shape factor may be high, the total uncertainty appears to be reasonable considering other uncertainties due to modeling which have not been included.
l SLC Test Tank (Sg) - The capacity for the SLC test tank is based on generic calculations for rigid equipment. This tank is supported on four columns and is not rigid. Based on inspection of this component during the plant tour, it appears to be very strong; however, the analy-sis performed for this tank is not applicable to the actual component. -
l 2-41 l
l v
, i l
. l j __ _ . _ _ . - - .
m - _ _m_.- ..-- 2
The capacity of_ the anchor bolts which attach the base of the four
! columns to the concrete floor, should be analyzed. The response factor should be recalculated taking into account the flexbility of the tank and the actual charactertistic of the four columns. If the tension force in the columns or anchor bolts control the capacity, the earth-quake component factor may be as low as 0.71 (as compared to 1.04 which 1
was assumed in the generic component ' analysis). Since t1e capacity may be controlled by a ductile element, a ductility value greater than 1.0 may be appropriate. In strmary, a component-specific analysis should be conducted for the SLC test tank.
Nitrogen Accumulator (SLC) (Sg) - The nitrogen accumulator is described in the calculations as an 18-inch diameter by 48-inch high tank which is anchored to the floor with six bolts. After visiting the Limerick plant, the reviewers are uncertain if the nitrogen accumulator which they saw fits this -description. Since the capacity of this com-ponent is based on extrapolating an analysis from Susquehanna to the Limerick site, the similarity between the nitrogen accumulators at the two plants should be verified.
SLC Tank (S10) - The capacity for this tank is based on the buck-ling of the shell, which was the weakest mode of the various modes of failure which were checked. One other possible failure mode is tearing of the base plate flange through which the anchor bolts penetrate. This j failure mode apparently was not checked. There are no stiffening ele-ments in the vicinity of the anchor bolts, which may mean that tearing of the base plate flange is the weakest capacity. This possibility should be checked.
The uncertainty value for modeling error was asstaned to be 0.10 which is small. A value equal to 0.20 would be more appropriate; how-ever, this change would have a small effect on the frequency of core mel t.
2-42 3------- -~
+
l 440-V Bus /SG Breakers (Sti) - The capacity of this component was developed in a similar manner to the capacity for the hydraulic control !
unit, which also was based on test data from the Susquehanna nuclear power plant. The calculations, which were based on the ratios of the TRS to the RRS at different frequency values, are not clearly stated.
The minimian ratio was assumed to represent the 95 percent level of survival along with a 0.40 logarithmic standard deviation value. These two assumptions led to doubling the minimum ratio. The final value is close to the average ratio (however, calculations of the average ratio are not apparent). It is interesting to note that the uncertainty value for the spectral shape factor is only 0.08 which is much less than the value of 0.29 obtained for the hydraulic control unit (see comments above for the hydraulic control unit).
In sumary, the fragility parameter values for this ccmponent appear reasonable, but it was not possible to check all the calcula-tions. Since this component is a significant contributor to the mean frequency of core melt, a specific analysis should be conducted for this component.
440-V Bus Transformer breaker (S 12I '
125/250-V DC Bus (St3),
4-KY Bus /SG (S14) -The capacities for these three components are the same and are based on the fragility analysis of the diesel generator circuit breakers. The only difference between the capacities of these three components and the diesel generator circuit breaker capacity is
, that the former components are in the reactor enclosure, while the later component is in the diesel generator building. Couuments concerning these three components are the same as given below for the diesel gener-ator circuit breakers.
t Because these three components contribute signficiantly to the mean frequency of core melt, a specific component analysis should be con-ducted for each.
l 2-43 v
.-p%-< ,g 4 y= mar -- ' * * ' " ~
~
l l
Diesel Generator Circuit Breakers (S15) - The capacity of the
! diesel generator circuit breakers is based on an ana?ysis of test data for the Susquehanna plant. The approach used to develop the capacity factor is identical to the approach used for the hydraulic control unit (see consents above). The same issues for that component also apply to the diesel generator circuit breakers (and also the three components above, i.e., 512' S13, and St4).
Since this component is a significant centributor to the mean frequency of care melt, a specific analysis should be conducted for this component.
Diesel Generator Heat and Vent (S16) - The capacity of the diesel generator heat and vent is supposedly based on the fragility of the exhaust fan supports which are assumed to be the critical link. How-l ever, the actual fragility parameters are based on generic passive c
-flexible equipment. The calculations for this class of equipment were specifically formulated for tanks and heat exchangers. It is stated in the calculations that shock test data indicate the capacity is 9.5g for -
the handling units; thus, the values used are conservative. However, since this component is a significant contributor to the mean frequency of core melt, a specific analysis should be conducted.
RHR Heat Exchangers (S17) - The capacity of the RHR heat exchanger was obtained by scaling Me capacity factor for the same component at .
the Susquehanna nuclear power plant. It is assumed in the calculations that the response factors for Susquehanna and Limerick are the same.
The controlling element is the lower support bolts.
The earthquake combination factor is 0.93, appears to be high since the columns supporting the RHR heat exchanger are located at the four corners of a square pattern. Since tension in the bolts is significant, the factor will be somewhere between 0.71 and 0.93.
2-44 V-m -e w - m , rr- z
~
. .. .n . . - - - - - - - - - - - - - - - - - - - -
This component does not appear to be a significant contributor to the mean frequency of core melt; hence, small changes in the values of the capacity factors for the RHR heat exchanger do not appear to be critical .
~
2.1.3.8 General Fragility-Related Coments i
The following comments are made in order to inform the reader of potential issues which because of their philosophical nature may not be resolved in the near future. Also, minor issues and errors which were found during the review are documented for completeness. The reader is directed to Reference 3 which gives a more' detailed discussion of some of these general issues.
As discussed in the previous sections, there are cases where the uncertainty values seem to be. low. In particular, modeling errors appear many times to be smaller than what was expected. In Section 5.3.1.4 of the LGS-SARA, it is stated that the coefficient of variation for equipment response factors is about 0.15. Since this factor is ve'ry sensitive to the relationship between the equipment fundamental frequency and the frequency corresponding to the peak of the floor response spectrum, it is easy to visualize cases where a slight shift in
, frequency could mean a factor of 2 or 3 (or even more) in the value of the spectral ordinate. Thus the 1.ogarithmic standard deviation for response should be developed on a case-by-case basis.
In general, the uncertainty in some of the parameters has been understated. In particular, there is uncertainty in' using a simplistic analysis to obtain the capacity of a component which was not recognized in the LGS-SARA. On the other hand, the median capacity values are probably on the high side. These two effects likely are self-compensating.
4 No uncertainty was assigned to the ground response spectrum factor used in the analysis. By definition this implies that this is the 2-45
{
'~
_ . _ _ _ _ ~ E 7.. Z ~ C I ' - - I '~~ '~ - ' ~~ ~ ~ ~~~ '
~
I absolute best (within the context of the analytical model) that can be
! achieved; hence, there is no motivation ever to conduct site-specific studies to improve the estimate of the frequency content of the seismic input. Although Limerick is a rock site, there is still uncertainty in the ground response spectrum which should be included in the analysis.
It is believed that a reasonable value for uncertainty, if included, would have a small effect on the frequency of core melt.
The documentation of the basis for the fragility values does not carefully distinguish between the categories of infomation which were used. The use of subjective or data-based infomation (either analysis or testing) should be specifically noted to infom the reader. In addition, sensitivity analyses should be performed to indicate the robustness of the assumptions. This is particularly applicable to Chapter 3 where the fragility, hazard, and systems information is com-bined to produce the core melt frequency distribution.
The issue of dependency and its affect on the core melt frequency distribution was considered in the review of the LGS-SARA. Except for sequence 3s T E uX, it appears that any additional capacity or response '.
i related dependency effects would not have a significant impact on the 4
mean frequency of core melt. For the case of Tss E uX, Section 2.1.3.6 discusses the implications if additional components were added to the l
series expression. For the current Boolean expression for the T E3sVX i
sequence, if any additional dependency exists, the frequency of core melt would decrease. As discussed in Reference 3, there are potential dependency effects which could effect the fragility values for cable
' trays and piping systems, although it is likely tnat the current capa-city values account for these effects.(3)
Another important issue is the use of ductility factors for one degree of freedom (SDOF1 models to represent multidegree of freedom (M00F) structures or equipment.(3) Research is required to resolve this issue. At the present, not enough uncertainty is generally assigned for this situation.
, 2-46 s
i 'e
i As discussed in Section 2.1.1, design and construction errors are l1 not systematically recognized and quantified in the LGS-SARA. This is a particularly important consideration for components in series which could lead to a' major failure if only one of the components fails. At i
best, the results of a seismic PRA can only be used to make relative
! comparisons.
l One concern which was raised is potential leakage through internal components caused by seismic motion, thus bypassing a closed value barrier. This probably is not a major problem but should be formally verified by PECo. The MSIV and purge and vent valves are important i
examples. Also, the type of SRV used at Limerick has a history of j sticking randomly in the opened position (i.e., failing to close after I
the signal is received). The possibility that seismic motions could increase the likelihood of this type of failure should be addressed.
l l The effects of soil pressures on the buried walls of the reactor i enclosure were not explicitly addressed in the LGS-SARA. Because the walls have as much as 40 feet of fill against them, they should be
- investigated to determine if the fill reduces the capacity of the reactor enclosure walls.
i The potential for secondary components failing, falling, and impac-
{ ting primary safety-related components apparently has not been syste-matica11y addressed since the plant is still under construction. The potential effects of block walIs failing has been considered. Other components could also be a potential hazard. At the completion of l construction, secondary components should be reviewed and their capac-l ities incorporated into the LGS-SARA f f they are weaker than the primary l components already considered.
[ On page 5-15 of Appendix B of the LGS-SARA, the value 648 K in.
should be 648,000 K-in. This is believed to be a typographical error.
2-47 s
- - - . - . . - - '- m +==.e . . . - %, _. . _ _ _ . .
y,. n m y.. __ ,.y ,.% ,
.,.s,-.- ., g y ,9, .
,,.._y 3 my.,,....,, , p ..w 7
On page 5-60, the damping factor for valves appears to have been included twice (once for the piping and once for the valves). It was explained by SMA that only one fac' tor was used for both piping and for valves and is based on adjusting the damping used in the original design analysis (i.e, 0.5 percent) to a median-centered value (i.e., 5 percent).I24I i
Toward the completion of the preliminary review, Section 10.1.6.5 was brought to the attention of JBA (other parts of Chapter 10 were not reviewed by JBA). In this section, the effect of earthquakes on the t
effectiveness of evacuation was quantified for the various accident classes. 'The argument for limting upper-bound accelerations on the hazard curves given in Reference 18 was incorrectly used to establish that below 0.61g effective peak acceleration evacuation will not be impeded. This value was then used to develop the percent of occurrence when evacuation would be affected by earthquake. Although the arguments
~
in Reference 18 are appropriate for establishing upper-bound accelera-tion limits for the hazard curves, the rationale has been incorrectly reversed. The result of this error means that the percentages of affected evacuations are much higher than given in Table 10-7. pECo should reexamine the percentages and establish more realistic values and incorporate them in the offsite consequence analysis.
2.1.3.9 Closure l
The LGS-SARA differs from the IPPSS and ZPSS in t.at the mean frequency of core melt is dominated primarily by five elec*rical compo-nents in series, which have nearly the same median capacitie;.. In
- contrast, nonelectrical components and structures controlled the results
- of the IPPSS and ZPSS. (Note that in Section 3.1.2, discrepancies concerning the Boolean equations are discussed which, if correct, may-mean that the electrical components do not dominate the results.)
2-48 l .
.)
i
The capacities for the LGS-SARA electrical components are based on generic tests and are not component specific. This approach is reason-able as long as the components do not control the final results. Since the electrical components are significant contributors, a more detailed analysis should be conducted. The recommendations given in Section 4.1.3 are directed to this goal.
, 2.1.4 References to Section 2.1
- 1. Pickard, Lowe, and Garrick, " Indian Point Probabilistic Safety Study," Prepared for Consolidated Edison Company of New York, Inc., and Power Authority of the State of New York, Copyright 1982.
- 2. Pickard, Lowe, and Garrick, " Zion Probabilistic Safety Study,"
Prepared for Consolidated Edison, Co., not dated.
- 3. Kolb, G. J., et al., " Review and Evaluation of the Indian Point Probabilistic Safety Study," Prepared for U. S. Nuclear Regulatory Comission, NUREG/CR-2934, December,1982.
- 4. American Nuclear Society and the Institute of Electrical and Electronics Engineers, "PRA Procedures Guide," Vol.1 and 2, U.S.
Nuclear Regulatory. Commission, NUREG/CR-2300,1983.
- 5. Cornell, C. A., "Probabilistic Seismic Hazard Analysis: A 1980 Assessment," Proceedings of the Joint U.S.-Yugoslavia Conference on Earthquake Engineering, Skopje, Yugoslavia,1980.
- 6. Philadelphia Electric Company, Limerick Generating Station Final l
Safety Analysis Report, 1983.
2-49
~
f ,
i
.. . . . _ _ _ . . -, -- - - - - - =~ - -
L
- 7. Cook, F. A., D. Albaugh, L. Brown, S. Kaufman, J. Oliver, and R.
Hatcher, " Thin-Skinned Tectonics in the Crystalline Southern
, Appalachians: C0 CORP Seismic Reflection Profiling of the Blue Ridge and Piecmont," Geology, Vol . 7, pp. 563-567,1979. a
- 8. Seeber, L., and J.G. Armbruster, "The 1886 Charleston, South Carolina Earthquake and the Appalachian Detachment," Journ. ,
Geochys. Res., Vol . 86, No. B9, pp. 7874-7894,1981.
- 9. Tera Corporation, " Seismic Hazard Analysis," Prepared for U.S.
Nuclear Regulatory Commission, NUREG/CR-1582, Vols. 2-5,1980.
- 10. Diment, W. G., T. C. Urban, and F. A. Revetta, "Some Geophysical Ancmalies in the Eastern United States," in The Nature of the Solid Earth, Ed. E. C. Robertson, pp. 544-572, 1972.
- 11. Sbar, M. L., and L. R. Sykes, " Contemporary Compressive Stress and Seismicity in Eastern North America: An Example of Intraplate Tectonics," Geol . Soc. Am. Bull ., Vol . 84, pp.1861-1882,1973.
- 12. Fletcher, J. B., M. L. Sbar, and L. R. Sykes, " Seismic Trends and Travel-Time Residuals in Eastern North America and Their Tectonic Implications," Geol . Soc. Am. Bull ., Vol . 89, . pp.1656-1976,1978.
- 13. Yang, J. P. and Y. P. Aggarwal, "Seismotectonics of the North-eastern United States and Adjacent Canada," Journ. Geophys. Res., .
Vol . 86, No. 86, pp. 4981-4988,1981.
- 14. Street, R. L., A. Lacroix, "An Empirical Study of New England Seismicity: 1727-1977," Bull. Seis. Soc. Am. Vol. 69, pp.159-175, 1979.
2-50
, , , , , , a . . . . . . ~ . ame- = = * = = = = =- =*N"***"** -"** * ' ~~ ~
--, - -a - - -, -_- - - - - - - - a. ----_ - - - - - - - - - _ - - - _ . - - - - - - - - - - - - - - _ _ - - --,.-----_.--
I
- 15. Nuttli, O. W., "The Relation of Sustained Maximum Ground Accelera-tion and felocity to Earthquake Intensity and Magnitude," Report l 16, Misc. Paper S-7-1, U.S. Army Engineer Waterways Experiment Station, Vicksburg, Miss., 1979.
- 16. Cornell, C. A. and H. Merz, " Seismic Risk Analysis of Boston,"
Journal of the Structural Division, American Society of Civil Engineers, ST10, 1975.
- 17. Trifunac, M. D. and A. G. Brady, "On the Correlation of Seismic Intensity Scales with the Peaks' of Recorded Strong Ground Motion,"
Bull . Seis. Soc. Am. , Vol . 65, pp.139-162,1975.
- 18. Xennedy, R. P., " Comments on Effective Ground Acceleration Estimates," SMA Report 12901.04R, February, 1981.
- 19. R. P. Kennedy, et al., " Engineering Characterization of Ground r
Motion Effects of Charateristics of Free-Field Motion on Struc
- tural Response," SMA 12702.01, prepared for Woodward-Clyde Consultants,1983.
- 20. Riddell, R., and N. M. Newmark, " Statistical Analysis of the Response of Nonlinear Systems Subjected to Earthquakes," Depart- ,
ment of Civil Engineering, Report UILU 79-2016, Urbana, Illinois, August, 1979.-
- 21. McGuire, R. K., " Transmittal to Mr. Howard Hansell of Philadelphia Electric Company," July 12, 1983.
- 22. Kennedy, R. P., et al., " Subsystem Fragility," Lawrence Livermore National Laboratory Report prepared for the U. S. Nuclear Regula-tory Commission, NUREG/CR-2405, October, 1981.
2-51
,,.,,e d -
meew'*a h e*
- - - * * *-
- epumme e.-
--e.w_ . , , . _ _
m,
_s
_ a_%__ww-.a-_ .._ _%,_.e-+ _._ ,_
- 23. Bohn, M. P., " Interim Recommendations for a Simplified Seismic Probabilistic R_isk Assessment Based on the Results of the Seismic Safety Margins Research Program," Lawrence Livermore National Laboratory, Draft, April 15, 1983.
I
. 24. Letter from R. D. Campbell to J. W. Reed, dated July 11, 1983.
4 4
e l
l
\
esp 52 -
,[ ,, , - - " s-. iW A-.e.s-^ 4 = * **-N * '
^
~
Table 2.1 Comparison of Mean Frequency of Core Melt Values Contribution to Mean Frecuency of Core Melt Secuence Aporoximate Analysis LGS-SARA Values T 33 E UX 4.0-6* 3.1-6 TRsB 9.5-7 9.6-7 T 3RPV 4.4-7 8.0-7 TECC 33m2 6.0-7 5.4-7 TRC 3gm 3.5-7 1.4-7 TEW s3 1.1-7 1.1-7 Total 6.5-6 5.7-6 4.0-6 = 4.0x10-6 e
B 2-53 9
. .. . . . . . . _ . _ . _.. . . . , - - - , --. - ~ - - - - - - - - -
--7-~<7 _ . __ _ ,, _
Table 2.2 ,
Hazard Curve Contribution to Mean Frequency of Core Melt Contribution to Hazard' Curve Mean Frecuency of Core Melt Percentage Decollement 3.0-6 46.2 Piedmont, % x = 6.3 2.6-6 40.1 Piedmont, thax = 5.8 5.7-7 8.8 Northeast Tectonic 2.4-7 2.7 Crustal Block, Mmax = 6.0 6.2-8 1.0 Crustal Block, Mmax = 5.5 1.5-8 0.2 .
Total 6.5-6 100.0 2
I .
2-54 -
e*" * *re-**' - o gaan _
. t3 --
'"
- gg g, 9N * * ~
l
- , Table 2.3 Hypothetical Mean Frequency of Core Melt (Based on Individual Hazard Curves)
Individual Mean Frequency Ratio to Hazard Curve of Core Mel t 6.5-6 val ue Decollement - 3.0-5 4.6 Piedmont, kax = 6.3 1.7-5 2.6 Piedmont, Mmax = 5.8 3.8-6 0.58 Northeast Tectonic 8.0-7 0.12 m Crustal Block, M,,x = 6.0 4.1-7 0.06 Crustal Block, M,,x = 5.5 1.0-7 0.02 aum f
r 2-55
~~
e l
i
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e s8"4 cc aC m
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. * %c 9m t a ,- ',
i g
c
- Do + m .,e. e n 2 a ..e
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=
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("
- i
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o 9
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o - g i -
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o 6
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[ d 2-56
1 l
l 10-2 Histcric Seisnicity Curies
- 1. Imin
- IV. Ima = X, b = 0.45
- 2. Imin
- IV. Ima = X, b = 0.72
- 3. Imin
- IV Ira = VI, b = 0.72 10-3_
LGS-S;RA Seisnicity Gr/es
\
4 Decollemnt
- 5. Crustal Block, M = 5.5
$ \ N 4 \ N 5 10 \
W \
N 1
[e \ N g N.
x %
M \ N e
h N 4 5 10-5 -- \
? \ 2 8
e \
\
\
10-6 \S 3
10-7 , , ,
0 0.20 0.40 0.60 0.80 Peak Ground Acceleration (g)
Figure 2.2 Comparison of various historical seismicity curves and the LGS-SARA seismicity curves from Appendix A for sustained-based peak acceleration for the Decollement l and Crustal Black, M=5.5 seismogenic zones.
l l
! .2-57 l l
i l
l 1
2-58 2.2 FIRE 2.2.1 Deteministic Fire Growth Modelira, 2.2.1.1 Introduction A deterministic fire growth model is used in the Limerick SARA to provide fire growth times. These times then serve as input to the probabilistic model from which the likelihood of a particular fire growth stage'is detemined given an initial size fire. The deteministic model contains the methodology which explicitly incorporates the physics of enclosure fire development.
The Limerick SARA uses the computer code COMPBRN[1,2] as its determinis-tic fire growth model. Briefly, this code is a synthesis of simplified, quasi-steady unit models resulting in what is ccmmonly called a zone approach model.
A detailed evaluation of this code and its application in the Limerick SARA appears later in this review. There are many other ccmputer codes [3-7]
which use the unit-model approach to model ccmpartment fire development. Of ;
particular interest is the DACFIR Code [8] developed at the University of Dayton Research Institute, which models the fire growth in an aircraft cabin as it progresses frem seat to seat. This is analogous to the problem of fire spreading from cable tray to cable tray as analyzed in COMPBRN.
At this point some' gener'al thoughts are deemed warranted on the compluity of fire phenomena and the state of fire science with regard to enclosure fire development. Computer models of enclosure fire development appear capable of
- predicting quantities of practical importance to fire safety, provided the ,
model is supplied with the fire-initiating item's empirical rate of fire growth and the effect of external radiation on this rate. ' As a science, how-ever, we cannot predict the initiating item's growth rate due to relatively poor understanding of basic combustion mechanisms. Questions and doubts have
- even been raised regarding the ability to predict the burning rate of a non-l spreading, hazardous scale fire in terms of basic measurable fuel properties.
However, while awaiting development of meaningful standard flammability tests and/or'more sound scientific predictions, realistic " standardized" fire test f
w y , , -
e 2-59 procedures should continue to be fonnulated for empirically measuring the rate of growth of isolated initiating items, the attendant fire plume, its develop-ment within an enclosure, and the convective and radiative heat loads to
" target" combustibles. Thus, in lieu of large-scale computer codes to assess the fire hazard in an enclosure, the unit-problem approach (as used in COMPSRN) is about the best that can be taken at the present time.
However, due to the state of infancy of fire modeling alluded to above, many judgemental assumptions in both modeling and physical data must be made in order to model fire development in the complex enclosures existing in nuclear power plants. Additional canplexity is introduced when one considers the fuel as electrical cable insulation rather than the more commonly con-sidered fuels such as wood or plastic slabs, wlich may have a more unifonn cam-position than cable insulation.
In fact, as discussed later, some of the models used in COMPBRN are non-physical . That is, while usually leading to results which are highly con-servative, these models do not adequately reflect the dependence on the physi-cal parameters which are evidenced in experimental data. Other mocels, assumptions, and omissions in the application of CCMPBRN to the Limerick, SARA are either conservative or non-conservative.
This canbination of non-physical models and conservative as well as non-conservative assumptions leads to very large uncertainties in the determittis-tic modeling process. It 'is therefore also difficult to quantify the effects of these uncertainties on the probaoilistic analysis, since the latter uses the results of the detenninistic analysis as input. Indeed, as a general com-ment, one wonders whether more is gained by making gross judgemental assump-tions, using them in an uncertain detenninistic methodology and " cranking" the results through a probabilistic analysis, than would be gained by making direct judgements on the risk of fire. In any case, we will evaluate the modeling and assumptions of the COMPBRN code and its application in the Limerick SARA in following sections. Section 2.2.1.2 briefly summarizes our concerns with the deterministic modeling, while Section 2.2.1.3 gives a more
, detailed discussion of each item. Some suggestions for reducing the uncertainties are given in Section 2.2.1.4.
1 m w. -w .m%m,=m-4ey ,
..ases_a .eg,=w - - = -m.= -e-e-- -s-.
%=a w .....-w.. ----e- -. e,. .. - .
e 2-60 2.2.1.2 Summary Evaluation of Deteministic Fire Growth Modeling 1
l The detenninistic methodology contained in the computer code COMPBRN[1,23 is used in the Limerick SARA to evaluate the themal hazards of postulated fires in terms of heat flux, temperature, and fire growth. This code employs a unit-model approach which is acceptable given the current state of the art in enclosu~re fire modeling as discussed in the previous section. However, we find some of the sub-models contained in the code to be non-physical, some as-sumptions highly overconservative, while other assumptions and applications yield non-conservative results. The uncertainties arising " rom the ccmbina-tion of these counterbalancing models and assumptions are difficult to quantify, but if forced to draw a conclusioil we feel the deteministic analy-sis as applied to the Limerick plant is generally on the conservative. side.
However, we also wish to restate that we do not feel that the counterbalancing i
of a non-physical, non-conservative model or assumption with another non-physical model or assumption, no matter how conservative, leads to a result which is useful in quantitative tems.
Based on our initial review of the deteministic fire modeling in the l
Limerick SARA, we have identified the following items of concern, which wil1 be discussed in more detail in the next section.
The burning rate model is probably the most important source of uncer-tainty in the COMPBRN code. The methodology employed is not realistic and can i lead to results which are dependent on the arbitrary choice of the size of
! " fuel elements" into which the fuel bed is discretized. Instead the f.:.1 burning rate should be dependent on the instantaneous size of the fire. Also use has not been made of existing cable flammability data.[9,10] II:is difficult to detemine if the cable insulation burning rates obtained by this method are conservative or non-conservative. For the postulated transient-combustible oil fire the burning rate considered appears overconservative with respect to ti.at reported in the literature.[ll3 Another example of non-physical modeling is the fuel element ignition time relationship. This model yields a finite fuel ignition time even if the inci-dent heat flux is considerably below the critical' value of 20 kW/m2 found necessary to initiate cable insulation damage in experiments.[12] The model q m- s og u.vsi ? = Tr--weg 9*w e -- .<--
= ,,- -
- t - ia.-. - -
9
o . :
l
. i 2-61 l l
assumes a constant input heat flux even when cables in a convective plume are considered. Convective heat flux must be a function of the difference between the plume and target temperatures, and must therefore cecrease as the target fuel heats up. A cable damageability criteria based on a critical heat flux.
i and an accumulated energy, as discussed later and in Ref.12, would be more appropriate. The model used in CCMpBRN leads to highly conservative cable ignition times.
The model used to calculate the radiative heat transfer frem the flame to a target object is also overly conservative. The radiative heat flux obtained frcm this model is much greater than that obtained frcm a classical Stefan-Boltzmann model, wherein the heat flux is a function of the flame gas tempera-ture to the fourth power. The CCMpSRN model also neglected the attenuation of the heat flux with distance due to intervening hot gas or smoke. The mcdel neglects, too, the partial reflection of the impinging radiative heat flux frem a target fuel element, as well as re-radiation, convection, and other losses. ,
Additional conservatism is introduced by assumptions made concerning the three stages of fire growth. The second stage considers fire growth to adjacent cable raceways once an initial raceway is ignited. The analysis assumes adjacent cable raceways are separated frcm the initial fire oy the minimum-separation criteria specified for redundant safety-related cable raceways (5 feet vertically and 3 feet horizontally). In other words, only one calculation of fire spread time is made for this configuration and the results are applied to all plant areas considered. This will yield a highly conservative upper bound calculation. Growth stage three assumes damage to redundant cables separated by 20 feet and up to 40 feet and those protected by fire barriers. Redundant raceways separated by more than 20 feet from the initial fire were assumed to be damaged in a time interval equivalent to the damage time of a fire barrier taken as a 1 inch thick ceramic-fiber blanket.
l This appears conservative since raceways separated by this distance would usually be damaged by convection in a stratified ceiling layer, and therefore there should be some dependence on the height of the raceway from the ceiling, l
those closer to the ceiling failing earlier than those below. Intermediate growth stages between stages two and three might be appropriate.
, _.. ,, . - s,,op e..-~=oee-mm ~
ese-*- *~ 0*~ ** *
- e p -e g e ...% -we--
-,"'.e-* * * * ' '*~
r-T yy a --, t---sy ? - + -==
a 8; 2-62 Another area of uncertainty concerns the quantity and size of the assumed transient-combustible fires. The Limerick SARA assumes three possible trans-fent combustible configurations; 2 pounds of paper 1 foot in diameter, i quart of solvent 0.5 fout in diameter, and 1 gallon of oil 1 foot in diameter. No rationale is given for this selection. It is certainly possible for larger quantities or ccmbinations of these fuels to exist in nuclear power plants. A distribution of varying quantities would be more appropriate. Also, it is not clear that given 1 gallon of oil that a 1 foot diameter pool represents the most severe hazard. A larger diameter pool will give a larger heat release al-though for a shorter duration. The damage sustained by the target cable may be a function of this combination of heat flux level and duration of imposi-tion.
Scme considerations emitted frcm the Limerick SARA would tend to make the i.nalysis non-conservative. These include the effects that enclosure walls and corners, in close proximity to the initiating fire, have on the convected heat flux and the possibility of cable damage due to convection in a stratif.ed ceiling layer.
2.2.1.3 Detailed Evaluation of Deterministic Fire Growth Modelina 2.2.1.3.1 Fuel Burning Rate The COMPBRN code [13 models the specific burning rate, 6", of the fuel, which is equivalent to the mass loss rate in combustion, for fuel surface con-trolled fires as in" = m" + Cs 6" ext .
(2.1)
The term in"o is defined as a specific burning rate constant and the second tem represents the effects of external radiation on the burning rate.
The specific burning rate constant is assumed to represent the effects of
- flame radiative heat flux to the surface, q"fl, r, and surface reradia-l tion, q" loss S"=(q"f1,p-d" loss)b 3
, where L is the heat required to generate a unit mass of vapor. Note that the use of HF , the heat of combustion of the fuel,- in Eqn. (4.4) of Ref.1 is incorrect. The correct formulation is given by Eqn. (3) of Ref.13.
l l
l
- . . , , ~~ ~, '~ ~
.--,.* e
2-63 Note that if the externally applied heat flux, q" ext, is zero, the object will burn ,at a constant rate given by in"=5"o. The consideration of
&"o as a constant for an element of fuel burning during the early growth stages of a fire is questionable. For non-charring combustibles, such as PMMA or Plexiglas, experimental data indicates that 5"o is indeed a constant.
However for complex solid fuels such as electrical cables, this may not be the case. Also, the burning rate is a function of the size of the fire through 6"fl,r and q" loss. Tha mass loss rate of a small sample of PE/PVC cable, subjected to a constant external heat flux, is shown in Figure 4.4 of Ref.10.
The mass loss rate is certainly not constant with time as would be indicated by Eqn. (2.1) with &"o and 6" ext constant by definition.
In CCMPSRN, Eqn. (2.1) is applied to each .small square " fuel element" into which the individual cable trays (super modules) have bt.en discretized. The fire is assumed to initiate in one element and spread to adjacent elements when their ignition criteria is reached due to the incident radiation frcm the initial fire. A const. int value of A"o = 0.002 kg/m 2-sec. is chosen for each element. This methodology results in a non-physical condition when the ccmplete cable tray is considered since the specific burning rate becanes. a function of the arbitrary numoer of elements into which the tray is divided.
For instance, if a fuel element was burning in infinite space with no externally applied heat flux, then according to Eqn. (2.1) its burning rate would be h" tot *S"o. However, if this fuel element is divided into two contiguous subelelements (1) and (2) with equal areas A/2 and with the fl'ame of subelement (1) supplying the external heat flux to subelement (2) and vice versa, then according to Eqn. (2.1)
&"totjm"o=[5"o+Csk" ext 3 (2.3) where we have tacitly assumed that 6" ext,1*k" ext,2*4" ext Likewise if the element we.re divided into n subelements with each j-th element supplying an external heat flux to every other element, by definition, the
. progressive. total burning rate when each of the j-subelements became involved -
will not be equivalent io the total burning rate had all the subelements been involved initially. This indicates that care must be exercised in using Eqn.
(2.1) to predict the ensuing development of a fire along an individual cable A ~_.- tray.
.._______M_. T ~TT __1 ~ ~ -~ ~ ~'~~T"~ ~~ ~~ ~
- 2 "
2-64 Intemediate scale data for the EPR/Hypalon cable used at Limerick is given in Fig. D-18 of Ref. 9. The cable weight loss for the twelve trays considered increases with time and a steady burning rate of 6.7 kg/ min was eached after about 37 minutes. This translates into a specific steady state burning rate of 0.008 kg/m2-sec. Use of such data and that of Ref.10 could remove some of the uncertainty of the present model.
For transient ccmbustibles, the fuel is not discretized and the specific burning rate is assumed to be the constant steady state value, rit"o. Table D-4 of the Limerick SARA gives values of m"o for paper and oil of about 0.061 kg/m2-sec. Hopefully, the value for paper is a misprint and should be 0.0062 kg/m2-sec. The value for oil seems somewhat conservative since Ref.
11 gives a value of 0.04 kg/m2-sec.
2.2.1.3.2 Fuel Element Ionition In the COMPBRN code, a fuel element is considered ignited simply if its surface temperature exceeds a critical ignition tempercture, T*. Addition-ally, the fuel elements are modeled as semi-infinite slabs and the losses frem the fuel to the environment due to re-raciation and convection are neglected.
An expression for the ignition time, t*, is obtained by solving the heat conduction equation, following page 75, Ref.14, for the condition of a constant imposed surface heat flux, d'o.
t* = ( r/4a)[k(T*-To)/d"o]2 (2.4)
Tnis expression is not physically corre'ct since it implies that an ig-nition time will be reached no matter how small a value of heat flux is ap-plied. Cable flammability test data [12] shows that cables are generally not damaged unless the heat flux is above a critical value of about 20kW/m2 due to heat losses at the surface.
Also, the assumption of constant imposed heat flux .is overly conserva-tive since the heat flux received by an object is a function of the object surface temperature, Ts, which increases with time as the object is exposed to the external flux.
, .. . _ , . - . - - - . ---~4.--~~--- -
- ~ K ;
'~' " ^'
t t 2-65 For instance, in the case of an oil fire 10 feet beneath a cable tray considered in the Limerick SARA, the convective heat flux at the cable surface will be q"o = h[Tp1-T] s (2.5) where Tp1 is the plume temperature at the cable height, T 3 is the cable surface temperature, and h the surface heat transfer coefficient. There fore, the surface heat flux will decrease substantially as the temperature of the cable surface approaches the plume temperature. The CCMPBRN code assuces the surface temperature remains at its initial value for the duration of the fire.
For the 1 foot diameter oil pool fire considered in the Limerick SARA, we 1 estimated the plume temperature at 10 feet above the fire using three methods.
These include two correlations of convective heat flux by Alpert,[15,16]
(one of which was used in CCMPBRN[13) and a more recent plume correlation by Stavrianidis[173 The plume temperatures thus cotained range between 370*K and 450*K. These low values of plume temperature indicate that cables witnin the convective plume and located 10 feet aoove the fire, would never reach their designated critical ignition temperature of 340*K. .This indicates the overconservativeness of Limerick SARA wnich predicts caole igntion in 4 '
minutes for this target / fire source configuration.
- Of course, one must also consider the radiative heat. transfer from tne flame to the target (the electrical cables) in order to predict the time required for the cables to achieve this critical ignition temperature. In this regard, audit calculations, using the method described in Ref.18, yields a radiative heat flux, q"r, of 0.42 kW/m2 . This is based upon use of the following equation
, q"r = (oTf14 /w) (Ap /12 ) e (2.6) where e is the Stefan-boltzmann constant; Tf1 is the flame tenperature (1255'K)[173; 1 is the distance of the target from the radiating body (with a flame height of 5 ft(16] and a cable height of 10 ft,1 is equal to 5 ft, and Ap is the flame projected surface area. The emissivity, c , was assumed h
- eg a
w -,
g g,g y, a wee t a meew a e == s<=e - - -w
2-66 to be 0.3 (the sum of a gaseous value of 0.2 and a luminous soot value of 0.1). This value of radiative heat flux, when added to the previously ~
calculated convective heat flux, then yields a value of ignition time, t*,
(via Eqn. 2.4) markedly higher than the 4 minutes stated in the Limerick SARA.
Even using the radiative heat flux model, as described in CCMPBRN, yields a similar value of radiative heat flux which is lower than that required to achieve the critical ignition temperature of 840*K within 4 minutes. In COMPBRN, the radiative flux is given by k"r=Fo.f1hr/Afl (2.7) ,
where Fo _f1 is the shape factor between the object and the flame, A f1 is the flame surface area, and r is the neat radiated by the fire wnich is expressed as
-he = yh. .
(2.8)
In the above expression, y reflects the radiant output fraction (v=0.4 as assumed in Ref.1) and 6 represents the total heat release rate of the fire.
In order to reconcile this wide disparity between ignition times reported and those calculated by the methods described above, "back" calculations using Egn. 2.4 indicated that an imposed surface heat flux, q"o, of approximately 12 kW/m2 is required to achieve a t* of roughly 4 minutes. This value is obtainable using the COMPBRN model, if Af1 in Eqn. 2.7 reqpresents the projected flame area (or pool area in this case) and not the flame surface area. This is clearly inconsistent, with the metho' dology used to derive Egn.
2.7.
These audit calculations clearly point out that the results of the Limerick SARA are based upon an overconservative estimate of critical times to reach cable ignition.
4
+
or
,p ._ %4 .. ,,. - e
- 1ewmi - " ' * ' * " ' " * ** ~ -
,~ . .
x -
2-67
.s=
N Even fr; the event that the radiative heat flux dominates the convective heat flux, the target will not absorb the total flux since significant amounts will be convected away, If a proper model for convective heat transfer, Eqn.
(2.5) is used, on e the surface temperature increases above the plume tempera-ture, heat will ' ec convected away from the target reducing the effects of radiation.
The selection of 840*K as the scontaneous ignition temperature for EPR/
Hypalon cable is also somewhat conservative since Table 3-1 of Ref. 9 pre-
- sents experimental data snowing that the critical temperature at or below which ignition cannot be achieved is 893*K for piloted ignition and is con-siderably higherifor spontaneous ignition. Actually, as stated by Siu,[13 the concapt of a thresnold ignition' temperature is somewhat i'nprecise. Ex-perimental data generally exhibit significant variations with further uncer-tainties arising if ill-defined cable insulation compositions are involved.
The crucial issue is not whether the fuel surface reaches a certain tempera-ture level, but rather if the heat gains by the pyrolyzing gases are great
~
enough to overcome the losses and trigger the combustion reactions, and if the c resulting heat of gaseous combustion is great enough to sustain the reaction.
Lee [123[has developed a set of cable damageability criteria along these
~ '
lines. For'.an applied heat flux, the time for spontaneous ignition is defined in tenns of a. critical heat flux, i;"cr, at or below which ignition cannot be initiated and an accumulated energy, E, required for sustaining ignition.
E t=E/(h" ext-dcr) _
(2.9)
Fig. 2.1 (attached) shows test data [123 for the inverse of time to 1
piloted ignition plotted'vs. external . heat flux for EPR/Hypalon cable. The
' slope of the straight line is 1/E. Also plotted is the ignition time model, Eqn. (2.4), using a critical spontaneous ignition temperature of 840*K. The
! COMPBRN model is more conservative ~ than even the piloted ignition data
- . s especially. for low levels of external heat flux, i.e., a given external heat flux willegive an earlier time to ignition than the data. Also, while the data shows no ignition below a' heat flux of about 20kW/m2, the model pre-dicts an ignition time for all values of heat flux. The 10 minute ignition o
time for stage,two,-self-ignited cable raceway fires is indicated for
^~ reference. .
4 _
4 I,e
- i
.] .
gn. ._ _ _ _ _
.__._.....n _ _ _ . _ _ _
2-68 2,2.1.3.3 Fires Near Enclosure Walls or Corners' The COMpBRN code 'does not consider the effects that the close proximity of
~
walls or corners of an enclosure can have on the temperature distribution in the convective plume of fires. The gas temperature at an elevation above the fire will be increased by the presence of walls by a magnitude that can be theoretically estimated by considering initiating fires having " equivalent"
+
heat release rates of 2 and 4 tices the actual heat release rate for walls and corners respectively. The neglect of this effect will have a non-conservative effect en fire growth calculations, especially in Fire Zone 2 where cable trays are stacked against the "J" wall.
Evidence of tne increased gas temperatures at a given elevation above a fire is available in the literature. In Ref.16, Eqns. (3) and (4) illustrate the concept of equivalent heat release rates mentioned above. Fig. 6 of the same reference shows test data of the fire positioning effects on ceiling temperature. On page 119 of Ref.19, the average plune temperature rise is found to increase by factors of 1.75 and 2.5 for fires tdjacent to walls or corners respectively. Finally, Table A-1 of Ref. 20 sht vs the upper layer gas temperature f s likewise affected by burner locations nea walls and corners.
The increased ' gas temperatures in the presence of walls results f om the effects of reduced cool air entrainment, which results in higher flam6s due to the additional distance needed for fuel vapor / air mixing. We are concerned with the distribution of energy, 'not just the maximizing of the overall energy. Even.though the code considers ccmplete combustion, wnich maximizes the heat release rate and the temperatues near the fire, the wall effect causes local temperature increases which must be considered to yield a con-servative result.
2.2.1.3.4 Stratified Ceilino layer 5
The application of the COMpBRN code in the Limerick SARA failed to con-l sider the stratified hot gas layer near the ceiling of enclosures even though such a model is included in the code. This assumption that enclosure effects are minimal may be valid since the fires considered are small with respect to i
, , , [Q * -e ,n
~ .~e,-vw ~~~5~~~*~~~~~'*'"' '~~ ~^~ "" " ~~~ '^~
2-69 i-the size of the enclosure. However, in small fire zones, as the static inver-ter room, the hot gas layer near the ceiling could preheat the non-burning fuel elements and reduce their time to ignition. Some substantiation of the
, neglect of this effect should be included in the analysis.
The consideration of thermal stratification might also eff&ct the defini-tion of fire growth stages in the Limerick SARA. It is conceivable that unprotected cables near the ceiling, but horizontally separated by more than 20 feet frem an initiating fire, could ignite prior to a cable closer than 20 feet but considerably below the ceiling. This would tend to have portions,of fire growth stage 3 thead of fire growth stage 2.
The ceiling gas layer model in CCMPBRN is based on a simplified steady gross heat balance. A unifom gas temcerature is assumed througnout the upper hot layer. Alpert[15] indicates that the ceiling gas temperature decreases with distance frca the ceiling, as well as with radial distance from tne plume axis. More recently, Newman and HillC21] have developed a transient cor-relation for the heat flux below the ceiling of an enclosure containing a pool fire, which includes the effects of forced ventilation. This correlation shows a decrease in heat flux with distance below the ceiling, but contrary, to Alpert, it indicates very little dependence on lateral separation. These works indicate that consideration in the Limerick SARA of all unprotected trays with greater than 20 feet horizontal separation as equivalent in damage rating to a fire barrier as being an oversimplification.
2.2.1.4 Reccmmendations for Imorovi tg Fire Growth Medeltnq The previous sections have detailed some of the concerns w2 have regarding the sometimes non-physical, usually over-conservative, deteministic fire growth modeling in the Limerick SARA. There are four ma,jor areas where we feel the modeling can be made more realistic, thereby reducing the resulting uncertainties. These are the cable burning rate model, the fuel element ignition time model, the flame radiant heat transfer model, and the surface temperature dependence of the convective heat transfer model.
l v
2-70 Incorporation of recent test dataC9,10] on cable flammability into the determination of the burning rate of the EPR/Hypalon cables should give a more realistic representation of fire growth. Similarly, the use of a cable ignition /damageability criteria,(123 based on a critical heat flux and an accumulated energy, wodd yield cable ignition times more consistent with test data. An improvement of the mcdel for calculating the radiated heat flux re-ceived by a fuel element, by using an appropriate f1ame ar'ea and by con-sidering attenuation due to hot gases and soot, will result in me.e realistic fire growth scenarios and establish a more correct proportionality between convective and radiative heating. Finally, the convective heat transfer model should take into account the instantaneous temperature of the surface of the object being heated. Tnis will reduce the convective heat absorbed as the object heats up and will allow for convective cooling if tne object tempera-ture exceeds the temperature of the local fire plume.
2.2.2 Probabilistic Fire Analysis Review For the Limerick Generating Station (LGS) Unit 1, the Severe Accident Risk Assessment (SARA) study reports that fire accident sequences constitute a sig-nificant portion of the overall public risk. Based on our review of the docu-ment, we have found no evidence which contradicts the conclusion that the risk of fire is significant. However, based on our understanding of the state of the art in fire PRA, and the existing inadequacies in both physical and proba-bilistic modeling in this area, we would like to avoid any judgement based on the quantitative results presented in the LGS' report. In addition, the expec-ted large uncertainties associated with the quantitative results would suggest that less importance be given to the numbers. Hence, the scope of our review is two-fold: first to identify the ex' isting inadequacies in physical and .
+
probabilistic modeling in fire PRAs in general; and second to review and com-ment on the existing LGS report for the fire risk assessment.
The generic comments associated with the physical modeling of fire growth have.been discussed in Section 2.2.1. The level of conservatism used in the deterministic analysis has also been discussed. In addition, fire growth modeling during the suppression phase will be described in the following l
sections dich basically indicate that the LGS approach is again highly conservative. Concerning the specific approach and data implemented in LGS fire risk assessment, we have concluded that:
~
7-- 7
2-71 l
- 1. The approach taken for systematic identification of critical pla'nt areas is sound, and the LGS fire hazards analysis appears to have identified all these areas.
' ~
- 2. The LGS fire analysis has adopted an appropriate data base' for es-l timating the frequency of fire in Nuclear Power Plants (NPPs).
! 3. The LGS analysis has generated plant-specific fire frequencies using the data base and has taken into account the specific features of the plant. In a few cases these estimates are unconservative.
- 4. The LGS analysis appears to have identified all important safety com-1 ponents and cabling which are located in the critical fire areas ex-cept for Zones 44 and 47. .
- 5. The event trees for panel fires generated by the LGS analysis snould be modified to take into account the layout of the panels with respect to the critical portion of tne zone.
- 6. The cumulative suppression distribution function generated in the LGS
! report does not seem to agree with available data.
l
- 7. Suppression probabilistic codeling seems to be very conservative and is not representative of the actual case.
- 8. The LGS analysis doe.s not quantify the uncertainty of the final re-suits. The uncertainty bounds generated are merely judgemental.
Consistent with these conclusions, the following section discusses each item in detail.
i 2.2.2.1 Evaluation of Significant Fire Frecuencies in General locations
, In this part of the LGS analysis, the frequencies of fires in general locations were estimated based on historical fire occurrence data in NPPs.
The general locations for LGS were identified from the Fire Protection and Evaluation Report (FPER). The data base adopted appears to.be suitable for estimating the frequencies of fires in NPPs. The point estimate frequencies 1
calculated for the general locations seem to be reasonable, but the uncer-tainty bounds were not detemined. Tne frequency of fires for the individual fire zones was then calculated using the ratio of the weight of combustible material contained within a zone to the total weight of combustible material 4
. , . . -w .. . + . 4 mm =>% .geme,w-+-e.*,wwe6- es e *~ ert h**'d="*~*- ***s.w ++ es , see ++e -^ * - * *-"* --"*- -e*-- *
- a,
2-72 in the general location. 'There is no justification for using tnis ratio for estimating the specific zone fire frequency. However, the results of these estimations were used for the systematic identification of critical fire zones j through screening analysis, rather than the detailed fire risk assessment.
For the detailed fire risk assessment, the fire occurrence frequency witnin each zone was estimated based on three different mechanisms of fire initiation. These are: self-ignited cable fires, transient ccmbustible fires, and distribution panel fires. Following are ccmments regarding eacn type of fire occurrence frequency estimation.
2.2.2.1.1 Self-Ignited Cable Fires Three incicents of cable raceway fires nave been reported in tne cata base i for NPPs. Two of tnem spread beyond one cable tray and were estimated to burn for 30 minutes before being extinguisnea. The LGS report indicates that all these cable fires were attributable to bad cable splices and underrated
- cables. Based on a review of the LGS data given in Tables 0-1 and 0-2 of their submittal, incicent 43 (Table 0-1) does not seem to nave been caused by underrated cables or bad splices. Hence, we cannot agree with the five-foId reduction of self-ignited cable raceway fire frequencies as indicatec in tne
- LGS report based on the Limerick protection measures and flame retarcant cables. It appears to us tnat a three-fold reduction should have been implemented for cable raceway, self-ignited fire frequencies in the Limerick plant.
- In order to estimate the frequency of fires within ene individual fire l zones, the frequency per raactor year was weignted according to tne fraction of cable insulation weight in that zone to the total cable insulation weight in the control structure and reactor building. We cannot follow the logic behind this fractional weignting factor. In our view, tne nummer of -
conductors and splices, tne voltage / power ratings, geometric factors, etc., f may be more suitable for weighting the frequency of fire in each fire zone, I rather than simply the insulation weignt. This matter of concern indicates l
f that large uncertainties are present in the fire frequency estimates of i
various zones. !
, . , _ . - , , . - - - , . . , - - - - , r-. _- . - - - -
2-73 2.2.2.1.2 Transient Combustible Fires Three types of transient ccmbustible fires were included in the analysis.
The quantity'and the a'rea of each type of transient combustible were con-sidered to be fixed. Tne state of the art for fire risk analysis is to con-sider various quantities of transient combustibles each with an assigned probability distribution. Hence, the effective damageability area and the critical propagation time for transient combustible fires are expected to be in the form of a distribution. Considering tnat no data are available, tne frequency of fires for transient ccmoustibles estimated in the LGS report seems to be reasonable.
2.2.2.1.3 Power Distribution Panel Fires The estimated frequency of fires cccurring in power distribution panels was estimated cased on five reported fires tnat have occurred during 564 years of reviewed U.S. LWR experience. The point estimate of fire frequency witnin a power distribution panel was derived from tnese data and seems to be reason-
- able. ,
, 2.2.2.2 Screening Analysis A systematic approacn is used in the LGS report for identifying the critical fire areas. In this approach it is assumed that upon the occurrence of a fire in a zone, all the equipment and cables in tnat zone will be disaDied. The core melt probability was then recalculated and was multiplied by the frequency of fire occurrence in that zone to provide a measure for screening analysis. Using this approach, the LGS fire analysis appe,ars to have identified all the critical areas in the plant. The quantitative reassessment of their results are beyond the scope of this review. Based on
[ our review of the FPER and the use of engineering judgement, the critical fire areas identified by the LES report seem to be reasonable.
e b
. M O
4 1.
. _ . . . . _ . _ - . . _ - . _ . - - ,- --~.- ~ . - -
f
~ ~
2-74 2.2.2.3 Probabilistic Modeling of Detection and Sucaression The probabilistic suppression / detection model used in the LGS study in the form of a cumulative probability distribution to predict the probability of failing to extinguish the fire within a time interval is based on actual plant data for automatic detection and manual suppression.' It is indicated that the data base for cable insulation fires reported by Flming, F al.[22] was used to construct the suppression probability distribution. This document was reviewed and the cumulative suppression / detection was reconstructed based on our interpretation of the data. A comparison of the curve constructed by BNL with the curve given in the LGS report is made in Figure 2-2. Table 2-1 presents the data used by BNL. It is our understanding that in the LGS esti-mate of the suppression success probability, the self-extinguished cabinet fire incidents were included. In our opinion, the LGS report should not :ake credit for the data on self-extinguished cabinet fires when estimating the suppression success probability for the cable raceway fires. In addition, the LGS report constructed the cumulative suppression probability distiribution with the assumption that the longest suppression period is 1.3 hrs (based on the longest suppression period observed in the data base). We feel it is more appropriate to obtain a distribution fit to the data rather than the " eyeball fitting" used by the LGS report. In our analysis, the legnomal, exponential, and Weibull PDFs were considered as the likely. candidates. The chi-squared goodness of fit for both the BNL and the LGS data indicates that the para-metric Weibull distribution is the best choice. A cumulative Weibull distri-bution F(x) can be defined by two parameters, n and e, and is given by F(x) = 1 - exp (-x/e)" (2.10) i The estimated (a,n) values for the BNL and the LGS data are (0.615,13.5) and (0.458,6.83) respectively. A comparison of the original LGS curve with the modi fled LGS curve and the BNL curve is given in Figure 2.2. Between the time interval of 30 to 75 minutes, the curve I obtained by the Weibull fit to the LGS data is essentially the same as the curve II,-obtained by the " eyeball fit" in the LGS report. Outside the above interval, the difference observed
, is not expected to result in any significant change in the final fire PRA results. However, the curve III obtained by the Weibull fit to the BNL data shows that the LGS estimate of suppression success probabilities is higher at all times than curve III.
y% w-,,,& --,,,,w - , - -
y ~< --o, ,---w_. <4 - , - m-+ -
9-4
l 2-75 1 In the LGS report, similar to other conventional probabilistic risk as-sessments, it is assumed that fire growth and suppression are two independent processes, and they are treated separately. This is one of the most important deficiencies of existing fire risk analyses which usually results in very con-servative values for fire-induced risk. The interaction between the fire growth and suppression will be discussed qualitatively in Section 2.2.2.4.
The probability calculated by the LGS report for fire propagation out of a distribution panel was considered to be 1/25 = 0.04. This estimation was made based on the data base which indicates that all of the five reporteu distribu-tion panel fires were self-extinguished and none of them propagated out of the panel. It was conservatively assumed that one of these fires had the poten-tial to propagate. In addition, a five-fold reduction was considered, based
' on engineering judgement, to give credit to the IEEE 383 qualified flame-retardant cable insulations. This reduction may not be justified. The com-bustibility of cable insulation can best be described through the sensitivity j of the cables to various themal environments, expressed as the change in generation rate of combustible vapor per unit change in the flux received by the combustible. This value is usually denoted by "S". The value of S is, l 0.17(g/kg) for EPR/Hypalon and 0.22 (g/kJ ) for PE/PVC cable insulation [23,10] ,
Hence, a maximum factor of 2 may be creoited because of flame retarcant caole insulations.
Additionally, during a visit to the plant, it was noted that some of the panels are air-tight. For these panels, we feel the probability of fire
[ propagation is negligible, therefore the value used in the LGS report is conservative. For panels with louvers or openings, the value used in the LGS -
report may be unconservative. In general, we do not expect the impact of panel fires to change appreciably if more detailed analyses were perfomed.
2.2.2.4 Py babilistic Modeling of Plant Damage State Generally, three stages of fire growth and corresponding states of shut--
down equipment damage were evaluated in the analysis. The first stage con-sidered is damage to components in the immediate vicinity of the source of fire. The second stage considered is fire growth to adjacent unprotected
%.l l
,,, . . . .. . - . . - - . _ . . . - - - - - - . - , -- - - . - - - - - + + = - ~ - - -
--1m w a --
g- g --w- .y y---- er w- -
-g-iw---- -
~
2-76 cable raceways separated from the initial fire by minimum separation criteria (5 ft. vertically and 3 ft. horizontally). The third stage of fire growth re-presents fire of sufficient severity and duration to damage the mutually re-dundant shutdown methods which may have cabling with a separation distance of at least 20 feet or protected by fire barriers. There are certain inherent assumptions in the analysis. These are:
- 1. The rate of fire growth is not dependent on the suppression.
- 2. A 20 ft. seoaration is considered to be equivalent to 1/2 hour fire barriers (1 in, thick ceramic blanket).
- 3. Cable raceways separated by a distance of 40 ft. or more from the fire source were considered undamaged by the fire.
4 It was assumed that long-term heat removal systems -not required until 20 hours2.314815e-4 days <br />0.00556 hours <br />3.306878e-5 weeks <br />7.61e-6 months <br /> into the fire-induced transient could be recovered by oper-
~
ating valves manually and operating pumps locally. The probability of failure of the operator to perform these recovery actions was con-sidered to be 10 times greater than human errors ascribed to internal events.
Given these assumptions, the LGS report analyzed the impact of fire in various critical zones as identified through the screening analysis.
Identification of various equipment damaged in different fire growth stages could not be verified by the BNL review group due to lack of information and time limitations. However, based on a limited identification of various critical components and systems in different fire zones by means of the information gathered frem LGS-FPER and the plant visit, we concluded that in most cases, the LGS report identified the components properly. There are two exceptions th'at are given as follows:
- 1. In Zone 44, BNL has identified seven distribution panels and motor control centers. These are distribution panels 100201, 100202, 10D203 and motor control centers 108211,108212,10BV215 and 108216. We have also concluded that a fire in distribution panels 100202 and 100203 l
l
- l. .
9
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W 7 -e
2-77 would affect the operation of the HPCIS, and a fire in distribution panel 100201 would affect the operation of the RCICS. Hence, there are three critical panels in_this area. The LGS report indicates that there are six distribution panels in this area and only two of them are critical (100201 and 100203).
- 2. During the plant visit, it was noted that in Zone 47, General Equip-ment Area, there is a boostar fuel pool cooling pump in the vicinity of the northeast corner, which is the critical area in this zone.
This pump was not identified in the LGS report. Therefore, its poten-tial for intitiation and progression of fire causing an adverse effect on the cables in this area was not considered.
Before presenting our cccments on each critical fire zone, it is approp-riate to discuss further the inherent assumptions used in the LGS report as mentioned earlier in this section. More specifically, we would like to dis-cuss the interacting nature between fire growth and suppression activities. In the LGS report, it was assumed that a fire can progress regardless of sup-pression initiation, but terminates with some probability after an expected time which is required for successful suppression. The lack of physical med-eling for the suppression phasa of a fire scenario appears to be one of the weakest links in the analysis. We are aware that this deficiency exists in other' fire PRAs and it seems to be a conventional practice, usually resulting in very conservative estimates for fire impact on equipment and cabling.
While reevaluation of the results given in the LGS report, taking into account proper detection and suppression modeling, is beyond the scope of this re-view, it seems necessary to discuss the basis for such analysis.
' In t!ie analysis of a fire scenario, initiation time for detection and sup-pression is of great importance. Detection and suppression can be achieved
' either manually or automatically. In a detailed fire PRA, both detection time and suppression initiation time should be expressed in the form of probability distribution function (pdf). For the automatic suppression and detection re-sponse, some design charts are available which graphically, or through some equations, detamine the response time vs. the spacing, ceiling height, and heat release rate. [24,25,26] If detailed fire growth modeling, with the as-sociated uncertainties of various fire parameters, is available for a specific w ..
e m . , , . < -%se,u-a.~ee--+-; *.*u
- mym- e=.* -.em .. ee-=ww. =. _
. o 2-78 scenario, the detection and suppression response may be directly estimated In the form of pdfs. If detailed fire growth modeling is not available, a generic response can be considered by assuming the two extreme fire growths -
(slow, fast)asdefinedinRef.[24]. In this case, the lower and upper bounds for response time may be determined assuming fast or slow fire growth, respectively. These bounds may be used to define a pdf for'the response. The response time for the inititiation of the manual suppression may be estimated by means of available data on response time during fire drills and scme en-gineering judgement. The modeling of a fire growth during the suppression phase can be very complicated depending on the governing mechanism of the process (heat removal, chemical reaction, oxygen removal.) However, for the purpose of fire pRAs, a ccmbination of simplistic models, coupled with em-pirical correlations, may be used. For example, the' effect of sprinkler sys-tems on fire growth may simply be modeled in the form of global energy bal-ance.[27]
In conclusion, the time in which fire can retch various stages of growth is dependent on suppression initiation time. There is a strong belief that fire cannot grow significantly once the suppression has begun. In the LGStre-port, it is conservatively assumed that probabilities of various stages of growth can be determined using the time period for the completion of success-ful suppression, rather than the initiation of supp'ression. This is a very conservative assumption and at present the effect of this conservatism on the final results cannot be evaluated.
2.2.2.4.1 Zone Soecific Comments In addition to the generic comments made in previous sections, there are
, additional zone specific comments that may impact the results of the fire PRAs
! given in the LGS report. These comments are mostly associated with the layout l of different components in various critical zones and they are based on the review of the FPER and the plant visit.
l S
.. -. % ..-%---- -g--e+ -w *w===e--~e -~~----i. --e- - - - - - - + * * * - - m--- ==~---aens+-em-- -e ===. - = m.s
2-79
- a. Zone 44, S4feguard Acces Area (CH=36f t. , A=8930 f t.2, ASD=357.2, S=M).* In this zone, th...e are a total of seven motor control centers (NCC) and distribution p 1els. Four of these panels are located close to the critical corners. These are distribution panels 100202 in SW, 100203 in NE, 100201 in , and MCC-108211 in SW (Drawing M118, Rev.)
The event tree associate with the panel fires should be modified.
- b. Zone 45, CRD Hydraulic E f pment Area (CH=25f t. , A=12860 f t.2, ASD=676.8f t, S=M/A) The oly critical panel which is located in the NE corner is the MCC-1CB. 4 The other panels are not located in the vicinity of the NE corne. ; Drawing M119, Rev. 19). The event tree as-sociated with the panel t es should be modified.
- c. Zone 47, General Equipmer Area (CH=7, A=9800 f t.2, ASD=490 f t. , S =
M/A). Based on the draw- 1 M-120, Rev.18, none of the distribution panels, load centers, or Jtor control centers are located in the vicinity of the critical : corner. Theref ore, the event tree as-sociated with panel fires n this zone should be modified. The only component that may result n a fire hazard and which is located in the NE corner of this zone, i a booster f uel pool cooling pump. .
= CH is the ceiling height,, A is t i floor area, and ASD is the area per smoke detector. The "S=M" represents nu nel suppression where "S=A" represents automatic suppression.
- g.
-,s... . . . - . - - , - - - - - ~ - - - - - - ' ~ ' ' ' ' ' ~ ' ' ' ' ' ~ ~ ' ~ ~ ~ ~ ~
-^'*TG.**'
2-80
+
2.2 REFERENCES
- 1. Siu, N.O., "Probabilistic Models for the Behavior of Compartment Fires,"
School of Engineering and Applied Science, University of California, Los Angeles, Ca., NUREG/CR-2269, August 1981.
- 2. Siu, N.0., "CCMPBRN - A Computer Code for Modeling Ccmpartment Fires,"
School of Engineering and Applied Science, University of California, Los Angeles, Ca., UCLA-Et:G-9257, August 1982.
- 3. Mitler, Henri E. and Emmons, Howard W., " Documentation for CFCV, the Fifth Harvard Ccmputer Fire Code," Harvard University, Cambridge, Ma.,
October 1981.
4 .~ Quintiere, J.G., " Growth of Fire in Building Ccapartments," ASTM Special Technical Pub. 614, 1977.
- 5. Tatem, P.A., et al, " Liquid Pool Fires in a Cceplete Enclosure," 1982 Technical Meeting, the Eastern Section of the Cembustion Institute, i
Atlantic City, N.J. , December 14-16, 1982.
l 6. Zukowski, E.E. and Kubota, T., "Two Layer Modeling of Smoke Movement in Building Fires," Fire and Material, 4, l',1980.
- 7. Delichatsios, M.A., et al., " Computer Modeling of Aircraft Cabin Fire.
Phenomena," FMRC J.I. OGONI.BU, Factory Mutual Research Corp., Norwood, Ma., December 1982.
- 8. MacArthur, C.D., " Dayton Aircraft Cabin Fire Model Version 3," Vols. I and II, University of Dayton Research Institute, 1981.
- 9. Sumitra, P.S., " Categorization of Cable Flammability, Intemediate-Scale Fire Tests of Cable Tray Installations," EPRI NP-1881, Electric Power Re-sear:h Institute, Palo Alto, Ca., August 1982.
- 10. Tewarson, A., Lee, J.L, and Pion, R.F., " Categorization of Cable-
. Flammability, Part 1: Laboratory Evaluation of Cable Flammability Para-meters," EPRI NP-1200, Electric Power Research Institute Palo Alto, Ca.,
October 1979.
- 11. Towersnn, A., " Fire Behavior of Transfomer Dielectric Insulating Fluids," 00T-TSG-1703, prepared for U.S. Dept. of Transportation, Trans-portation Systems Center, by Factory Mutual Research Corp., Nomood, Ma.,
September 1979.
~ 12. Lee, J.L., "A Study of Damageability of Electrical Cables in Simulated Fire Environments," EPRI NP-1767, Electric Power Research Institute, Palo Alto, Ca., March 1981.
- 13. Tewerson, A., "Physico-Chemical and Combustion / Pyrolysis of Polymeric
- Materials," N85-GCR-80-295, prepared for U.S. Dept. of Commerce, National
-Bureau of Standards, Center for Fire Research by Factory Mutual Research l Corp., Nomood, Ma., November 1980. '
.j L 3, -...,( ,-== . . -.e zw s=r e s = % *** - ^ - * " ' " ' ~ " '
- -6'Nd 4# -* < '
. . - , ~ - .- ,- .
2-81 1
14., Carslaw, H.S. and Jaeger, J.C., " Conduction of Heat in Solids, 2nd Ed.,"
0xford Clarendon Press,1959.
l
- 15. Alpert, R.L., " Calculation of Response Time of Ceiling-Mounted Fire Oe- l tectors," Fire Technology, Vol . 8,1972, pp.181-195. '
- 16. Alpert, R.L. and Ward, E.J., " Evaluating Unsprinklered Fire Hazards,"
FMRC J.I. No. 01836.20, Factory Mutual Research Corp.. .Norwood, Ma. l August 1982.
- 17. Stavrianidis, P., "The Behavior of Plumes Above Pool Fires," a thesis presented to the Faculty of the Depart =cnt of Mechanical Engineering of 4
Northeastern University, Boston, Ma., August 1980.
- 18. Orloff, L., " Simplified Radiation Modeling of Pool Fires," FMRC J.I. No.
OE1EO.BU-1, Factory Mutual Research Corp. , Norwood, Ma. , April 1980.
- 19. Zukoski, E.E., Kubata, T. , and Categen, B., "Entrainment in Fire Plumes,"
Fire Safety Journal, Vol. 3,1980/81, pp.107-121.
- 20. Steckler, K.D., Quintiere, J.G., and Rinkinen, W.J. , " Flow Induced by Fire in a Compartment," National Bureau of Standards, N8 SIR 82-2520, September 1982.
- 21. Newman, J.S. and Hill, J.P., " Assessment of Ex;:osure Fire Hazards to Cable Trays," EPRI-NP-1675, Electric Power Research Institute, Palo Alto, Ca., January 1981. -
- 22. Fleming, K., Houghton, W.J., and Scaletta, F.P., "A Methodology for Risk Assessment of Ma;or Fires and its Application to an HTGR Plant,"
GA-A15402, General Atomic Company, San Diego, Ca., 1979.
- 3. Tewarson, A., "Damageability and Cembustibility of Electrical Cables " ,
paper presented at FMRC/EPRI Seminar, Factory Mutual Conference Center, Norwood, Ma., December 1981.
! 24. SenDmin, I., et al, "An Analysis of the Report on Environments of Fire Detectors," Ad Hoc Committee of the Fire Detection Institute, 1979.
- 25. Newnan, J.S., " Fire Tests in Ventilated Rooms - Detection of Cable Tray and Exposure Fires," E.PRI NP-2751, February 1983.
- 26. Hill, J.P., " Fire Tests in Ventilated Rooms - Extinguishment of Fire in Grouped Cable Trays," EPRI NP-2660, December 1982.
- 27. Levinson, S.H., " Methods and Criteria for Evaluation of Nuclear Fire Protection Alternatives and Modifications," Ph.D. thesis, Rensselaer Polytechnic Institute, Troy, N.Y., December 1982.
'% e*
4
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2-82 l
r
( TABLE 2-1 Suppression Data and Calculations Performed f or Suppression Success Probability Self-Ignited Cable Raceway Fires Index* Plant Name Time to Bring Fire Type of Type of Under Control (hr) Detection Sucoression 58t Browns Ferry 7.0 Automatic Manual 23 Zion 2 1.3 Manual / Automatic Manual / Automatic 25 San Onof re 1 0. 7 Manual Manual 24 San Onefre 1 0.5 Manual Manual 8 Kewaunee 0.5 Automatic / Manual Automatic / Manual 28 Three Mile Is. 2 0. 5 Manual Manual 37 Vermont Yankee 0.5 Automatic Manual 42 Nine Mile Pt. 1 0.05 Manual Manual 46 Oyster Creek 0.05 Manual Manual 27 Trojan 0.05 Manual Manual
- Indices are the same as those in Fleming's report.[22]
tThe fire occu*rrences during the construction phase or those that were self-extinguished and confined to a cabinet were not included. In addition, the Brown's Ferry fire indicated above is not included in our analysis.
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=!
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O 10 2C 30_.40 50 60 70 80 EXTER; \L HEAT FLUX (kW/m2) i Fig. 2.1. Pi ted ignit- 1 of EPR/Hypalon cable under various external l
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~~
Fig. 2-2: Fire suppression model aeg r=w w w .w. -- .%w.. ..-,.wew,--- = .<=-.e- e see- =- *=4r= = --eep.-- ===www-e*==e--e+
- _. -._ r _g... -
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- 1 3-1 1
3.1 Seismic The objectives of this section are to provide: (1) a brief description of the methodology and assumptions adopted in the LGS-SARA (1) report in the i
quantification of seismic accident sequences, (ii) BNL review comments of parti:ular critical areas of the LGS-SARA document. Results based on BNL modifications are also presented whenever simplified estimations can be made to illustrate the effects of tne modifications. This section is divided into two parts. Section 3.1.1 addresses tnose plant frontline systems which are identified in tne LGS-SARA report and the method of quantification by wnich system unavailabilities, including the seismic contributions, are evalue:ad.
Section 3.1.2 summarizes ene seismic event tree approach and tne seismic accident sequence analysis.
3.1.1 Plant Frontline Systems This section is ccmprised of two subsections. Subsection 3.1.1.1 pre ,
sents an overview of the LGS-SARA approach in modeling frontline systems. It also summarizes the assumptions made pertaining to systems and components of the systems in the evaluation of the seismic contribution to the system un-availability. Subsection 3.1.1.2 provides the BNL revisions to the frontline system models and the results thereof. A discussion of the assumptions and the LGS-SARA approach to system fault trees are also included. .
3.1.1.1 Overview of the SARA Aoproach in Frontline System Modeling The system analysis part of the LGS-SARA effort is based extensively on I
tne structure and contents of the LGS-PRA(2). This includes using the LGS-PRA frontline system fault trees in the description of the random failure of the various systems. In addition, these fault trees also provide the basis for the development of the seismic related failures. Finally, the components l
that appear in the LGS-PRA system fault trees constitute, in part, the group of components for which fragility evaluations were conducted.
l t 8 I
.~
- - ~ - - - - ~ ~ - -
..._.2_. i
[
3-2 ,
l LGS-SARA purported to have examined fragility of two groups of com-
, ponents: those tnat are contained in the LGS-PRA system fault trees and those which are identified to have tne potential of significantly influencing the likelihood of core damage from seismic events, such as the reactor vessel and
! other related structure. A decciled discussion of component fragility is presented in Chapter 2.1. These components are then ranked according to the
! acceleration capacity of eacn 1':em; those with a median ground acceleration capacity of greater than 1.56 g were not considered since they are deemed to have a far higher ground acceleration capacity than tnose predicted for the reactor site. Based on this criterion, a final list of 17 components are selected to be used in the LGS-SARA evaluation. Table 3.1.1.
- Each seismic frontline system fault tree developed in tne LGS-SARA an-alysis is made up of two parts
- the first part, wnich leads to the failure of l 2he system, corsists of the random inaependent failures evaluated in the I
LGS-pRA; the second part includes all the pertinent seismic-related failures as determined using a specified criterion. This criterion for inclusion as a 1
seismic-related failure requires that the component appears in Table 3.1.1.
j The random independent failures for each system, as calculated in the LGS-PRA, are treated as a basic event in the seismic system fault tree. For both the HPCI and the RCIC syster, failure of the condensate storage tank (CST) neces-sitates the transfer of the water source from the CST to the suppression pool
- and is included in the fault trees. A total of eight seismic fault trees were developed for tne LGS-SARA study and they include the following
- high pressure coolant injection (HPCI), reactor core isolation cooling (RCIC), low pressure coolant injection (LPCI), low pressure core spray (LPCS), residual heat re-moval (RHR), standby liquid control (SLC), automatic depressurization system (ADS) and emergency power. An example of HPCI seismic system fault tree is t
I given in Figure 3.1.1.
i 3.1.1.2 BNL Revision and Review of Frontline System Fault Trees Fault Tree Approach The inclusion of random independent failures into the seismic fault trees represents a more realistic approach than those that focus solely on seismic- .
related failure events. In some circumstances, these random independent .
. -~ _ . - . - - . . . -- - - . . - _ - . - . . . . - - _ __ . . - _ .
~
3-3 failures wnen coupled with a substantial reduction in the operatior's e.bility to follow procedures due to high stress conditions resulting from an earthquake may contribute significantly to core damage.
i BNL reviewed the LGS-SARA mocularized system fault trees developed for the seismic analysis. These fault trees were prepared based on the list of 17 components identified to be more susceptible to seismic event related failure.
On p. 3-1 of tne LGS-SARA, it is stated that the internal system fault trees provide, in part, tne list of components for wnich fragility functions were developed and that additional items were included when they were deemed to have tne potential for significantly influencing the likelihood of core damage from seismic events. BNL agrees that consideration of only those components that are identified in the internal event system fault trees does not ensure that all important seismic sensitive components have been incluced. Since in the construction of the internal event system fault trees, depending on the levels of details at which the trees are developed, approximations may have.
been made to reduce the complexity of the trees. For instance, in modeling the faults of an injection train, the piping faults could have been excluded in the fault tree. Consequently, wnen it is used in the seismic assessment, dependence on piping failure would not have been properly evaluated. It is not clear from the report that a systematic search was conducted and wnat criteria were used to select,those components for fragility evaluation to ensure that all components sensitve to seismic events are included in the analysis.
In the .nodularized system fault tree approach, inter-system and support system dependences are not explicitly modeled. LGS-SARA did include the com -
mon mode failure of the diesel generators as a means of failing the systems.
Preliminary review of the fault trees appears to indicate that connon mode diesel failure is one of the dominant scenarios to core damage. However, by not explicitly modeling the dependences, other contrioutions to core damage may be lost in this approximation process. The BNL review of Limerick internal event report (7) assessed that contributions from including the support system dependence constitutes a 60% increase. It is judged in the context of a seismic event that these dependence contributions will be quite
insignificant.
i e
- . .. .w ,
,-m --. m + - . - . - ~ -
3-4 In addition to those dependences discussed earlier, there is one de-pendence which involves failure to transfer water source from CST to sup-pression pool. This operation is required for both the HPCI and tne RCIC sys-tem wnenever there is a low CST level. An operator failure to transfer given that a low CST level is likely to affect botn nign pressure systems. This de-pendence should be included to pr operly reflect its impact on the final re-sults.
Electric Power The failure of the electric power system is modeled with tne failures of seven components; namely, two faults leading to 't ne loss of the 440V power supply, three faults resulting in the loss of tne diesel generators, one lea-ding to losing the 4KV bus, and one in the loss of DC power.
For both the HPCI and RCIC systems, loss of control power due to failure of tne DC bus, is assumed to disable the systems. In principle, it is pos-sible to operate the two hign pressure systems in a total blackout condition:
for an extended period of time with the operator manually providing the con-trols necessary. Nevertheless, in the event of an earthquake, BNL concurs tnat the LGA-SARA assumption may be more realistic.
In the LGS-SARA Appendix 8, it is estimated that at 1.0g, no significant damage to the diesel generator fuel oil tanks is expecs'ted and that at ac-celerations somewnat in excess of 1.0g, failure of attachments would be likely.
It was identified in the Indian Point external event PRA review (3) (p, 2.7.1-15) that the diesel generator fuel oil tanks are major contributors to core damage frequency. The data reported in the Indian Point Safety Study 4 I
for the diesel generator oil tank is of a generic nature and the median ground acceleration capacity is estimated to 1.15 g. In light of this information, it is pertinent that the LGS-SARA report includes a more detailed analysis on l the diesel generator fuel oil tank to show that they have the capacity much greater than 1.0g to justify their exclusion from the system fault trees.
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+ 3-5 Human Error i
In the seismic part of LGS-SARA, it was reported that in estimating tne error rates for operator actions required during seismic accident sequences, the probability of failure within a given time scale was increased by a factor of 10 subjected to a maximum probability of 1.0. The basis to this selection of a factor of 10 comes from the fact that an earthquake of sufficient intens-ity to damage reactor systems will initially disturb the performance of the operators and raise doubts in their minds about the performance of instru-mentation and controls. The eartnquake may also lead to component failures that are not normally encountered in plant operations and therefore, may require innovative actions on the part of the operators.
It is BNL's judgment tnat during and subsequent to an earthquake, tne operators ability to follow procedures, to diagnose problems or to,take cor-rective actions depends on tne intensity of the earthquake. Given the limited information that is available in this area, it is often difficult to quantify the likelinood of failure under these unusual circumstances. However, one would expect tnat an increase in the human failure probability is warranted.
Moreover, tnere are three factors wnich also play a significant part .in determining the human failure probability. One of them is tne avail'ablity of
- reliable instrumentation. Subsequent to an earthquake, with alarms and ar.-
nunciators sounding, it may be difficult for an operator to adequately assess the true plant condition given tnat some instrumentation might give erroneous information. This is by far a more cnallenging situation and it significantly increases the complexity of the situation that confrents the operator. This
, . . . ._ _ type of scenario may cause two types of human failure in addition to those normally considered in the LGS-PRA; they are namely: 1) as a result of false instrument rea' dings, tne operator is misled and follows the wrong procedure in securing the plant, 2) as a resul* of wrong and confusing information, the operator may be misled to err in an error of commission.
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9
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. _ . . . . . ~ _
3-6 The second contributing element to tne subject of human failure that was not addressed in LGS-SARA is the impact of aftershock upon tne performance of tne operator to discharge his responsibility. Based on seismic data, the probability of occurrence of aftershock decreases fallowing an expociential type of pattern; in other words, the aftershock is most likely to occur right after the first quake and that likelihood decreases as a function of time in an exponential type manner. If an aftershock occurs uitnin tha time frame wn:n operator action is critical, it may further impede his ability to respond to the demands of thc plant.
The thirc area entails the subject of display in:.trumentation, wnicn is intended to provide tne operator pertinent information to help nim to under-stand the status of the plant. Display instrumentation could be in the form i
of lights, chart recorder, annunciators, alanns, etc. In the event o'f an eartnquake or an aftersnock, the failure modes of tne display instrumentation l could be: 1) display information wnien is inconsistent of other indicators 2) .
loss of display function. LGS-SARA should furnish a discussion on this sub-ject to ensure that failure of display instrumentation has been investigated i and is deemed not to impact significantly on final results.
BNL concludes that the factor of 10 increase in the human failure probability in some instances may be reasonable while it may also prove to be 1
conservative or non-conservative in otner instances depending on the situa-tion wnich confronts the operator. An example of how the absence of readily available and reliable information can affect the operator's ability to pursue th] proper course of actions is given for the CST. The CST is calculated to have a median ground acceleration capacity of 0.24 g wnich is comparatively H Iow in light of the other component values. It constitutes one of the two water sou'rces from which the HPCI and the RCIC take suction. Failure of the CST would necessitate a transfer of the suction from the CST to the sup-pression pool. As for the HPCI, this transfer process is automatic, i.e.,
giv:n that tnere is a low CST tank level, an automatic switch over will be initiated; however, for the RCIC, this transfer is a manual operation. The
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. , i 1
3-7 ,
failure mode of tne CST water lev'el sensors given that the CST is failed was not addressed in LGS-SARA. Nevertheless, one could postulate the following:
- 1) tnat despite tne failure of the CST, whether it be rupture or toppled over, the level sensors give a low level reading; 2) that in the failure of the CST, tne level sensors are damaged and that erroneour or misleading in' formation re-sults. Preclusion of one or the other would require a more detailed investigation of tne failure modes of botn the CST and the level sensors.
If scenario 2 occurs, it implies that the information given to the oper-ator is misleading, and hence in tnat case, the failure probability for the operator to respond preperly shculd be close to unity and should not be based on an arbitrary rule of thume - a factor of 10. It so happens that when tnis factor of 10 is applied to tne HpCI and the RCIC transfer from CST to sup-pression pool, the human failure probability is unity. But there are other human operations within these trees as well as other system fault trees which should be examined on a case by case basis to determine the respective human failure probability, for instance, manual failure to re-start system, failur'e to transfer service water, etc. A detailed discussion of this impact upon system unavailabilities is deferred to the next sectiori.
Final'ly, it is important to note that LGS-SARA did not convey to the re-viewers that the increate in human error was applied consistently to all the
" pertinent basic numan events. BNL reviewed the LGS-PRA system fault trees and identified a number of manual operations which are omitted in the seismic sys-tem fault tree consideration, for instance, the manual failure to initiate HPCI, failure to manually initiate the LPCS and others. A more detailed investigation of the system fault trees should be conducted and pertinent findings on manual errors should be included in the modularized system fault trees. .
Relay Chatter
, It is reported in LGS-SARA that low accelerations cause a momentary inter-ruption of control circuits and. power supplies (typically from relay-contact chatter); however, relay-chatter is dismissed as a means of leading to system
, failure based on the fact that the operator can intervene and reset the cir-l q.; cuit, hence restoring the system to 1'ts initial state.
I"
- y. , , , ,,
3-8 It appears tnat tne question at issue here is not wnether the relays will chatter or at wnat acceleration will they begin to chatter, but what credit should be given to the opeator to reset them if relay cnatter occurs. If in one part of LGS-SARA, it is maintained tnat human error, in the event of an eartnquake, should be modified Dy a factor of 10 to reflect tne increase in stressful conditions, it would seem consistent that these human responses to reset relays be given the similar treatment in assessing their failures. BNL is of tne opinion that given there is relay enatter, failure on the part of tne operator to reset would result in tne equivalent of a relay failure.
If one wants to quantify the impact of relay-chatter upon the system f ailure, then one would have to ascertain relay fragility information for the various kinds of relays. In the Indian Point study (4), it is stated tnat
' relay-cnatter occurs at 1.2 g and that it presents no major difficulty The .
SSMRP data (S) snow tnat enatter occurs at as low as 0.759 (spectral accel-eration). Moreover, for certain relay chatter which results in a breaker ,
trip, reset of the system may be readily possible at tne control room; nowever, there are those relay trips which may require resetting at local panels and this causes a substantial increase in the failure probability of human to reset. It is important that LGS-SARA provides additional analysis on the fragility of relay chatter and its impact upon various systems. Failure of numan action required to reset relay, heace leading to relay failure, should also be considered. '
Finally, there is the underlying question that, in view of the different relay trips, the operator is presented with a scenario for which he has not been trained ano for wnich no procedure has been written, wnat is the probabil-ity that he will perform adequately to reset the relays. Attempts to answer this question should be furnished in LGS-SARA to support the premise that the operatoqcanindeedinareasonabletimeresettherelaysandrestorethe system.
An example to illustrate these points can be found in the SLC fault tree.
There are two relays per SLC pump, for example, K4A and KSA for train A. K48 and KS8 for train B, etc. If chatter causes these relays to terminate the' l
. - _ ..=ew ego
, **I M9
3-9 operation of the three SLC pumps, tnis will lead to a direct failure of the SLC system. Furtnermore, in tne redundant reactivity control system, relay chatter may cause the failure of all APRM cnannels wnicn in turn will result in failure to initiate the SLC explosiva valves and the SLC pumps. In an ATWS accident event, the time available to an operator to respond to these enal-lenges is also significantly reduced to the order of minutes. In light of tnis information, tne impact of relay cnatter upon the SLC system snould be evaluated in more detail.
Transients A list of tne LGS-SARA tean random faiTure values and the nomenclature is given in Table 3.1.2; the first column of values are tnose given in tne LGS-SARA report. Tne second column tabulates the values used in tne internal ev-ent risk assessment study, LGS-PRA. The third column denoted by NUREG/CR-3028 enumerates those values generated by BNL in the review of the LGS-PRA. The last column represents values that BNL believes snould be used in the LGS-SARA study. It is quite obvious that between the first and second column, differ-ences in values can ce noted. Despite the fact that little explanations are furnisned 11 LGS-SARA to address the differences for both hign pressure sys-tems, these differences are miniscule. But for the low pressure system (V)
. and tne manual depressurization (X) function a more detailed discussion is warranted.
As stated repeatedly in LGS-SARA, the seismic evaluation was done based extensively on the LGS-PRA; therefore, it is reasonable to assume unless noted otherwise that the nomenclature used would also correspond to that of LGS-PRA.
The manual depressurization function, X, denotes the failure on the part of the operator to depressurize the reactor in a timely manner using the Auto-I matic Depressurization System (ADS). The low pressure injection function (V) f represents either the failure of the ADS hardware or a simultaneous failure of l the LPCI system and the LPCS system. In the LGS-PRA, the X function unavail-ability is calculated to be 2 x 10-3; the V function value is estimated by BNL based on the LGS-PRA unavailability of the LPCI and LPCS systems given tnat tnere is a loss of offsite power and a failure to recover offsite power
.[ j to be 2.65 x 10-4 According to the information provided on p. C-15 of
- ' u - voes wan
- -a. -W - -.
1 3-10 ,
Table C-6 of the LGS-SARA, it appears tnat tne V function ~ defined in tne report only consists cf the LPCI and the LPCS system; this notion is further confirmed in the Boolean expression of X = XR+A shown on p. C-14 of Table C-5. Xg is defined in the report as the random failure of X and A, as the loss of electric control and motive power. Since the manual action to depres- s surize the reactor does not require electric control or motive power, hence, ,
it is possible to argue tnat the haroware failure of the ADS is lumped with the X function without mucn impact on the function unavailability. However, this is not consistent with what nas been presented in the LGS-PRA, and tray result in misleading conclusions of dominant sequences. The impact of properly including the ADS hardware failure within the V function for various accident sequences will be addressed in Section 3.1.2.
Quantification of the RHR system with the loss of offsite power and no '
, recovery was not performed by BNL nor by PeCo and hence no value' is reported in LGS-PRA and NUR.EG/CR-3028. The most substantial tr.cesase between'tne LGS-SARA and NUREG/CR-3028 internal event values occurs witn the V function -
a factor of 3.7 followed by a factor of 3.0 increase for.tM X function.
The comon mode diesel generator failure probability was reported in the earlier revisions of tne LGS-Pita as 1.88 x 10-3 and it is for this reason that this value was used in the NUREG/CR-3028. Revisions received subsequent to the report of NUREG/CR-3028 modified the unavailability to 1.08 x 10-3 '
claiming that it was a typographical error in the earlier versions. In LGSe SARA, a diesel generator common mode failure mean value of 1.25 x 10-3 was reported. BNL agreed that the 1.88Ix 10-3 value is overly conservative.'
Recsently, there have been studies (6) conducted to attempt to better evalu-ate the diesel common mode unavailability, and values lower than 1.0 x 10-3,-
have been suggested. As for tai 3 review, BNL will use the ;1.25 x 10-3 for l comparison purposes. '
i
, The increase in numerical values for the HINIA and RIN3 is due' to tee -
following: HINIA and RIN3 represent failure to provide flow from suppression pool given that tne CST water is unavailable. The major difference-between the two events lies in the manual action required to perform the operatton for -
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~
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3-11 the RCIC system, FSAk p. 7.48. Consequently, if a factor of 10 increase is assumed, the manuel error for failure to transfer becomes unity and dominates
- tne failure of the RCIC system. Because of the automatic transfer function in the NPCI, a simila'r increase in the manual error results only in minimal in-crease'in'the system unavailability. If, instead, a numan factor of 7.5 is used, tne RIN3 will be 0.15, wnereas, HINIA would remain unchanged.
/. , Ar$6ther irta,jor change that is evident based on tne factor of ten increase
~
is in the manual depressurization function. This increase results merely from applying th'e'numan error factor of 10 to the NUREG/CR-3028 value of 6x10 3 ItabpearsthatforHINIA,RIN3andX,theincreaseduetothehuman error factcr was-not included in tneir values as it should be.
t Antici$atec Transients Witnout Scram .
~
In tne event that an earthquake odcurs which results in an ATWS, LGS-SARA analysed the sequence using a loss of offsite power ATWS event tree. A set of mean failure' values that wastur,ed in the LGS-SARA analysis is shown on Table 3.1.3. There are~ four columas of values in the Table, the first three of them present values used in the LGS-SARA study, the LGS-PRA and the BNL internal
/
o event review, NUREG/CR-3028, respectively. The last column represents values which BNL believes should be used in the LGS-SARA analysis. It should be I
pointed out that the values presented in Table 3.1'.3 are representative j' nume,ers; one should refer to these reports for more detail information.
LGS-SANA values for both the HPCI and RCIC failure values are in general hb lower than thesd of t'he LGS-PRA and NUREG/CR-3028. The increase for RCIC, s RR 'Is abo'ut 'a factor of 6.6 times. As for the ADS inhibit function, the
- . LGS-SARA val'ae is 8.0 x 10-3 versus 2.0 x 10-2 from NUREG/CR-3028; another
! factor of 10 increase due to the intense stress level for the operator brings.
the final value (last column)'ta 2.0.x 10-1 There is no disagreement on l
the va' lues of UH as assessed by BNL ano LGS-SARA. Little increase is noted j -
between'the LGi-SARA and BNL ' values for the }SLC system, however, LGS-SARA l value is about a factor of 10 larger than the LGS-PRA value. W2 was re-ported in LGS-SARA to 'be 0.1 rather than the 0.14 used in the LGS-PRA. The,
- s. _ ,
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diesel generator common mode failure, HINIA and RIN3 failure values are described in the previous paragraphs.
The value selected for the mechanical failure of the scram system
, increased to 1.5 x 10-5 The variable PCR is defined in the text of LGS-SARA to nave a value of 0.2. No description is provided as to wnat PCg is.
Failure'to scram is defined in LGS-SARA as:
Cg = (1-PCR ) CR + PCR (33*35
- S7)
'. telephone conversation with PeCo revealed that the PCg is a judgment factor applied to the seismic failure of the reactor internals arid CRD guide
- tubes,,see Figure 3.1.2. PeCo stated that in the event of an earthquake, only l 20% of the time would the failure of the CRD guide tubes or tne reactor inter-( nals cause a failure to scram. Since there is insufficient information on how this PCR value is obtained, it is difficult for BNL to render some sort of judgment on its validity.
Another area of concern is in the treatment of random failure to scram; BNL believes that, given that there is a challenge to the scram system, the failure to scram probability should no't be weignted by a factor of (1-PCR )*
h
,0.8. Also in the telephone conversation, PeCo explained that they attempted to preserve the scram failure probability from the 0.8 reduction by increasing the scram failure probability from 1.0 x 10-5 to 1.5 x 10-5 It is sug-gested tnat a more detailed documentation of~ these points by PeCo be provided in LGS-SARA For the purposes of sequence quantification to be presented in the next section, failure to scram is defined by BNL as folicws:
1 i
Cg = CR+S3+SS+S* 7 Finally, it is suggested that a detailed discussion be provided in l
i LGS-SARA to identify and reconcile differences in the random failure values used in LGS-SARA and LGS-PRA.
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l t 3-13 3.1.2 Accident Sequence Analysis
, This section addresses the definition of accident sequences and the quantification of core damage probability given that there is an earthquake.
Section 3.1.2.1 briefly describes the approach and methodology used in 4
LGS-SARA for accident sequence definition and core damage quantification.
Section 3.1.2.2 contains results of the BNL review.
3.1.2.1 Overview of LGS-SARA Accident Secuence Analysis LGS-SARA examined various fragility , estimates (provided in Appendix C) and concluded tnat the offsite power system was most susceptible to an eartnquake wnich, wnen failed, would result in an initiating event. Failure of pipes and valves causing an initiating event is dismissed as highly improb-able in light of the significantly greater capacities of these components. It is for this reason that LGS-SARA maintains that the frequency of a seismically induced LOCA (be it large, medium, or small) is quite insignificant. The simultaneous occurence of an earthquake and a random LOCA event is also estimated to be a few orders of magnitude smaller than the loss of offsite power event. Therefore, only the seismic-induced loss of offsite power was investigated as a credible initiating event.
The event tree method was used to define the accident sequences. A total of three event trees were developed: the first event tree depicts the suc-cess or failure of a number of critical functions whose operation or inopera-tion greatly affects the analysis to be followed, see Figure 3.1.3. This tree is made up of five functions, namely, the seismic event initiating frequency, reactor pressure vessel, reactor and control building, and reactor scram.
i Failure of the reactor pressure vessel given that there is an earthquake leads directly to core damage. The nature of the failure was identified to be initially the failure of the vessel supports which, in turn, results in severing of all four steam pipes. To mitigate such a breach of the reactor
- coolant boundary is far beyond the capability of the ECCS.
er l
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- s4- w- -- _____mh ---m.. .es. has
3-14 t -
Given that the reactor pressure vessel stays intact, failure of tne re-actor and control building will result in core damage regardless of whether there is a successful reactor scram or not, Sequences 4 and 5. If, however, the reactor building does not fail, then failure of offsite power coupled with either successful or unsuccessful scram would lead to transfers to Figure 3.1.4 and Figure 3.1.5 respectively. -
The event tree presented in Figure 3.1.4 is identical in structure to tnat of the internal loss of offsite power event. Systems wnien are re-quired to mitigate the event are assessed and accident sequences are defined.
In Figure 3.1.5, the mitigation of an ATWS event is presented. It is again identical in structure to tne one given in the LGS-pRA for loss of off-site power.
Inputs to these event trees for individual systems are based upon the modularized system fault trees and a discussion of these trees is provided in Section 3.1.1. The quantification of these event trees were performed using-the computer code SEISMIC. The Monte Carlo method is used in the code to simulate the failure probability of seismic and random failure of components and accident sequence frequency is then calculated based on the Boolean expression inputed for that particular sequence. Median and mean values, and confidence levels of the sequences are also evaluated and those fcr the dominant sequences are reported.
3.1.2.2 BNL Review of Accident Sequence Quantification BNL reviewed the event trees and assumptions which enter into the de-velopment of these trees. Review comments are presented in this section. A number of areas were identified which warrant further discussions and they are also presented in this section. As a result of the .evisions made to the modulari, zed system fault trees, estimates of their impacts on respective accident sequence core damage frequencies are described.
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. 1 3-15 Methodolog The event tree-fault tree methodology employed in the LGS-SARA represents a widely practiced approacn used witnin the nuclear industry today to assess accident sequences and core damage frequencies. BNL agrees that it is adequate in evaluating risk indices within tne context and requirements of today's ri< assessment studies. ,
The LGS-SARA analysis based extensively on the approacn and results of the LGS-PRA. Two event trees from the LGS-PRA were adopted to analyse the seismic initiating event. They are the transient and ATWS loss of offsite j power trees. While BNL agrees that these trees will model the loss of offsite power event adequately if caution is exercised in addressing the depencent failure of components due to an earthquake, additional information should be included in LGS-SARA to establish the basis why the seismic event evaluation can be based extensively on the internal event analysis. In other words, justification should be presented to show that external event accidents do not warrant separate event trees to model the different scenarios. Rationaleori wny tne LGS-PRA event trees were used should reflect these concerns.
'As described in Section 3.1.2.1 of this review and in Chapter 3 of LGS-SARA, the loss of offsite power due to failure of the switchyard ceramic insulators (median ground acceleration capacity of .20 g) was identified to be the major initiating event ' contributor.. Failure of the reactor and control building and of tne reactor pressure vessel has also been included in the
- consideration of initiating an accident event. BNL agrees that these are important initiating scenarios that should be investigated.
- Nonetheless, it is not clear from what is reported in LGS-SARA that the search for initiating events went-beyond those components and some structural In particular, it is not obvious that effort was devoted to ex-members.
amining the non-safety related equipment or equipment which is not important for a safe shutdown of the plant to determine if they could beconie initiating event contributors given that there was an earthquake. Tnese two types of equipment ar'e not subjected to the same rigorous seismic qualification
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r 3-16 standaras as otner seismically qualified components. Depending on the capacities of these non-safety components, an earthquake with low ground accelerations may cause a reactor trip without failing the switcnyard ceramic insulators. Such an event will initiate a transient which should be evaluated by event trees similar to those presented in Figures 3.1.4 and 3.1.5. The difference between the event trees is that tnere is offsite power in this case. In the event that a transient does not occur given an earthquake, then the sequence is a success event.
An example to illustrate these points is the feedwater system. It is not a system that is required for a safe snutdown of the plant nor is it a safety-related equipment, nowever, if an earthquake occurs control, relays and other components of the feedwater system may generate a trip of the system rnsulting in an a reactor transient.
In Figure 3.1.3, the event T , ssequence number 1, was treated as an OK sequence. A note at the bottom of the figure states that a seismic event that does not lead to the loss of offsite power is considered to be benign and is '
I adequately accunted for in the turbine trip initiating event.
Given that there is an earthquake, if offsite power is still available, the event tree presented in Figure 3.1.3 does not model the plant response beyond that point. In principle according to the event tree, the reactor is not even scrammed and therefore, there is no need for it to be transferred to tne turbine trip event tree. However, if there is failure of non-safety equipment or tripping of the eguipment offline which results in a plant transient, such as the loss of feedwater due to seismic, then the event tree snould be further developed to define the accident sequences. The internal event turbine-trip event tree is not appropriate since the mitigation system considered will not include the necessary seismic failures. The new event tree wil1 be similar to the one which is in Figure 3.1.4 with certain random j failure values modified to reflect the availability of offsite power. BNL estimated that by transferring Ts to this new event tree, the only sequence wnica may contribute to the overall core damage would be TsuX. The core damage probability is estimated to be in.the order of 10-7 to 10-8, D
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= > * ~ ~ " ' " " * " * * * ' ' '-'
, 3-17 In the event that given the reactor transient, there is a failure to scram, an event tree similar to Figure 3.1.5 should be developed. It is conceivable tnat tne contribution to risk due to class V sequences may not be negligible. It is recommended that these considerations of additional acci-dent sequences sho'uld be addressed in the LGS-SARA.
Not Event Quantification LGS-SARA stated that non-failure states are included in tne Boolean expression of the accident sequences and therefore in the quantification process. BNL performed some preliminary estimates of the core damage probability for tne six dominant sequences as identified in Table 3.1.4 and the results are also provided in the table. The values under the LGS-SARA column come directly from Table 3.2 of the LGS-SARA report, whereas the BNL
~
estima~tes reflect two assumptions made to assess the core damage probability of the respective sequences based en the Boolean expression given on Table C-3 to C-5 in the LGS-SARA. One of these estimates, (the second column), did not, include the NOT events that appeared in the Figure 3.1.3. They are namely the NOT of RB , Cg anc RPV. Good agreement was obtained between the LGS-SARA results and the second column BNL preliminary estimate values. The final column delineates BNL results when the NOT events were included. These estimates show a major reduction for two sequences: an order of magnitude change for the sequence T3S E UX and a factor of 2 decrease for sequence T3 B R. As far as the remaining sequences are concerned, _little impact is observed.
Attempts were made to identify the cause of this apparent difference in results. A number of possibilities have been identified; 1) the preliminary
! estimates performed by BNL was not accurate; the agreement between LGS-SARA values and the BNL estimates without NOT events (second column on Table 3.1.4)
{ is fortuitous, 2) the NOT events were not properly included in the PeCo quan-tification, 3) the computer code SEISMIC did not perform these quantifications l properly. If indeed the NOT events were not included in the peCo quantifica-
! tion, then BNL results (last column on Table 3.1.3) indicates a substantial I
impact reduction for the T E33 UX and T RS B sequences. Furthermore, the w'
. . _ ~ , . . .
3-18 core damage frequency would be more evenly distributed over tne six sequences as indicated in the last column.
ADS Seismic Failure A discussion of how LGS-SARA modeled the ADS in the seismic system fault tree is given in Section 3.1.1. It :s inferred tnat the failure of tne ADS hardware is included in the definition of X wnich is tMe manual depressuriza-tion function, X=A+X- R XR represents t1e random failure of the manual depressurization function; A comprises seven different types of electric failures. They include the loss of the 440 V power supply, the 4KV supply, the diesel generators and the DC I power. LGS-SARA conservatively assumed that the failure of all these events wculd lead to a failure of the ADS hardware. In essence, only the failure of the DC power supply would lead directly to an AOS failure. It is, of course, obvious that the availability of AC power provides added assurance of the reliability of the DC power supply; however, failure of the 440V bus' does not result in failure of the AOS. It is for this reason that LGS-SARA is conser-vative when it assumed that X = A + XR*
Since NOT events are considered to be important in sequence quantifica-tion, they should be included in the sequence evaluation. However, as a re-sult of a conservative definition of X, this may lead to non-conservatism in other sequences. This effect may not necessarily manifest itself in the 1
change of the core damage frequency, but it may well havi substantial impacts on the risk evaluation. -
For instance, if accident sequence T S3 E UX, sequence %. 6 in Figure 3.1.4, is calculated by assuming _ either the 440V, the 4KV, the diesels or the DC power will fail the function, then a NOT-X event will imply that these various types of power supplies are available. This represents a non-con-servative departure from the system modeling, since the operation of ADS can only imply that DC power is available. This will tend to underestimate sequences TSS E U, T 33 E UW and 7 33 E UV. The impact may reside in underestimating the contribution to accider.t class IS, whereas the change in
,.f. - ==a ~ * " " * * - ~ ~' "~
l 3-19 core damage probability may be inconsequential. Other risk indices, such as, -
latent and acute fatalities may be affected differently. .
One of tne approaches to address this concern is to integrate the ADS hardware with the low pressure injection function, V consistent with tne LGS-PRA definitions.
Secuence Ouantification Tne focus of this discussion will be primarily on the six dominant sequences identified by LGS-SARA and on other sequences which BNL believes will reflect some impact on the risk indices.
(I) Dominant Sequences ,
If the modifications in Section 3.1.1.2 and this section are included in the sequence quintification, only three of the six dominant sequences are significantly affected. Table 3.1.5 enumerates the changes in core damage frequency given a modification in system unavailability for each of the dominant accident sequences. The core damage frequencies tabulated on Table 3.1.5 are preliminary estimates only. The first column identifies the six dominant accident sequences. Two of them are ATWS events: TECC 3332 and TSgg R C . The_value in parenthesis following each sequence name is ,
the core damage frequency as calculated in LGS-SARA. The second column depicts the system which is modified when the sequence is re-quantified. The value in paranthesis denotes the revised system unavailability. The last column is the core damage frequency as a result of the requantification. The sequence T33 E UX is calculated in LGS-SARA to have a core damage frequency of 3.1 x 10-6 and if the manual depressurization function random failure is modified to the new BNL value of 6.0 x 10-2, BNL estimated that the core damage will increase to about 4.0 x 10-6 It is assumed in the calculation that bes,ide X, all other components retain their values as suggested in LGS-SARA. Similarly, if only the U function is modified, an increase from 3.1
, to 3.8 x 10-6 is observed. Increases in the failure to transfer from the CST to the suppression pool produce similar results - 3.8 x 10-6 If all of these modifications are integrated into the accident sequence T 33 E UX, the s.j i
_ g w % ,,- . . , . . . . . - _ - . u ~ '
3-20 total core damage frequency is about 5.2 x 10-6; approximately a factor of 1.7 increase.
If one assumes that both the HINIA and the RIN3 become unity and the other system values are those of the LGS-SARA, then the core camage frequency for tne accident sequence T E 3SUX becomes 4.0 x 10-6 In other words, in the event tnat there is a total failure to transfer from tne CST to the suppression pool, for causes wnich may be human dependence failure or failure of all CST level sensors, the core damage frequency increases by a factor of about 1.3.
The other two sequences that are affected are the ATWS sequences. BNL revised the definition of Cg to reflect a more prudent approach in view of the lack of information in LGS-SARA on the definition of mechanical failure to scram. The BNL definition ~is given as follows:
i l Cg=CR+33+35+3' 7 This definition leads to an increase of about a factor of 5 for both the TECCS 3 g 2 and the T R CS g 3 accident sequences.
There is no impact for tne remaining three cominant sequences as a re-suit of the modifications in Table 3.1.2. The tctal core damage frequency is increased by slightly less than a factor of 2. This increase does not include the contribution from considering the NOT events.
! (II) T3S E UV. Accident Sequence The core damage frequency of the accident sequence T33 E UV is calculated in LGS-SARA to be 5.9 x 10-9 The Boolean expression of this se-quence can be written as follows:
TSSE UV = T 3 Ng S E guTV C~
= T3641M T I 3 E TUV.
If one uses the definitions of V and X provided in LGS-SARA, the following expression will. result:
T3S E UV = T364M T I I I ISR3 31 2 17 + I E Y R H R1Rgg
+ TR IS1317HR (3.1) -
R R + IRIS 12R 3VG
, , , _**' $%_n d. "?$? "
~* *
- y __
^
~
- .2 --
3-21 where Sj , i = 1,2,..17 are the seismic-induced component failures; a de-tailed listing is given in Table 3.1.1. The bar above eacn variable denotes a NOT event. 3C is the mechanical failure to scram; A is seismic failure of the electric power system; the subscript R denotes random failures. H and R represents the HPCI and the RCIC system respectively. G denotes tne combina-tion of transfer and high pressure system failures, and is defined as follows:
G = HINIA
- RIN3 + HINIA
- Rg + RIN3
. LPCI
- LPCS + AUS, waere LPCS and LPCI are tne same as those defined in LGS-SARA, and wnere the added term ADS is the sum of the ADS nardware random failure Ag and the electric power A, the following Boolean expression is ob-tained:
T33 E UV = T T 5 ERI 364M I S A + T RS 132517
+ISRHYR I g R R + R1R I S R HgSt7 .
+TSRHARIRRR+TS3R12 G Vg
+IS3R12 GAR VR represents tne random failure of the LPCS and LPCI systems. Comparison of the two expressions in Equations 3.1 and 3.2 indicates that except for NOT-A, Equation 3.2 contains all the terms of Equation 3.1 and, in addition, there are three more terms which are not in Equation 3.1. These terms contain a failure of the electric power system and failure of the ADS hardware given the loss of high pressure injection. BNL did not estimate the contribution of this sequence as a result of the modifications made. It is suggested a more detailed analysis be provided in LGS-SARA to better identify the contribution of T33 E UV to core damage and to the final risk.
! (III) Other'ATWS Sequences
, BNL reviewed the LGS-SARA ATWS event tree and found that, beside those two dominant ATWS sequences, T E C C3 S g 2 and Sgg T R C , the contribu-tion to core damage from other ATWS sequences defined in Figure 3.1.5 is relatively small. However, with the BNL definition of C g , there will be about a f.ctor of 5 increase for all the ATWS sequences in Figure 3.1.5. The l
_n __
3-22 1
total ATWS core-damage frequency reported in LGS-SARA is 8.1 x 10-7; by.
eliminating the PCR and using the BNL Cg definition, the total ATWS core damage becomes approximately 4.0 x 10-6 This does indicate that the ATWS results are quite sensitive to the parameters used to define the failure of scram. PECo believed that, based on.the failure modes defined for the reactor internals and tne CRD guide tubes, it will be conservative if one assumes that failure of these components woula cause directly a failure to scram. BNL tends to agree that there may be conservatism inherent in the definition of failure modes of these components and would encourage additional analysis be provided to support the LGS-SARA assumptions. A refined analysis in this area is needed since it will have significant impact on the acute and latent fatalities, j Examination of ATWS function unavailabilities provided in Table 3.1.3, reveals a number of major increases in the random failure probabilities: a factor of about 1.6 for the HPCI; a factor of approximately 6.6 for the RCIC; .
a factor of 25 for the ADS inhibit function; and a factor of 1.4 for the W 2
functions. BNL did not perform any re-assessment of those accident sequences whicn are affected by these modifications, however, due to the change in magnitudes of some of these functions, and the fact that significant contri-bution to risks comes from the Class IV events, it. will be prudent to evaluate the effects of these cnanges upon the results on core damage as well as the final risks. Sensitivity analysis would also provide neipful insight in the .
evaluation of these accident scenarios.
(IV) Summary BNL did not reassess the final core damage frequency as a result of all the proposed changes. A few of the areas identified requires more detailed analysis whereas others need additional infonnation to substantiate the as-sumptions. Re-quantification of some changes were made 'wnerever it was possible and results are discussed earlier in this section. It appears that i for these modifications investigated, at most a factor of two changes to the core damage frequency is observed. In view of the large uncertainty associ-ated with the seismic accident sequences, tnese changes in magnitude do not s constitute any significant impact on the core damage frequency, but their effects on the acute and latent fatalities may be significant.
' ~
.. ,e . ~ I. ,,;... p_ * - ~
~~' '~ ' " " ~
3-23 Table 3.1.1 Significant Earthquake-induced Failures Yeaian grouna Failure cause acceleration No. Component or mode capacity S g S g
9 St Offsite power (500/230-kV Ceramic insulator 0.20 0.20 0.25 switchyard) breakage S2 Condensate storage tank Tank-wall rupture 0.24 0.23 0.31 S3 Reactor internals Loss of shroud support 0.67 0.28 0.32 S4 Reactor enclosure and control structure Shear-wall collapse 1.05 0.31 0.25 SS CR0 guide tube Excess bending 1.37 0.28 0.35 S6 Reactor pressure vessel Loss of upper support 1.25 0.28 0.22 bracket S7 Hydraulic control unit Loss of function 1.24 0.36 0.52 S8 SLC test tank Loss of support 0.71 0.27 0.37 Sg Nitrogen accumulator (SLC) Anchor-bolt shearing 0.80 0.27 0.20 S10 SLC tank Wall buckle 1.33 0.27 0.19 S11 440-V bus /SG breakers Power circuit 1.46 0.38 0.44 S12 440-V bus transformer Loss of function 1.49 0.36 0.43 breaker S13 125/250-V de bus Loss of function 1.49 0.36 0.43 S14 4-kV bus /SG Breaker trip 1.49 0.36 0.43 SIS Diesel-generator circuit Loss of function 1.56 0.32 0.41 S16 Diesel-generator heat and Stnactural 1.55 0.28 0.43 vent S17 RHR heat exchangers Loss of lower support 1.09 0.32 0.34 (anchor bolts) s.
i _ _ . . . _ . _ . ~ .
- I- _ _.
3-24 Table 3.1.2 Mean Values for Random System or Function Failures used in Transient Events LGS-SARA LGS-PRA NUREG/CR-3028 BNL
. . (this review)
HPCI 8.8X10-2 0.07 0.1157 0.1157 RCIC 7.6x10-2 0.07 0.07 0.07 V 1.0x10-4*** 2.7x10-4 3.7x10-4* 3.7x10-4*
W 2.6x10-4 - --
2.6x10-4**
HINIA 1.0x10-2 1,0xio-2 1.0x10-2 1,oxio-2 RIN3 7.0x10 3 7.0x10-3 .
7.0x10-3 1.0 OGC 1.25x10-3 1.08x10-3 1.88x10-3 1.25x10-3 X 2.0x10-3 2.0x10-3 6.0x10-3 6.0x10-2 With ADS hardwar,e and no offsite power.
- With no offsite power and only RHR.
HPCI - High Pressure Coolant Injection RCIC - Reactor Core Isolation Cooling V - Low Pressure Injection Function W - Containment Heat Removal Function HINIA - Failure to Transfer from CST to Suppression Pool in HPCI RIN3 - Failure to Transfer from CST to Suppression Pool in RCIC i
DGC - Diesel Generator Conson Mode Failure l X - Manual Depressurization.
i l
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- p e ee e g -M* *
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-=.:_
. l l
l 3-25 Table 3.1.3 Ail 4S Mean Random Failure Values SARA LGS-PRA NUREG/CR-3028 BNL (this review)
HR 8.8x10-2 0.1 0.14 0.14 i
Rg 7.6x10-2 0.5 0.5 0.5 D 8.0x10-3 2.0x10-4 2x10-2 2x10-1 Ug 2.0x10-3 20x10-4 2.0x10-4 2.0x10-3 C12 1.6x10-2 1.5x10-3 1.4x10-2 1,4x10-2 Cg 1.5x10-5 1.0x10-5 1.0x10-5 1.0x10-2 HINIA 1.0x10-2 1.0x10-2 1,0x10-2 1.0x10-2 RIN3 7.0x10-3 7.0x10-3 7.0x10-3 1,0 W2 0.1 0.14 0.14 0.14
- 0. 2 PCR -- --
1.0
'0GC 1.25x10-3 1.08x10-3 1.88x10-3 1.25x10-3 HR HPCI random failure RR RCIC random failure D ADS inhibit failure UH Failure to control reactor vessel level 8 C12 Failure of two or three SLC pumps Cg Scram failure-mechanical W2 Failure of both RHR PCg Fraction of events that lead to scram failure *
- Definitiion not given in LGS-SARA inferred from modularized system fault tree.
\
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3-26 Table 3.1.4 Dominant . Seismic Core Damage Sequences Sequence BNL Estimates Class LGS-SARA W/o NOT Event w/NOT Event TS3E 0X I 3.1x10-6 4.0x10-6 4xio-7 T3Rg IS 9.6x10-7 9.5x10-7 5x10-7
( TSRPV . S/III 8.0x10-7 4.4x10-7 4.4x10-7 TECCS3g2 III/IV 5.4x10-7 6.0x10-7 4.5x10-7 TSBR Cg IS 1.4x10-7 3.5x10-7 3.5x10-7 TEW 33 II/IS 1.1x10-7 1.1x10-7 1.1x10-7 Total 5.7x10-6 6.5x10-6 2.3x10-6 Table 3.1.5 Dominant Seismic Sequences with BNL Changes Sequence (Core System Modified Damage Probability) Core Damage *
(Unavailability) Frecuency f
TSS E UX (3.1x10-6) X (6x10-2) 4.0x10-6 U (8.1x10-3) 3.8x10-6 HINIA, RIN3 (1x10-2,1.0) 3.8x10-6 AlI cambined 5.2x10-6 T3SE CgC2 (5.4x10-7) CM 3.0x10-6 -
TRCS S g (1.4x10-7) .CM 1.8x10-6 TSPRV (8.0x10-7) --
8.0x10-7 TSSE W (1.1x10-7) -
1.1x10-7 TRS B (9.6x10-7) 9.6x10-7
, Total 5.7x10-6 1.2x10-5
- Based on LGS-SARA sequence values.
- Sum total of T 33 E UX (cambined) and the other 5 sequences .
i
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l 3-27 i
j FAE.URE TO PROVEE ACECUATE FLCU FRCN HPC:
I wm l 4 1 l 1 l RANCCM MIXTURE CF SC3MI:
DCEPC40ENT RANO@ AM SCSMIC FATLURES FA Lungs g33U,C3 I II Il I i i FAILURE T0 sCsMI:
SUPPLY WATER FRan SUPPRESSION PCCL RUPTURE I II r i I
C I
o LOSS of FAILURE 70 SUPPRESSION POOL l ouE TO SCISMIC TRANSFER ON
'AILuRC CV RHR N/t VATER Loss i I i w I Figure 3.1.1 Reduced fault tree for the HPCI system, t.
[ N. .
- --e- w- ---.= .== .mm m - . e . , . . , ,, , , , ,
- - ~ . . . .~ . . . _ _ _ _
l 3-28 FAILURE TO SCRAM i I CRAM ST*;NAL NO SC5in PRESC4T SEFCRt seNAL 9tESENT scrut Mtot. *EFCRE SCRArt 3 1 01. srAIt.
RANOCM . LURE SCSMIC CF $ AM ,
, STS M FAILURES I '
l l l C 3
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- FAILURE CY FAZLuat Cr DTERNALS CRO GUIDE MYORAULIC CSMIOUD SUPPCRT), TU8t3 c:yeTROL UNIT I II I l i i
i Figure 3.1.2 Reducec 'ault tree for failure to scram.
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3-32 3.2 Fire The objectives of this section are to give a brief presentation of the LGS-SARA approach to quantification of the accident sequences generated as a consequence of fires in the different critical zones along with the cor-responding results, to describe the SN modifications to tne quantifica-tion, and to present the revised results. This section is organized as fol-lows.
- Section 3.2.1 summarizes the LGS-SARA approach to quantification of ac-cident sequences and presents the mean values for the frequency of core-damage for the different fire zones. Section 3.2.2 presents the detailed BNL review of the different fire types for two fire zones
- Fire zone 2, for which the fire growth event tree is similar in structure to all other fire zones with 4
the exception of the second fire zone described here; i.e., fire zone 25. In this section the fire growth event trees for all other fire zones are also presented, but the details are given in Appendix A. In Section 3.2.3 a sum-c mary of the review results is presented.
3.2.1 Overview of the LGS-SARA Accident Sequence Quantification For each critical zone the LGS-SARA (1) report identified the following steps used in the quantification of accident sequences:
- 1. Identification of potential initiating fires within the fire zone; the j following types of fire were considered: -
- a. self-ignited cable raceway fires,
- b. self-ignited fires in power distribution panels, and
- c. transient combustible fires.
- 2. Evaluation of the frequency of each of the above types of fires within the fire zone.
-, 3.
Subdivision of the growth of fires into several intennediate stages between ignition and damage to all safe shutdown systems served by ca-bling or components located within the fire zone.
. .s e
. ., _ A. . m eo% w - %.mw< ' ' ~ " '
. _. 1 i*, *,
r 3-33 F
- 4. Evaluation of each fire growth stage in terms of (a) the probability of failing to suppress the fire before reacning each stage, and (b) the shutdown systems tnat remain undamaged at each stage.
- 5. Evaluatior, of the conditional probability of core melt at each stage of fire growth, taking credit only for the reliability of systems not already camaged by the fire. This was achieved by mcdifying tne fault and event trees developed in the LGS-pRA(2).
- 6. Evaluation of tne core damage frequency associated with individual
~
fire growen stages by combining the frequency of failing to suppress the fire at eacn stage of growth and the associated probabilities of core damage from random failures of tne undamaged systems.
- 7. Sununation of the core damage frequencies associated with each damage stage for all types of fires to obtain the overall fire-induced care
. damage frequency for the fire zone.
Based on the above described steps a fire-induced core melt frequency of 2.3x10-5/yr was obtained in the LGS-SARA report; the breakdown of the con-tribution of the different fire types for each fire zone is given in Table 3.2.1 (LGS-SARA Table 4.6, modified to correct some typographical errors).
l .
3.2.2 BNL Revisions in Quantification of Accident Sequences The BNL review of the LGS accident quantification considered each of the steps identified in Section 3.2.1. Review of steps 1 through 4 is described in detail in Section 2.2,'and the ma'in disagreements found in this review are sununarized below.
- a. A reduction factor of five in the frequency of self-ignited cable raceway fires was used in the LGS-SARA report. As described in Sub-section 2.2.2.1.1, the BNL review indicates that a reduction factor of three is more appropriate, based on the existing data base.
i
- b. It is the BNL judgement that the probability of' fire suppression suc-l cess is overestimated in the LGS-SARA report. Based on the discus-sions in Section 2.2.2.3, the following probabilities of failure to l . extinguish a fire in t minutes, P(t), are used in the BNL review:
L i
._ . . . . N . . ... . . . . - . . . . . . . . ..
- i 3-34 P(10)=0.43 P(30) = 0.195 P(60) = 0.08 The following values were used in the LGS-SARA report: 0.40, 0.15, and 0.04, respectively.
Review of step 5, evaluation of conditional probability of core melt at each stage of fire growth, is based on the SNL review of the LGS-PRA(7)
(NUREG/CR-3028). It 13 noted that a computer reevaluation of systen unavailacility or core camage fault trees was not mace; only nand calculations were perfarmec.
The approacn used in the reevaluation of steps 6 and 7 is essentially the same as used in the LGS-SARA report; the results of the review of steps 1 through 5 are used in the BNL review.
In the following sections, a detailed review of accident sequences for fire zones 2 and 25 is described, along with the respective fire growth event trees for the other zones. In this review, the following will be presented for each fire type: frequency of fire, fire-induced transient, undamaged mitigating systems, and dominant sequences for each fire growth stage.
3.2.2.1 Fire Zone 2: 13kV Switchgear Room
- a. Quantification of Fire-Growth Event Tree for Self-Ignited Cable-Raceway Fires. ~
The fire growth event tree for fire zone 2 is shown in Figure 3.2.1, and the evaluation of the branch point probabilities is discussed below.
Event A: Frequency of Cable-Raceway Fires The frequency of cable-raceway fires is computed by multiplying two quantities: (1) the ratio between the weight of cable insulation in this zone (8736 pounds) and the total weight of cable in the reactor enclosure and control structure (172,799 pounds) and (2) the frequency of cable fires per reactor year:
, , , ******?O***
I ,
3-35 8,736(lb) 172,799(ib)x3 (5.3x10-3 ) = 8.9x10-6/yr, where the frequency of cable fires per reactor year is 5.3x10-3 and, the reduction factor of 3 is based on the BNL analysis of the data base as discussed above.
Event B: Undamaged Systems Mitigate Acciaent Given Fire-Growth Stage 1 -
Since most of the cabling in this fire zone is associated with balance-of-plant (BCP) equipment, loss of the power conversion system for inventory makeup and long-term neat removal was assumed. At this stage all safety-related equipment is undamaged, and the dominant accident sequences and their conditional probabilities, based on the BNL review of the LGS-pRA (NUREG/CR-3028), are as follows:
. Class I QUX = 4.9x10-5 QUV = 1.5x10-6
. Class II QW = 9.4x10-6 Total (Event B) = 6.0x10-5 Event C: Fire Suppressed Before Damaging Unprotected Raceways It is considered unlikely that a cable-tray fire would be suppressed before damaging cables in conduits that are not protected by a ceramic-fiber blanket. A failure probability of 1.0 is assigned to this event (the same as
, in the LGS-SARA report).
Event D: Undamaged Systems Mitigate Accident Given Fire-Growth Stage 2 Thi.s stage represents damage to all safety-related equipment except that i
associated with shutdown methods A and B (Table 4.1 of LGS-SARA), which are served by protected cabling. The dominant accident sequences and their conditional probabilities are as follows:
i ep '
f
,, e,,,,, .,
u-..++wn +w -
__ ._ .. b.- - -
---o o 3-36
. Class I QUX = 4.9x10-5 QUV = 6.6x10-5
. Class II QW = 4.5x10-3 pQW = 4.5x10-5 ,
. Total 4.7x10-3 Event E: Fire Suppressed Before Damaging Protected Raceways This event is concerned witn the probability of failing to suppress this fire before protected cables serving snutdown metnods A and 8 are damaged.
This is equivalent to failure to suppress the fire witnin one nour after the fi re. This probability is equal to 8.0x10-2, using the BNL curve given in Section 2.2.2.3;'tne LGS-SARA uses a value of 0.04.
Event F: Undamaged Systems Mitigate Accident Given Fire-Growth Stage 3 Fire growth stage 3 represents damage to all safe-snutdown systems served by the equipment in the fire zone. From the description of this zone it is
- clear that such damage would result in a loss of all systems required for safe shutdown and the resulting conditional probability of core molt is thus 1.0.
D. Quantification of the Fire-Growth Event Tree for Equipment-Panel Fires.
The fire-growth event tree for panel fires is also shown in Figure 3.2.1 and the evaluation of the branch probabilities follows.
Event A: Frequency of Panel Fires The BNL review agrees with the frequency of panel fires as calculated in LGS-SARA, i .e. , 1.8x10-3/yr.
Event B': Undamageo Systems Mitigate Accident Given Fire Growth Stage 1 Since the panels in this zone serve BOP equipment, the initiating event is loss of the power conversion system and the quantification of this event is identical with that described for Event B in Section a.
- Na"*
. - _ . - - * * .e-w
.- s_ .. ,-* '
.*M**-*-* ' *
- l 1
l 3-37 '
Event-C: Fire Suppressed Before Damaging Unprotected' Raceways The probability for fire propagation out of a distribution panel was considered to be equal to 0.04 in the LGS-SARA report. In Section 2.2.2.3 of this report there are some qual,itative comments about how this value was obtained. However, the BNL review does not change this value.
Events 0, E, and F Given that a fire has propagated from the panel in which it originated to adjacent cable raceways, the quantification of the conditional probabilities associated witn events 0, E, and F is identical witn chat described in Section a.
- c. Quantification of the Fire-Growth Event Tree for Transient-Comoustible Fires.
The fire growth event tree for transient-combustible fires is also presented in Figure 3.2.1, and the evaluation of the branch probabilities ,
follows.
Event A The BNL review concludes that the frequency of transient-combustible fires -
given in the LGS-SARA report seems to be reasonable; this probability is equal to 1.3x10-5/yr.
Events B, C, 0, E, and F The evaluation of the canditional probabilities associated with Events B, C, D, E, and F is identical with that described in Section a.
3.2.2.2 Fire Zone 25i Auxiliary Equipment Room
, a. Self-Ignited Cable Fires.
The frequency of self ignited cable fires in the raceways of the auxiliary equipment room.was determined in the same way as described for Event A in Section 3.2.2.1.a. This frequency is given by:
l
"*.* * = ' * * -
h, 4,, %+ w *W 6 ewh elme me.
m== 4
- -'y--= 'm'aew-e e'
- -- ,.7
\
r 3-3'8 -
4,400(1d) 172,739(1b) 3x (5.3x10-3 ) = 4.5x10-5fyr In the LGS-SARA report it is argued that based on fire analysis of raised floor sections, a fire initiated in one section will neither propagate through installed combustible material (cable insulation), nor cause any damage to cabling in adjacent floor sections. Thus, the max 1' mum fire damage that coulo
~
result is the loss of one division of safe-shutdown equipment, and assuming the most demanding transient, MSIV cicsure, th'e dominant accident sequences and their conditinal probabilities are:
. Class I QUX = 5.6x10-4 ^
QUV = 9.2x10-5
. Class II '
QW = 7.8x10-7
. Total = 6.5x10-4 ,
Using these conditional probabilities, the resulting frequency of core.
melt is: (4.5x10-5) x (6.5x10-4) = 2.9x10-8fyr,
- b. Self-Ignited Cable Fires.
The frequency of cabinet fires in the auxiliary equipment room is estimated as 1.75.x10-4/cabi net-year,. This auxiliary equipment room has four cabinets where fires may cause sign,1ficant d'amage to safe-shutdown systems.
Assuming a fire in any of those cabinets would, destroy the contents of that cabinet, the following equipment would'st111 remain undamaged:
~ _
- 2. Means of Reactor Depressurization
- 3. ,The LPCI System (Two Trains) ~
- 4. The Core Spray System (0ne Train)
- 5. The. RHR System -
J s
j
/ # \ #
..i.__.2.,_ ..----...~.--A-.--. - - - - - - - + - + *
-w---- -
.g,. '
, n
, - e.r
- 'f * -
y N - *8 -
3-39 I
~
Assuming the initiating event is a transient with isolation frcm the power conversion system (LGS-SARA assumption), the following are the dominant
' accident sequences with their conditional core' celt prooabilities:
. . ' ' Class I ,
CUX = 5.6x10-4
, QUV.= 2.9x10-0
. Total <
5.9x10-4 b' 'The core melt frequency resulting from self-ignited panel fires is tnerefdre; 4 x (1.75x10-4) x (5.9x10-4) = 4.1x10-7/yr.
$r c. .Transtent-Combustible Fires.
The frequency of ' transients-combustible fires were estimated as follows:
Trash-can Fire ' = 3.4x10 4/yr Solvent-can Fire = 3.4x10-4/yr Oil Fires = 3.4x10-5fyr i
i Heat transfer analysis was used to evaluate cable temperatures resulting
~ '
'from external-expcsure fires, and based on this analysis, locations within the fire zone'wnere fires may be significant contributors to core melt were
, identified, and the area associated with each location is given in Table 3.2.2
-(Table 4t .4 of LGS-SARA). Using the results in Table 3.2.2 and the concept of critical location probability (the ratio of the area of the fire location and the total free area of'the auxiliary ekuipment room associated with Unit 1, excluding the arsa taken up by. cabinets), the core melt frequency is calcu-lated and given in Table 3.2.3. It snould 'be pointed out that the dominant sequences for each fire loca' tion are QUX and[QUV.
I l
3.2.2.3 Fire Growtn Event Trees for Fire Zones 20, 22,-24, 44, 45, and 47 The detailed description .of each event in the fire-growth event trees for zones 20, 22, '24, 44, 45, and 47, as well as their branch probability is given in Appendix A. In the following section, the review results for core damge frequency are presented.-
4F g
, v. es t-e f
f %
& ,. .. '.~' r l ': ..}.l. ,-
~ - %p.-y - --- --- --
=. .w
340 3.2.3 Review Resul'ts The core damage frequency for each fire zone and for each type of fire as obtained in this review is presented in Table 3.2.4. The most 'important re-suits are: .
- 1. The total core damage frequency res,ulting from fire-induced transients ootained in the BNL review-.is 5.2x10-5/yr, as compared to 2.3x10-5/yr reported in tne LGS-SARA report.
- 2. The difference between the BNL review and the LGS-SARA core damage frequency can be attributed to two factors: (a) the probabiitty of fire suppression in any given time, and (b) the reduction factor used in the calculation of self-ignited cable-raceway fires (see Section 2.2).
- 3. Most of the core damage frequer.cy comes from the fire growth stage 3 (about 85% the both BNL review, and about 31% in LGS-SARA). At this, fire growth stage, in almost all zones,all safe-snutdown systems are assumed to be damaged by the fire. Thus, the core damage frequency is determined by the initiator frequency and the probability of failing to suppress the fire within a given time interval. This indicates that the changes made by BNL in the accident sequence quantification (relative to the i.GS-PRA quantification) have a small impact upon the total fire-induced core damage frequency.
l 4. In the BNL review, about 67% of the total core damage frequency comes
~
from the self-ignited cable-raceway fires (about 57% in LGS-SARA).
- 5. In the BNL review, about 93% of the total core damage frequency comes from fire zones 2, 44, 45, and 47 (about 91% in LGS-SARA).
- 6. In the BNL review, about 97% of core damage is binned in the Class 1
' category (see LGS-PRA); the other 3% is Class II.
The results presented in Items 1, 3 and 4 show that the total core damage frequency is very~ dependent upon the modification made by BNL (Item 2 above).
-)
~ _ . -
. = . . . -- -= - --
3-41 Thus, calculations were performed to show the impact of these two modifica-tions, and the results are as follows:
- a. If the LGS-SARA prowability of failing to suppress the fire within 60 min. (0.04) is used instead of the BNL value (0.08), the total fire-induced core damage frequency would be equal to 3.6x10-5/ reactor year. ~
- b. If the LGS SARA reduction factor (RF=5), used in tne calculation of self-ignited cable-raceway firer is used, instead of the BNL value (RF=3), tne total fire-induced core damage frequencv would be equal to 3.8x10-5/ reactor year.
Anotner area where some sensitivity study is warranted is in the evalua-tion of human errors in case of fire-induced transients. Since 97% of the total fire-induced core damage frequency is due to failure of injection, two cases were analyzed here:
- a. Operator fails to depressurize the reactor (X in the accident sequences).
The results presented in Table 3.2.4 are based upon the value of X given in the BNL review of the LGS-PRA(7); 1.e., X=6.0x10-3 If this value is incre.ased by a factor of 10, the total fire-induced core melt frequency would be equal to 7.6x10-5 (an increase of 45%).
- b. Operator fails to initiate required systems from remote shutdown panel (pertinent to fire zones 22 and 24).
The results presented in Table 3.2.4 are calculated using a value of 1.0x10-3 for this error. If a human error probability equal to 1.0x10-2 was used, the total fire-induced core damage frequency would increase to 5.6x105(anincreaseof7.7%).
w g.
l t _ - . _ . . . . . - -_ - - -- . n-, . ._,.~:~.-.
I t.
i f
f .
A B C D E F Undamaged pg ,
Fise suppsessed Undasnaged Five suppsessed U M unaged systems l taelose dasnaging syssesnt belose dunaging systems
- N C/ panel snesigase 5,np,otected masse .se psosected musisate l accident >given saceway acculent given saceways accident given h, FGSI FGSI lFaisuse gives FGS 2) FGS2 IF suse geves FGS3l FGS3 l t c-a ,,,s a .
s, .,; s. .. ,
calde TC paael -
OK .
- i i
OK
, OK 8.0-2 8.9-5/1.3-5 / 1,0/1.0/4.0-2 1.0 cu 7.1-6 1.0-6 5.8-6 1.8-3 (see note) (see '
note) i
,i 4.7-3 '
I!
CM 4.2-7 6.1-8 1.4-7 ! !1 6.0-5 '. '
l CM 5.3-9 7.8-11 1.1-7 GS = yowth sa 7.5-6 1.1-6 6.7-6 Tea.: annual cos e melt iseesency -1.5-5 i Note: Because of the evaludaion of event E. ihe prob.besity of event c f is not included in the evaluation of the seques ca issquency.
- i Figwe 3.2.1 Firarerowth event tree for fire zone 2
! b . . .
?
__m. _ - . _ _ _ . _ _ _ _
~
__ m._.. . . . . __..- ~.1 _ _ _ _ '. _ -
j 343 o
Table 3.2.i Summary of Fire-Analysis Results Annual Contribution to Core-Melt Frequencya Self-Ignitec Transient-Caole Sel f-Ignited Ccmbusticles Fire Zone Raceway Fire Pane.1 Fire Fire lutal 2 12-kV switchgear room 2.4-6b 3.2-6 5.9-7 6.2-6 20 Static inverter room 5.0-8 3.5-8 1.5-8 1.0-7 22 Cable-spreading room 6.1-8 NAc 1.9-7 2.5-7 24 Control room Negligible 1.6-7 1.0-7 2.6-7 25 Auxiliary equipment room Negligible 1.0-7 2.6-7 3.6-7 44 Safeguard access area 4.2-6 1.5-6 4.1-7 6.1-6 45 CRD hydraulic equipment area 4.7-6 1.0-6 6.6-7 6.4-6
- 47 General equipment area 1.2-6 5.0-7 1.8-7 1.9-6 1.3-5 6.5-6 2.4-6 2.2-5 Contribution from all
, other fire zones 1.0-6 l
l Total annual core-melt l frequency from fires 2.3-5
, , a Point estimates
! b 2.4-6 = 2.4 x 10-6 c Not applicable I
l
- m. - . . _ _ _ ,_ _ -._ _ _ _
. . . . . . . . . . ~ . . . . . . . . . - , .
, 344 Table 3.2.2 Critical Locations of Transient Combustible Materials in the Auxiliary Equipment Room
- Area of location (m2)
Systems assumed to be un-Solvent-Can 011 damaged and capable of RPy Fire Location Fire Fire inveritory makeup Intersection of floor areas 0 2.4 LPCI train D, means of 10U792 (a) and 10U791 depressurization Intersection of floor (b) areas 7. 7 12 LPCI train D, means of 10U791 and 100793 depressurization Floor area 100795 (c) 0.6 2.3 LPCI trains B and C, means of depressurization Floor area 10U789 (d) 0.6 2.3 LPCI trains C and D, means of depressurization
- Table 4.4 of LGS-SARA S
4
- 9 r
t
\
l ./
i
, 4,W wad;pD9meDWuulp44 h,M -.-# 9v Jd OY4. . -*TU -4 ~~
^
. ~
3-45 Table 3.2.3 Evaluation of Sequence Frequencies of 011 Fires (Transient Lombustibles)
Probability ~b Criticala of Randon:
Annuala Location Equipment Core-MeltD
, Fire Locationa Frequency Probability Failure Frequency OIL FIRE Location a 3.4-5 0.01 0.022 7.5-9 Location b 3.4-5 0.05 0.022 3.7-8 Location c 3.4-5 0.01 0.014 4.8-9 Location d 3.4-5 0.01 0.014 4.8-9 SOLVENT FIRE Location a 3.4-4 0 0 0 Location b 3. 4-4 0.03 0.022 2.2-7 Location c 3.4-4 0.003 0.014 1.4-9 Location d 3. 4-4 0.003 0.014 1.4-9 Total 2.8-7c a From LGS-SARA Table 4.5 b BNL Review c The corresponding LGS-SARA value is 2.6-7.
9 gs l
.?. ..
3-46 Table 3.2.4 Summary of Fire-Analysis Results BNL Review Annual Contribution to Core-Melt Frequencya Sel f-Ignited Transient-Cable Self-Ignited Comeustibles Fire Zone Raceway Fire Panel Fire Fire Total 2 12-kV switengear room 7.5-6 6.2-6 1.1-6 1.5-5 20 Static inverter
- i room 2.4-7 7.5-8 4.3-8 3.6-7 22 Cable-spreading room 3.7-7 nab 7.4-7 1.1-6 24 Control room nab 4.8-7 2.2-7 7.0-7 25 Auxiliary equipment room 2.9-8 4.1-7 2.8-7 7.2-7 44 Safeguard access area 1.3-5 3.3-6 7.8-7 1.7-5 45 CRD hydraulic equipment area 9.6-6 1.8-6 8.6-7 1.2-5 i
! 47 General equipment t
, area 3.9-6 1.7-7 3.7-7 4.4-6 3.5-5 1.2-5 4.4-6 5.1-5 Contribution from all other fire zones 1.0-6 Total annual core-melt frequency from fires 5.2-5 a Point estimates b Not applicable N~
. . . . ..,,._A,S~~ ~~~~ '
e 3-47 REFERENCES
- 1. Philadelphia Electric Company, " Limerick Generating Station, Severe Ac-cident Risk Assessment," April 1983.
- 2. Philadelphia Electric Company, " Limerick Generating Station, Probabilistic Risk Assessment," March 1981.
3 Kolb, D. L., et al., " Review and Evaluation of the Indian Point Probabilistic Safety Study," NUREG/CR-2934, SAN 082-2929, December 1982.
- 4. Pickard, Lowe, and Garrick, " Indian Point Probabilistic Safety Study,"
Prepared for Consolidated Ecison Company of New York, Inc., and Power Authority of tne State of New York, 1982.
- 5. Kennedy, R. P., et al ., " Subsystem Fragility," SSMRP-Phase 1, NUREG/CR-2405, UCRL-15407, February 1982.
- 6. Battel, R. E. and Campbell, D. J., " Reliability of Emergency AC Power Sys-tems at Nuclear Powar Plants," NURFG/CR-2989, ORNL/TM-8545, July 1983.
- 7. Papazoglou, I. A., et al., " Review of the Limerick Generating Station Prob'a bilistic Risk Assessment," NUREG/CR-3028, BNL-NUREG-51600, February 1983.
e 8
s
~
4.1 Seismic Hazard and Fragility Recomendations 4.1.1 Introduction Many concerns have been raised in Section 2.1 in regard to the seismic hazard and fragility analysis. Recommendations for resolving these concerns are civen in this section. These recommendations are primarily directed to PECo and are based on discussions already pre-sented in Section 2.1. Rather than reDeating the background, each recomendation is oresented and followed by the applicable subsection in Section 2.1 which can be referred to for additional information. Al so ,
recommendations are made to the NRC to perform additional review tasks to complete the review of the LG5-SARA.
Section 4.1.2 gives the recomendations for the hazard analysis and Section 4.1.3 gives the recomendations for the fragility and associated system analysis concerns.
4.1.2 Seismic Hazard The following recomendations should be addressed by PECo. The numbers in parentheses at the end of each recomendation refer to the subsection of Section 2.1 which gives background information.
- 1. The delineation of zone boundaries in the Crustal Block hypo-thesis should be reconsidered. Specifically, a redefinition of Zone 8 is recomended that is better correlated to the pattern of seismicity in the vicinity of Limerick and the geologic structure of the Triassic Basin. (Sae Section 2.1.2.3.)
- 2. The possible occurrence of large-magnitude events (i.e., =M7.0) should be considered as an alternative hypothesis on maximum magnitude for each seismogenic zone. The distribution should be selected in consideration of recommendation 4, below. (See Section 2.1.2.3.)
4-1 1
. o
- 3. The uncertainty in Richter b-values should be considered in the seismic hazard analysis. Consideration should be given to the distribution of earthquake magnitudes based on the historical
- record in each seismogenic zone and expert opinion. (See Section 2.1.2.4.)
~
- 4. Justification should be provided for the estimate of the large-magnitude (i.e., M = 6.8) events considered in the hazard analysis. Specifically, the basis for assaning that the magnitude estimated for the 1886 Charleston, South Carolina earthquake is the largest event that can occur should be pro-vided. Also, the basis for not considering uncertainty in this parameter should be justified. (See Section 2.1.2.4.)
- 5. The implication of including the Cape Ann events in the Pied-mont source zone should be addressed. Consideration should include recent work that rejects the notion of a Boston-Ottawa seismic belt and the fact that the 1982 New Brunswict Canada event is included in the Piedmont province. (See Section 2.1.2.4.) .
The following recommendation is addressed to the NRC.
l
- 1. An independent analysis should be conducted to verify the hazard analysis results. Also, an independent quantitative evaluation of the impact of comments raised in this review should be perfomed.
4.1.3 Seismic Fragility l
The following recommendations should be addressed by PEco. The 1
numbers fr. parentheses at the end of each reconmendation refer to the subsection of Section 2.1 which gives background infomation.
4-2 .
.. on .~ *m ' " " ~ ~ ~ ~
- 1. Justification for using the 1.4 duration factor to increase the capacity of structures and the 1.23 factor to shift the hazard curves from a sustained-based peak acceleration to an effective peak acceleration should be providad. Specifically, the con-cern is the region of the Decollement hazard curve at and above 0.40g effective peak ground acceleration (i.e., in the region where the average magnitude is M6.0 or larger). (See Section 2.1.3.1.)
- 2. Justification should be provided for the median duration factor.- Specifically, the median value of 1.4 and the varia-bility associated with this factor should be addressed. A median value which is magnitude dependent (as used in the LGS-SARA) should be developed; Also the uncertainty components of variability of 0.08 should be increased.
- 3. The revised median capacity value of 0.90g for the reactor enclosure and control structure should be verified. (See Section 2.1.3;3.)
- 4. The assumption that the containment building will have an
~
effective damping value of '10 percent at the acceleration levels corresponding to the failure of the reactor internals, l CRD guide tube, and reactor pressure vessel should be justi- ,
fied. Both the damping values for the individual containment components (i.e., containment wall, pedestal, lateral support, and RPV components) and the combined system damping value should be addressed. For the latter concern, either a weighted
+
model damping calculation or a time history reanalysis of the containment /NSSS model should be conducted. (See Section 2.1.3.4.) Note that this recommendation has a lower priority 4-3 '
~. 4 . .
since the mean frequency of core melt would increase by only 10 percent for this effect.
- 5. The implications of impact between the containment butiding and the reactor enclosure should be addressed for the following concerns:
- a. Failure of safety-related electrical and control equipment located in the reactor enclosure.
- b. Failure of safety-related piping which crosses between the j two buildings due to relative displacements.
In addition, it should be verified that no safety-reltted components will be damaged by spalled concrete caused by impact of the two structures. (See Section 2.1.3.5.)
Finally, it should be verified that failure of small lines attached to the safety-related piping near the junction of the two structures and anchored to the reactor enclosure will not contribute to the frequency of core nelt.
- 6. In regard to the safety-related electrical components which significantly affect the frequency of core melt including, but not limited to:
e 440-V bus /SG breakers (S11) l e
440-V bus transfomer breaker (S12) e 125/250-V de bus (Sg3)
I 4-4
__,m, ,. ga r-
e 4-KV bus /SG (St4) e Diesel-generator circuit breakers (S15) identify the number of actual components, their locations, and their characteristics relative to the generic tests at Susque-hanna which were used to derive their capacities. Justi fica-tion should be provided for the number of each component type which should be included in the Boolean equation for sequence T3S E UX. Consideration should be given to the possible effects of capacity and response dependencies which exist. (See Se: tion 2.1.3.6.)
- 7. Justification should be provided that the test results for the Susquehanna components can be directly scaled by the ratio of the design SSE values for the two plants (i.e., Limerick and Susquehanna) and used to develop capacity values for the fol-lowing Limerick components:
e Hydraulic control unit (S7) e~
440-Y bus /SG breakers (Sit) e 440-V bus transfomer breaker (S12) e 125/250-V de bus (S13) e 4-KV bus /SG (514) e Diesel-generator circuit breakers (S15) 4-5 1
l -
l
, . . _; , _ c - , .-
Consideration should be given to the location of the components in the two plants, foundation conditions, and construction similarities. It is recommended that fragility parameter values specifically calculated for each of the above components at Limerick be dev' eloped. (See Sections 2.1.3.6 and 2.1.3.7.)
- 8. The capacity parameters for the SLC test tank should be based on a component-specific analysis which includes the dynamic characteristics of the tank and the actual geometric configura-tion. The capacity of the anchor bolts should be checked and the earthquake component factors derived ba' sed on the actual response and capacity characteristics. (See Section 2.1.3.7, Component Sg .)
- 9. The similarity between the nitrogen accumulators at the Limer-ii:k and Susquehanna plants should be verified since the anal-ysis from Susqeuhanna was used as the basis for the capacity of the nitrogen accumulator at Limerick. (See Section 2.1.3.7,
- Component Sg.) -
- 10. The possible failure of the SLC tank due to tearing of the base plate flange near the anchor bolts should be checked to verify that it is not the weakest capacity. (See Section 2.1.3.7, Component 510*I
- 11. A specific analysis should be conducted for the diesel genera-tor heat and vent which is based specifically on the character-istics of this component. (See Section 2.1.3.7 Component S16 I
- 12. Verification should be provided to document that the fragility values for valves include consideration of potential leakage through the internal components bypassing the closed valve 4-6
- e..=, .-
- - ^
. 2 . - _ -- - - - .
~
barrier. Specific consideration should be given to the MSIV and the purge and vent valves. Verification should also be provided to document that seismic motions will not cause SRVs to stick open. (See Section 2.1.3.8.)
- 13. Verification should be provided to document that soil pressures on the embedded portions of the reactor enclosure walls do not reduce the capacity of these walls and thus decrease the capa-city to resist in-plane lateral loads. (See Section 2.1.3.8.)
- 14. After construction of the plant is completed, a systematic review of the plant, including walkthroughs, should be con-ducted to locate secondary components which could fail, fall, and impact primary safety-related components. Analyses of potential failures should be conducted to determine whether the secondary components are weaker than the primary components already considered. (See Section 2.1.3.8.)
- 15. The percentages of occurrences when evacuation would be affected by earthquakes should be recalculated using realistic relationships between damage to civil structures and ground acceleration. (See Section 2.1.3.8)
The following recommendations are addressed to the NRC.
- 1. A followup review should be conducted to independently verify the capacity values used for the electrical components. A coordinated task between nuclear systems and structural engin-eers should be performed since these components are major l
contributors to the mean frer,uency of core melt.
i 4-7
%p
- Ma m a f
- 2. Other significant nonelectrical components are based on generic capacities. Independent, speciffc calculattons should be perfomed for the following: components since they are important to the final risk.
- Hydraulic Control Unit (S7 )
e Nitrogen Accumulator (Sg )
e Diesel Generator Heat and Vent (S16) 1 O
l t
4-8
- m. -- ---- -.- --ww.. -- -em+-- -- -- * * - - - ' - - - * - * '~ '~ ~~ * * '
. _ _ + ~ . . . . - _ . m. .. . _ _ . - - . _ . .
4-9 4.2 FIRE The methods to evaluate the risk due to a fire in a nuclear power plant
. (NPP), as described within the Limerick SARA, and as reviewed herein, can be divided into three categories for the development of ignition, detection, suppression and propagation models: physical models, point probability models, and probabilistic models. The Limerick SARA attempts, and.in our judgement rightly so, to use a hybrid of all three. A hybrid approach is indeed war-ranted. Physical models suffer frem the complexity of the large numcer of variables and relationships required to calculate a fire history. Point probability models suffer from small and inadequate data bases. While a com-pletely probabilistic approach also suffers from data base inadequacy, it, more importantly, suffers from an inability tn accurately model certain phases of fire development.
To put the issues of fire-development modeling in proper perspective, let us consider those components of the fire which are relevant in assessing fire growth: the burning object, the flame, the hot layer, the cold layer, the vents within an enclosure, target objects (other ccmbustibles), and inert surfaces (walls and ceilings). As Friedman [13 points out rather simplist.1-cally, 20 interaction vectors involving heat and material flux exist between
$hese seven components. Several of these interactions have multiple elements with positive feedback as a critical part of the fire growth phenomena.
Adequate knowledge of the various feedback loops should suffice, in prin-ciple, to permit description of t'he growth rate of the fire. However, in order to make safety assessments, it is also mandatory to have additional in-formation such as carbon monoxide and smoke content, for its impact on plant personnel safety. More importantly, from a public risk viewpoint, it is necessary to have information on the plant damage states as a function of fire growth.
Indeed, assessing fire risk .is a highly coupled, nonlinear, dynamic pro-cess. We at BNL are of the opinion that the state-of-the-art in fire modeling, coupled with such complex issues as systems interaction from automatic / manual suppression and human error, is such that probabilistic analyses which purport to quantify the safety of NPPs in the event of a fire have a wide range of uncertainty.
1
-s_-
i l
.- . . , & ~_n. nn , . -n -, . . - ~ .
4-10 Furthemore, the very conservative assumptions used in the Limerick SARA fire analysis (in most respects) may, without proper context, lead to a distortion of perspective for fire risk relative to other risks at the plant.
In some respects, assumptions and submodels that are touted to be con-servative are tantamount to gross violations in physical realities. Sevarai cases in point have been discussed in the previous sections - not linking a suppression model directly to the fire growth model; a mass-loss rate model that does not truly reflect the positive feedback of the various fire growth stages; an ignition-time model that does not adequately reflect the various heat-exchange mechanisms are some of the modeling inadequacies $1ch have been addressed directly.
The Limerick SARA on fire analysis has only considered intrazone fire propagation. A true assessment of fire risk must consider interzone fire propagation and all aspects pertaining thereto including the decilitating r
effect of smoke mig' ation. While this latter facet has no immediate bearing on component reliability, it should have immediate ramifications with regard to manual suppression effectiveness. Hence, smoke propagation is one aspect that should have been considered even if its level of sopnistication can be construed to be only on a par with the physical models used in ascertaining the thermal history.
In this connection, the mechanisms by which fire suppression systems (automatic and/or manual) can cause the failure of redundant or diverse safety systems should be considered in the assessment, again to a level of detail consensurste with the probabilistic/detenntnistic analysis that is applied to assess fire risk.
The foundation on which the fire propagation model, basically a one-room f
fire model, rests is sound. Various capartment fire models[2] have been developed and COMPBRN can be considered as one which lies within their spectrum of sophistication. COMPBRM, along with other fire models, uses a control volume, or " zone approach," in lieu of those models which discretize the governing differential equations directly, the so-called " field-model approach."
, eml _-_y_,
_y _. -. - - . - - ~
_ = _ . -. . . .
4-11 This zonal approach has several important advantages: (1) canputational simplicity, (2) ease of decoupling zones for independent investigation; (3) simpler conparison of theory and experiment for individual zones, and (4) easier conceptualization of the interaction between zones. Field models, however, in the long run should provide the most general, accurate, and detailed prediction of fire development. However, at present field models:
(1) are limited by computer capacity, (2) do not yet properly treat action-at-a-distance radiative energy transfers, and (3) are still awaiting a more rigorous treatment of buoyancy driven turbulence. Both the zone and field approach should, in BNL's judgement, be pursued with the field approach usad as a basis for " fine-tuning" the unit models that are built into the zone-model approach.
Zone models, like COMP 8RN, represent a nearer term engineering approach which is closely tied to experimental observations. However, a basic philo-sophical limitation in zone-model structure is in its emphasis in predicting roon flashover. For assessing nuclear power plant risk, predicting the onset of flashover is not as crucial as predicting the effects of in-place component vulnerability during the earlier fire-growth stages. As such, for canplet'e-ness a larger spectrum of initiating fire sizes must be incorporated into the
- I analysis.
Accordingly, several of the unit-models enpl'oyed in the zone approach re-quire improvement.[2] Other aspects of fire growth that are lacking in existing models (like COMPBRN) are needed. For direct application in assessing nuclear power plant fire risk, these additional models should reflect the possibility of (1) the effects of walls, corners, and obstacles on fire plume and thermal plume development, (2) the possibility of combustion of excess pyrolyzate within the stratified layer, (3) the effects of turbulent-induced buoyancy on plume development (4) intra zone mass and energy exchange, (5) and implementation of existing knowledge and correlation of fuel-f1muna.bility characteristics, specifically current cable flammability and j damageability indices.
T_:
4-12 -
Another keypoint regarding the practical use of a zone model in general, and COMPBRN in particular, is that the structure of the numerical code is not
" user friendly". Before one can use a code employing a series of unit models, one must have an awareness of the assumptions that are built into the analy-sis, the key physical parameters and their sensitivities, and finally a working knowledge of the state-of-the-art in fire phenomena and modeling.
e e
e 9
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e 4
e e
__ --_ _ f I_*'_ J_ --
_ 2*"~ ' *'r2 -
~ * * ~ - .
4-13
4.2 REFERENCES
~
- 1. Friedman, Raymond, " Status of Mathematical Modeling of Fires," Factory Mutual Research Corporation, FMRC RC 81-BT-5, April 1981.
- 2. Jones, Walter J., "A Review of Compartment Fire Models," NBSIR 83-2684, i April 1983.
e t
4
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p I
- er
_ _ - - ---__._-_z_-- - --__ --- - _
._ . .-~.--. - - - -.
A-1 APPENDIX A DETAILED REVIEW 0F THE QUANTIFICATION OF THE FIRE-GROWTH EVENT TREES In this appendix the' detailed review of the fire-growth event trees for the following fire zones is described:
- 1. Fire Zone 20: Static Inverter Room
- 2. Fire Zone 22: Cable-Spreading Room
- 3. Fire Zone 24: Control Room 4 Fire Zone 44: Safeguard Access Area
- 5. Fire Zone 45: CR0 Hydraulic Equipment Area
- 6. Fire Zone 47: General Equipment Area A.1 Fire Zone 20: Static Inverter Room - -
The fire growtn event-tree for all types of fires in zene 20 is shown in Figure A.I.
A.1.a Quantification of the Fire-Gcowth Event Tree for Self-Ignited Cable-Raceway Fires.
Event A: Frequency of Self-Ignited Cable-Raceway Fires.
This frequency is calculated in the same way as for Event A in Section 3.2.2.1.a. i.e.,
9,558(lb) 5.3x10-3 172,799(lb) ( "
- 3 Event 8: Undamaged Systems Mitigate Accident Given Fire-Growth Stage 1.
Based on the locality of the initial fire a reactor trip transient is
- assumed, with the lo'ss of one division of safety related equipment. The dominant seqences and their conditional probabilities are as follows:
. Class I QUX = 8.4x10-6
[ - QUV = 1.1x10-5 I ,
.O. '
I t - , . . .- - - -. _ _ . . . - . - -. -- , , , ,
A-2
. Class II PW = 4.5x10-5 QW = 5.4x10-6
. Total = 7.0x10-5 Event C: Fire Suppressed Before Damaging Unprotected Raceways.
The procability of this event is given by the probability of failing to suppress the fire within 10 min. (estimated time before damage to unprotected raceways). BNL value for this event is 0.43.
Event 0: Undamaged Systems Mitigate Accident Given Fire-Growtn Stage 2.
Fire growth Stage 2 represents damage to all safaty-related equipment except tnat associated witn shutdown method A wnich is served by cable raceways protected with ceramic-fiber fire blankets; also unaffected is equipment associated with tne power conversion system. The dominant sequences and their conditional probabilities are as follows:
. Class !
QUX = 8.4x10-6 QUV = 2.6x10-5
. Class II PW = 6.6x10-5 QW = 7.9x10-6
. Total = 1.1x10-4 Event E_: Fire Suppressed Before Damaging Protected Raceways.
The probability of this event is given by the probability of failing to suppress the fire within 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> (estimated time before damage to protected raceways). BNL value for this probability is 8.0x10-2, Event F: Undamaged Systems Mitigate Accident Given Fire-Growth Stage 3.
This stage represents damage to all safe-shutdown systems served by equipment in this zone. Only the power conversion system would remain undamaged to mitigate the accident. The dominant sequences with their ,
9
.__n_n,. ,93# ,' J- -.
9 enW '* ** '
_____.___________~______ -
A-3 conditional probabilities are:
. Class I .
QuV = 2.0x10-2
. Class II PW = 1.0x10-2
. Total = 3.0x10-2 A.1.b Quantification of the Fire-Growth Event Tree for Panel Fires.
Event A: Frequency of Panel Fires.
BNL agrees with the frequency of panel fires given in LGS-SARA, i.e.,
4.4x10-4/yr.
Event B: Undamaged Systems Mitigate Accident Given Fire-Growth Stage 1.
Based on the panels located in this zone, reactor trip transient is assumed, and the following equipment is assumed to have failed: HPCI, RHR -
Trains B and 0 and Train 8 of LPCS. The dominant accident sequencas, and tneir conditional probabilities are:
. Class I QUX = 8.4x10-6 QUV = 1.1x10-5
. Class II PW = 4.5x10-5 QW = 5.4x10-6
- l. . Total = 7.0x10-5 Event C: Fire Suppressed Before Damaging Unprotected Raceways.
The quantification of this event is identical with that for Event C in Panel Fires for Fire Zone 2 (see Section 3.2.2.1.b).
t
/
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=94-
l._ .
.T-- -
~ ~ ~~ ~^^
. .'. -- .. :. ^
A-4 Events 0, E, and F The quantificatio of the conditional probabilities associated with those events is identical with that described for self-ignited cable-raceway fires in Section A.1.a. .
A.I.c Quantification of the Fire Growth Event Tree for Transient-Combustible
.Fi res.
Event A: Frequency of Transient-Combustible Fires.
BNL agrees with the frequency of transient-ccmbustible fires as calculated in LGS-SARA, i .e. , 1.7x10-5fyr, Events B, C, 0, E, and F The evaluation of the conditional probability associated with Events 8. C, O, E, and F is identical with that described in Section A.I.a.
A.2 Fire Zone 22: Cable-Soreading Room The fire growth event tree for all types of fires in Zone 22 is shown in Figure A.2.
- A.2.a Quantification of the Fire Growth Event Tree for Self-Ignited Cable Raceway Fires.
Event A: Frequency of Self-Ignited Cable-Raceway Fires.
This frequency is calculated in the same way as for Event A in Section 3.2.2.1.a, i.e.,
~
35,526(1b) j 172,799(1b)x(5.3x10-3 3
) = 3.6'x10-4/yr.
Events B and C Since all fires are capable of damaging adjacent cable raceways, except those protected by a ceramic-fiber blanket, Event 8 is effectively omitted and l Event C is assigned a probability of 1.0.
1 im *7,",
'~ ~ '
)
A-5 .
Event 0: Undamaged Systems Mitigate Accident Given Fire Growth Stage 2.
The initiating event is a transient with isolation from the power-conversion system and tne only equipment potentially operaole is that associated with shutdown methods A and B. The dominant accident sequences and tneir conditional probaoflities are:
. Class !
QUX = 4.9x10-5 QUV = 6.6x10-5
. Class II QW = 2.7x10 4 PQW = 4.5x10-5
. Total = 4.3x10 4 Event E: Fire Suppressed Before Damaging Protected Raceways.
The protected raceways (serving shutdown methods A and B) consist of cable trays protected by a 1" thick ceramic-fiber blanket which is equiv3 lent to a 1/2 hour fire rating. Thus, Event E is assigned a probability of 1.95x10-1
'which is the probability of failing to suppress a fire witnin 1/2 hour.
Event F: Undamaged Systems Mitigate Accident Given Fire Growth Stage 3.
At this stage all safe-shutdown equipment dependent on cabling within this zone-is considered to be damaged. The only equipment that is potentially operable is that served by tne remote shutdown panel. Therefore, the dominant accident sequences and their conditional probabilities are:
. Class I QUX = 4.2x10-4 QUV = 2.2x10-3
. Class II QW = 4.0x10-4 l PQW = 6.6x10-5 .
. Total = 3.1x10-3 ,
s_
e
.e, ee-, y e -= ,=,e, em,-- =c- ,, , - . , , , , -e *
. _ .... .. - -.~. .. . - . - - - - . . . . - . - - - . . .. - ...- --- *
~
A-6 A.2.b Quantification of Fire Growth Event Tree for Transient Combustible
~
Fires.
Event A: Frequency of Transient-Combustible Fires.
BNL agrees with the frequency of transient-combustible fires presented in LGS-SARA, i .e. , 7.2x10-4/yr.
Events 8, C, 0, E, and F The quantification of all those events is identical with that discussed in the previous section (see Section A.2.a).
A.3 Fire Zcne 24: The Control Room Since there is no exposed cable insulation in the control room, the'only types of fires analyzed in this section are: Self-Ignited Panel Fires and Transient-Combustible Fires. '
A.3.a Quantification of Fire-Growth Event Tree for Self-Ignited Panel Fires. .
~
The fire-growth event tree for self-ignited panel fires is shown in Figure A.3. .
Event A: Frequency of Self-Ignited Panel Fires.
The frequency of significant panel fires in the control room was estimated to be 1.8x10-3 '
Event B: Undamaged Systems Mitigate Accident Given Fire-Growth Stage 1. ,
Thisstagerepresentsdamagethatisconfdedtothecabinetinwhichthe fire starts. There are 17 separate cabinets in the control ecom. . However,<
only fires in 3 cabinets can cause significant damage. Fires in one of thcs -
cabinets may disable all systems required for reactor shutdown with the' exception of equipment that may be controlled from the remote shutdown'paneT.
i Fires in. the other two cabinets will only disable the power-conversion system.
yd 4
.?,_
,a=
s mg
,_, 9.euemeses,fy **
Y .
'.M-
- m. - -
w - , ,
A-7 1
a , The transient resultir from any of these fires is Loss of Feedwater or
!*.SIV Closure and tne domir at accicent sequences, with their conditional probabilities are:
s
~
. Cl4ss I .
Qu'X w 3.1x10-5 '
QUV =' 1.340-4 _ _
~
. C1 ass 11 QW4 = 2. 5 x10-5 PQW = 3.9x10-6.
. Total = 1.9x10 4 Event C: Fire Suppressed ifore Propagating Beyond the Confinement of the Cabinet.
BNL agrees with the ev uation of the probability of a cabinet fire propagating beyond the cor nement of the cabinet as given in LGS-SARA, i.e.,
2.5x10-2 -
Event 0: Uncasageo Systerr Mitigate Accident Given Fire-Growth Statje 2.
In this stage it' is cc 'dered i that only the equipment which can be oper-ated'from the remote'shutc -n panel is potentially operable. The dominant ac-
~
, cident sequdnces and.their oncitional probabilities are iaentical with those
- y. calculated for Event F in' ction A.2.a. i.e., the total conditional core dam-
. age probability is egual't- 5.1x10-3 A.3.b Quantification og e Damage Probability for Transient-Combustible Fires.
BNL agrees with the qu tification of the frequency of transient-combustible fires which ca . damage safe-shutdown equipment in the control room. This frequency is e al to 7.2x10-5fyr, "Given the occurrence o- a transient-combustible fire it is assumed that the only potentially operat a equipment is that which can be operated from the remote snutdown panel. In ais case the dominant accident sequences and their m .
i, ryo "w
_he d'
'4 il.
A-8 conditional probabilities are identical with those calculated for Event F in ~
Section A.2.a; 1.e., the total conditional core damage frequency is equal to 3.1x10-3 So, the total contrioution of transient-combustible fires to the core amage frequency is given by:
(7.2x10-5) x (3.1x10-3) = 2.2x10-7/yr.
A.4 Fire Zone 44: Safeguard Access Area The fire-growth event tree for all types of fires in Zone 44 is shown in Figure A.4.
A.4.a Quantification of the Fire-Growth Event. Tree for Self-Ignited Cable Raceway Fires.
Event A: Frequency of Self-Ignited Cable-Raceway Fires.
This frequency is calculated in the same wa as for Event A in Section 3.2.2.1.1, i.e.,
28,290(lb) 5.3x10-3 172,799(LS)x ( ) = 2.9x10 4/yr.
3 Event B: Undamaged Systems Mitigate Accident Given Fire-Growtn Stage 1.
The accident initiating event was taken to be a transient with MSIV closure, and at this stage of tne fire the following systems would remain potentially operable 'RCIC or Hf system, the ADS, the RHR system (three trains), and the LPCS (one k si n). The dominant accident sequences and their conditional probabilitM s a:
l . Class I l
QUX = 5.6x104 QUV = 9.2x10-6 -
. Ciass II I
QW = 7.8x10-7
'~
( . Total = 6.5x10-4 -
p, _ gps W *"
. . . -e4. . - em* - -""
. I A-9 Event C: Fire-Suppressed Before Damaging Unprotected Raceways.
The probability of this event is given by the probability of failing to suppress tne fire witnin 10 minutes (estimated time before damage to unprotected racewyas). BNL value for the probability of this event is 0.43.
i Event D: Undamaged Systems Mitigate Accident Given Fire-Growtn Stage 2.
Given fire-growtn stage 2, the following equipment would remain potentially operable: the ADS and the RMR system (2 trains). The dominant accident sequences and tneir conditional probabilities are as follows:
. Class I QUX = 6.0x10-3 QUV = 8.2x10-3
. Total = 1.4x10-2 Event E: Fire Suppressed Before Damaging Protected Raceways.
The probability of this event is given by the probability of failing to suppress the fire witnin 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> (estimated time before damage to protected raceways). However, since only fires in two quadrants can grow to this stage, the probability of Event E is given by:
P(Event E) = 0.5 x Probability of Failing to Suppress tne Fire Within 1 hr.
= 0.5 x 8.0 x 10-2 = 4.0 x 10-2
~
Event F: Undamaged Systems Mitigate Accident Fire-Growth Stage 3.
In this zone, fire growtn Stage 3 repr'esents damage to all shutdown methods, and consequently the conditional failure probability of Event F is 1.0.
A.4.b Quantification of the Fire-Growth Event Tree for Fires in Power-Distribution Panels.
, ' Event A: Frequency of Fires.
The frequency of panel fires is determined from the number of panels l
-; e
- -m..
- l l-4 s 4 m-=_ y" " " ' ' * ' - ~ ~~ ~
A-10 multiplied by the frequency of fires for panel-year. As described in Section 2.2.4.1, seven panels are located in this zone. Thus, the frequency of panel fires is:
7 ,x (2.2x10-4) = 1.5x10-3fyr, Event B: Undamaged Systems Mitigate Accicent Given Fire Growth Stage 1.
In this zone only fires in three panels are capable of causing initiating events (turbine trip transient) and damaging mitigating systems. Such fires cause, at this stage, the loss of either the RCIC of the HPCI system. The detr.inant accicent sequences are:
. Class I QUX = 5.2x10-6
. Class II QW = 1.1x10-8 PW = 9.4x10-8 *
. Total = 5.3x10-6 Event C: Fire Suppressed Before Damaging Unprotected Raceways.
The evaluation of the probability of this event is identical with that for Event C in Section 3.2.2.1.b.
Event 0, E. and F Once the fire has propagated to cable raceways, the quantification of Events 0, E, and F is identical with that given in Section A.4.a for self-ignited cable Raceway Fires.
A.4.c Quantification of the Fire-Growth Event Tree for Transient-Combustible F1res -
Event A: Frequency of Fires.
BNL agrees with the frequency of fires calculated in LGS-SARA, i.e.,
1.7x10-5fyr, e
e 4 g, =
32 b.52 = **fa 8h"** *' "
_____--______--_-_-_____-----_---?---------*
. . . . . - . . . - + - - - -
A-11 Events B, C, 0, E, and F The quantification of these events is identical with that given in Section A.4.a for Self-Ignited Cable Raceway Fires.
A.5 Fire Zone 45: CRD Hydraulic Eouipment Area The fi.re-growtn event tree for all types of fires in Zone 45 is shown in Figure A.S.
A.S.a Quantification of the Fire-Growth Event Tree for Self-Ionited Cable-Raceway Fires.
Event A: Frequency of Self-Ignited Cable Reaceway Fires.
This frequency is calculated in the same way as for Event A in Section 3.2.2.1.a, i.e.,
18,637(1b) 5.3x10-3 172,799 W * ( ) = 1.9x10 4/yr.
3 Events B and C The quantification of these events is identical with that for tne same events in Section A.4.a.
Event 0: Undamaged Systems Mitigate Accident Given Fire-Growth Stage 2.
This stage represents damage to all safety-related equipment except that served by cable raceways or components protected by horizontal separation or ceramic-fiber fire blankets. The only quipment potentially operable is that served by shutdown method A or B (but not both). The dominant accident sequences and their conditional probabilities are:
. Class I QUX = 5.6x10-4 QUV = 1.9x10-3
. Class II QW = 4.0x10-4 QUW = 3.7x10-5 PW = 6.6x10-5 L
. Total = 3.0x10-3
A-12 Event E: Fire Suppressed Before Damaging Protected Raceways.
The probability of this event would be given by the probability of failing to suppress the fire within 30 minutes (time to damage to protected raceways).
However, only fires in one quadrant (northeastern) are capable of damaging equipment associated with botn snutdown methods. So, tne probability of Event E is given by: 1.95x10-1/4=4.875x10-2 Event F: Undamaged Systems Mitigate Accident Given Fire-Growth Stage 3.
This third stage of fire-growth represents damage to all safe-shutdown equipment, and the failure probability associated witn tne event is 1.0.
A.S.b Quantification of the Fire-Growth Event Tree for Self-Icnited Panel Fires.
Event A: Frequency of Panel Fires.
BNL agrees witn LGS-SARA evaluation of panel fires in this zone, i.e.,
3 x (2.2x10-4) = 6.6x10 4/yr. .
Event B: Undamaged Systems Mitigate Accident Given Fire-Growth Stage 2.
This stage represents damage that is confined to the panel in which the fire starts. Fires in two of the three panels can cause a turbine trip transient and at the same . time disan e one high pressure injection system (HPCI or RCIC) and one RHR train. Tc e dominant accident sequences and their conditional probabilities are as fr 'ows:
. Class ! .
QUX = 1.1x10-5 QUV = 1.8x10-6
. Class II PW = 1.3x10-7 QW = 1.6x10-8
. Total = 1.3x10-5..
a no m .7 . - ~
~
A-13 Since only two of the three panels can contribute to the accident sequences, the probability of Event 8 is: 1.3x10-5 x 2/3 = 9.0x10-6, Event C: Fire Suppressed Before Damaging Unprotected Raceways.
The probability of tnis event is identical with tnat for the same event in Section 3.2.2.1.b.
Event 0: Undamaged Systems Mitigate Accident Given Fir'e-Growtn Stage 2.
The quantification of this event is identical with that of Event 0 in Section A.S.a.
Event E: Fire Suppressed Before Damaging Protected Raceways.
Only fires in one of the three panels are capable of damaging protected raceways. So, the probability of this event is given by:
1 3
x Probability of failing to suppress the fire within 30 minutes (time to damage to protected raceways) =
1
-- x 1.95x10-1 = 6.5x10-2 3
Event F: Undamaged Systems Mitigate Accident Given Fire-Growth Stage 3.
This stage represents damage to all safe ' shutdown equipment, and the failure probability associated with this event in 1.0.
A.S.c Quantification of the Fire-Growth Event Tree for Transient-Combustible Fires.
Event A: Frequency of Transient-Combustible Fires.
BNL agrees with the frequency given in LGS-SARA, i .e.,1.7x10-6/yr.
Events B, C, 0, E, and F Given a transient-combustible fire that causes the ignition of cable
- , o Q/
, , - $-u,..,-e-e-.seee. -.e
_____-*]__ *7__-_'_
,P__ '
~ s A-14 trays, the evaluation of all tnese events is identical with that for the same events in Section A.5.c.
A.6 Fire Zone 47: General Equipment Area The fire-growtn event tree for all types of fires in Zone 47 is shown in Figure A.6.
A.6.a Quantification of tne Fire-Growth Event Tree for Self-Ignited Cable Raceways. -
Event A: Frequency of Self-Ignited Cable Raceway Fires.
Tnis frequency is calculateci in tne same way as for Event A in Section 3.2.2.1.a i.e.,
17,791(1b) 172,799(lb) x3 (5.3x10-3
) = 1.8x10-4/yr.
Events B, C, and 0 The quantification of these events is identical with tnat for the same event sin Section A.S.a.
Events E: Fire Suppressed Before Damaging Protected Raceways.
The probability of this event would be given by the probability of failing to suppress the fire within 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> (time to damage to protected raceways).
However, only fires in one quadrant (NE) are capable of damaging equipment associated witn botn snutdown methods. So, the probability of Event E is
-given by: 8.0x10-2/4 = 2.0x10-2, Event F: Undamaged Systems Mitigate Accident Given Fire Growth Stage 3.
This stage of fire represents damage to all safe-shutdown equipment, and the probability associated with this event is 1.0.
A.6.b duantification of the Fire-Growth Event Tree for Self-Ignited Panel Fires. .
Event A: Frequency of Self-Ignited Panel Fires.
J l .
'_..~".
._ l _, .. - . . - '-~~~2- ~~~
- 4. m =e !# - - ~ ' - --- I
I ?
. A-15 BNi. agrees with the frequency of panel fires given in LGS-SARA, i.e.,
5 x (2.2x10-4) = 1.1x10-3/yr.
Event B: Undamaged Systems Mitigate Accident Given Fire-Growth Stage 1.
Fires in three of tne five panels in this zone may be capable of causing an initiating event and disable one RHR train and one core spray train. The initiating transient was assumed to be an MSIV closure, and the dominant accident sequences and their conditional procabilities are:
. Class I QUX = 4.9x10-5 QUV = 8.1x10-6
. Total = 5.7x10-5 Since only fires in t'bree of the five panels are contributors to those sequences, the probability of Event B is given by:
x 5.7x10-5 = 3.4x10-5 5
Event C: Fire Suppressed Before Damaging Unprotected Raceways.
The quantification of this event is identical with that for Event C in panel-fires for Zone 2 (Section 3.2.2.1.a).
Event 0: Undamaged Systems 141tigate Accident Given Fire-Growth Stage 2.
Given that a fire has propagated from the panel in which it originated to adjacent raceways, the quantification of this event is identical to Event 0 in Section A.6.a.
Events E and F It is BNL judgement that, since none of the existing panels are located in the NE quadrant, the progression of the fire to fire growth Stage 3 is not
- possible in this zone.
I
--..:.. . . . . . ._ . . _ .. ... . . . . n .
A-16 -
A.6.c Quantification of the Fire-Growth Event Tree for Transient-Combustible Fires.
Event A: Frequency of Transient-Combustible Fires.
BNL agrees with the frequency of transient-combustible fires given in LGS-SARA, i.e., 1.7x10-5, Events B, C, 0, E, and F '
Given a transient-combustible fire that ignites cable trays, the quaatification of these events is identical with that in Section A.6.a.
e 1
e G
4 e
.h , b. , . . .-=a +
' ' ~
-4
- i **
7 t
.j A. 3 C i D E F pg Undamaged Fise suppsessed systems Undunaged Five suppressed undamaged cable /TC/ panel belore damaging systems belose damaging systems minisate. unpeosected mitigate accident given protected nutigate saceway accident given raceways accident given FGS) FGSt (Failure gives FGS 2) FGS2 IFailure gives FGS31 FGS3
}!
Cose Annual sequence
{8,' frequency cable TC panel OK ii '
l' i
OK i
t OK '
d.0-2
, 9,8-5/1.7-5/ 0.43/0. @ .0-2 1.0-2 1
CM 2.3-7 4.1-8 4.2-8 4.4-4 I8** I888 note) note.
1.1-4 '
CM 4.6-9 8 . 0- 1 t i 1.9-!
l 7 .0, 5 5 CM 7.0-9 1.2-9 3.1-8 1&
j S = For i sI 2.4-7 4.1-8 7.5-8
, yo,,,,non,,,,,, ,,,,,4"*"'Y "
Note: Because cf the evaluation of event E. the probatulity of event C 2.6-7 is not included in the evaluation of the sequence itsquency.
i l' :
1
- Figure A . I. Fire-growth event free for fire zone 20 l'
i
'i
" A a C 9 E F Undamaged
, ; pg,, g F6ee suppsessed Undunaged Fise suppsessed undarnaged systems cable /TC/ panel
. laelore damaging systems belose dunacing systems snitsgate unpsolected matigate psotected mitigate
. accident given raceway accidsnt given saceways accident S Iw88
'1 FGSI FGSI IFailure gives FGS' 2). FGS2 (Failure G ives FGS31 FGS3 Cose Annual sequence
{tf lesqssancy cable Tt' parul
- t OK lj i
OK OK 3' 1.95-1 .
1.0 3,g.3 g6-4/7.2-4 ,
N '
l 2.2-7 4.3 7
, I 4.3-4 L,M i 1.5-7 3.1-7 i f
. NA CM FGS = Fire growth stage 3.7-7 7.4-7 TC = Transicaa combustible , yes,1 annual cose snelt frequency =I 1-6 * -
Nota: Secause of the evaluation of event E. the psobability el event C is not included in Ilie evaluation of the sequssics itsquency.
f Figure A.2 Fire-growili avant tree for lism rone 22 . .
I /
O ,
i..
. . i 1
. 1 1
A 3 C D
Undamaged Fine suppressed I systems Undam e d before spreaqing systems
! mitigate beyond confinement managate
- Fie in Panel accidens given of cabinet "'" tent usan FGSI FGSI (lailuse gives FGS2) FGS2 Core
! Annual senience
- h. ','- frequency panel OK i
I
,i-1 1.8-3 j 3.1-3 -
CM g,4_7 ;
I, 1.9-4 -
n 4 CM 3.4-7 FGS = Fire growih stage - ~
TC = Transient condastMe ' 8-Constituation to CM from T/C fires 2.2-7 i l
Total core melt frequency 7,o_7 I
per year i
Figure A.3 - Fire-powth event tree for fire zone 24.
3
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Fleure A.4 Firargrowth event tree for fire zone 44
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B-1 APPENDIX B : Report of Professor Alan L. Kafka:
, A Critique of " Seismic Ground Motion at Limerick Generating Station," by ERTEC Rocky Mountain, Inc.
INTRODUCTION Although no theory has yet been developed that explains the cause of earthquakes in the eastern United States, seismologists and engineers. are still calle upon to assess earthquake hazards in this region. As the trends of urbanization and industrial 12ation spread throughout the East, the number of requests for earthquake hazards assessments increases. Seismologists must, therefore, respond to'the need for a technical evaluation of the current state of knowledge of earthquake processes at a given site, while also tempering hazard assessments with clearly expressed admissions of their inherent limitations. Thus, in the assessment of earthquake hazards at sites located in the East, two key issues emerge:
(1) A realistic assessment must emphasize that there is no deterministic model that describes the cause of earthquakes in the Eastern United States in general, or (certainly in most cases) at the site in particular.
(2) It is nevertheless incumbent upon seismologists to provide a practical guide for siting critical facilities that incorporates the prekent state of knowledge in the field.
" Seismic Ground Motion at Limerick Generating Station " a report prepared by IRTEC Rocky Mountain. Inc., can be evaluated in the light of these two issues. On the one hand, the report fails to state explicitly that very little is known about the cause of earthquakes in the East in general or at
, the Linarick site in particular. On the other hand, despite this significant - - -
9
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, B-2 omission plus a number of technical problems, the results contained within the report can still be of practical value in the assessment of the seismic hazard 9
at the Limerick Cenerating Station.
Iu particular, the recults shown in Figure 9 of the ERIEC report for the
" Decollement" hypothesis probably yield a conservative estimate of seismi-ground motion at the site. This conclusion is ironic, since " Deco 11ement" is possibly the most speculative of the four hypotheses considered. Nonacheless, the practical application of " Deco 11ement" is ultimately useful, since its essential feature (as far as the calculated seismic hazard is concerned) is that it creats the entire eastern seaboard as one seismogenic zone. This allows for the possibility that large earthquakes (M=7) could occur anywhere in that' area.
The inclusion of calculations of seismic hazard resulting from the other three hypotheses on seismogenic zonation (Piedmont, Northeast Tectonic tones, and Crustal Blocks) also provides insight into the seismic hazard at the Limerick site. The peak ground acceleration curves shown in Figure 9 for all four zonation models illustrate that a very wide range of hazard assessments results from the lack of knowledge of the cause of earthquakes in this region.
Nonetheless it is useful, from a practical point of view, to know how sensitive the resulting hasard evaluation is to changes in the geometry of seismogenic zones.
While these practical results can be gleaned from the . ERTEC report.
Section 3 (Seismogenic Zones) and Section 4 (Seismicity Parameters) contain a number of technical problems. Also, there is insufficient information in the report regarding the earthquake catalogues used in the study. These issues are discussed below.
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B-3 SEISMOGENIC ZONES Section 3 of the ERTEC report describes the seismogenic zones used in the hazard analysis. In this section, seismogenic zone is defined as "[a zone]...
within which earthquakes are considered to be of similar tectonic origin so that future seismic events can be modelled by a single function describing earthquake occurrences in time, space, and size." It is important to note that since the tectonic origin of all earthquak.es along the entire eastern seaboard is at present unknown, all of the hypothesized seismogenic zones discussed in the ERTEC report are highly speculative. The report does not mention this fact. Some fundamental problems with the two more recently proposed hypotheses are discussed below.
Decollement:
This hypochssis is based on an analysis of intensities reported for the 1886 earthquake in Charleston, SC (Seeber and Armbruster, 1981) coupled with results of seismic reflection studies of the deep crustal struct.are of the southern Appalachians (Cook at al., 1979). The seismic reflection profiles have revealed a continous shallow-dipping reflector beneath the southern Appalachians that has been interpreted to be a major deco 11 ament. The inferred decollement has been proposed as the boundary of a seismically distinct block of the earth's crust, i.e. the " Appalachian Detachment" (Seeber and Armbruster, 1981).
Historical earthquake catalogues for the eastern United States (e.g.
Barstow at al., 1980) show a rather low level of seismicity in the Charleston area, and the recent monitoring of the area with a dense seismo, graph network L
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I B-4 has also revealed a relatively low level of activity. Thus, studies of microsarthquake distribution, fault-plane solutions, and earthquake depth have not been very abundant in this region (R= milton, 1981). The hypothesis that the current seismicity in the vicinity of Charleston, SC is occurring along a major decollement surface is, therefore, not well supported by quantitative geophysical studies. The existance of an " Appalachian Detachment" should thus be considered as interesting speculation, but speculation tonetheless.
Furthermore, although pre 14mhary results from deep seismic reflection profiles in the northern Appalachians (e.g. Ando et al. , 1981; Brown et al.,
1982) have also revealed shallow-dipping reflectors, the lateral excent of these surfaces in the northeast does not appear to be as great as in the southern _ Appalachians. Thus, even if " decollement tectonics" were applicable to earthquakes in the southern Appalachians. I have seen no convincing evidence to suggest that this hypothesis should be applicable in the nortfiern Appalachians in general or in the vicinity of the Limerick site in particular.
Figure 5 of the ERTEC report shows the northern boundary of the Decollement zone at about 41*N. No reason for choosing this boundary was t
given in the report.
Crustal Blocks:
According to this hypothesih, the occurrence of earthquakes in the eastern United States is controlled by large crustal blocks. Supposedly, the boundaries of thes's blocks are seismically active and the interiors are relatively inactive. While this hypothesis seems reasonable in principle, and may eventually predict the locations of future large earthquakas, none of the crustal block models that have been proposed (e.g. Diment et al., 1979) l correlate very well with historical or instrumentally located seismicity. '
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l B-5 Lacking any definitive correlation with the only existing records of actual earthquakes, this hypothesis should be considered as interesting geophysical speculation worthy of further investigation, but -
like the " Decollement" hypothesis - speculation nonetheless.
SEISMICITY PARAMETERS Seismic Activity Rate:
The ERTEC report overstates to some extent the conclusions found in McGuire (1977). This is an example of how the report implies (at least in style, if not in fact) that more is known about eastern earthquakes than really is known. My incorpretation of the results of McGuire (1977) and the further studies on this topic by McGuire (1979) and McGuire and Barnhard (1981) is not that the historical race of activity is well determined.
Rather, the value - of these studies is that they show that even though the rates of activity in the East are poorly determined, a reasonable approach to hazard analysis for exposure times of about 50 years in this region is to assume a stationary model of the rate of seismic activity. This approach is useful only in' light of the current lack of knowledge of the cause of earthquakes in this region. Perhaps
- Sis approach should be referred to as being " reasonable" rather than " realistic" (see Table 1 of ERTEC report).
The ultimate test of such an approach to hazard assessments is, simply, the causative mechanism of earthquakes in the eastern United States. Perhaps the historical earthquake activity in China studied by McGuire (1979), for comparison with the eastern United Stat:.s. was anomalously stationary due to a process that is at present unknown. Future investigators may discover that
. . . . ._. . . . . . .. . ..-- . -. .o -
s 9 B-6 the rate of activity in the eastern United States during the past two centuries was anomalously low or high by an order of magnitude or perhaps even more. If, for example, seismic gap theory (proposed for seismic hazard studies in the vicinity f place boundaries; e.g. McCann et al., 1979) is found to be applicable intraplate earrhquakes, then there might be long periods of seismic quies '.c a premonitory to impending large earthquakes in this region.
Does the rate of ac .vity observed for the past 200 years in the East represent an intraplace ' riation of a seismic gap, or is this rate a result of many years of afteral :ks of a large earthquaks such as the New Madrid event of 18117 Such que :1ons can not be answered without a deterministic model of the cause of ear quakes in the East.
Mnin- Magnitude:
It is not clear wh. h hypothese,s are being referred to in the ERTEC report. that restrict t 2 recurrence of Cape Ann. Massachusetts type earthquakas to areas in few England; the author should have cited some references. I suspect, E aver, that the author is referring to an apparent association between' the .orthwest-southeast trend of seismicity in this region, and a landward .ctension of the New England seamounts - that was discussed by Diment e 't al (1972), Sbar and Sykes (1973), and Fletcher et al.
(1973). This trend er ses the Ottawa-Bonnachere graben and Mesozoic intrusions that postdata e initial separation of North America from Africa (Sykes, 1978). The assoc: :1on between the trend of seismicity (the so-called
, " Boston-Ottava seismic be: ') and these tectonic features (possible candidates for ancione zones of weakr is reactivated by the present-day stress field') has '
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been analyzed in detail by Sykes (1978). Further analysis of the correlation l
by Yang and Aggarwal (1981) showed that there are a number of reasons to question the existence of such a seismic belt.
The monitoring of earthquakes by a dense microearthquake network in the northeastern United States reveals a gap in the Boston-Ottawa trend that goes through Vermont (Yang and Aggarval,1981). This gap (althcugh not as distine'.)
l can also be seen in the historical record of seismicity (e.g. Chiburis,1981).
In addition, the pattern of crustal stress in this region appears to be different to the southeast of Vermont than to the northwest (e.g. Yang and Aggarwal, 1981). This observation suggests that earthquake processes may be t
different in the cluster of seismicity that lies to the southeast of Vermont than it is in the northwestern part of the Boston-Ottawa trend.
There is, therefore, no convincing geophysical evidence to support the existence of a Boston-Ottawa seismic belt within which earthquakes are of similar tectonic or'igin. Hence, I see no reason to exclude earthquakes near Cape Ann, Massachusetts from the Piedmont region. If the 1982 earthquake in New Brunswick, Canada is to be included in this province, as stated in the ERTEC report, then certainly earthquakes that occurred near Cape Ann should be.
l l LARGE EARTHQUAKES NEAR THE LIMERICK SITE i
Appendix B of the ERTEC report discusses the credibility of hypotheses that allow an earthquake of the size of the 1886 Charleston event to occur in the vicinity of the Limerick generating station. As stated in Appendix B, calculations of the . hazard at the site are . sensitive to the subjective i-m __ _ _
B-8 probability assigned to such hypotheses. In the main report a subjective probability of ten percent was assigned to the " Decollement" hypothesis, and this hypothesis can be considered to be representative of any hypothesis that treats the entire eastern seaboard as one seismogenic zone, thus allowing for an earthquaks the size of the Charleston event to occur at the I.imerick site.
Since no e:.pianation has been found for the cause of the 1886 Charleston earthquake, there .is no particular reason to exclude such an event from anywhere along tla eastern seaboard. Thus, a probability of ten percent may be an underestimate for the credibility of tectonic hypotheses which would allow a large earthquake (M=7) in the next 50 years in eastern Pennsylvania.
Perhaps the twenty-five to thirty percent probability for the scientific credibility of such an hypothesis (as suggested b'y at least one of the experts consulted in Appendix B) is not unreasonable. Also, in evaluating Appendix B, it would be useful to know the distribution of responses on this issue: 1.e.
how many of the experts assigned a high probability (25-30%), and how many a low probability (0%) to the credibility- of such an hypothesis?
EARTHQUAKZ CATALOGUES There is no mention in the ERTEC report of the fact that there may be a bias in the distribution of seismicity shown in Figure 1 due to incomplete repording and/or recording of events. While the lower bound of a =4.5 (MM b
intensity V-VI) that was used for the part of the study estimating seismic ground action seems appropriate, it is not clear to what extent the
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, B-9 incompleteness of catalogues for smaller events could effect other parts of the study.
Incomplete reporting could, for example, have an effect on the various studies of determination of seismogenic zones. The report states that, '
consistent with the level of effort available for this study, it relies heavily on the work of others (p.1). This approach is justified, and a serious evaluation of the completeness of the catalogues used is justifiably beyond the scope of the study. Nonetheless, the report should state that completeness af catalogues could be a problem. This omission, again, creates an impression that the phenomenon of eastern United States earthquakes is t
better understood chan it really is.
SUMMARY
AND CONCLUSIONS The general writing style of " Seismic Ground Motion at the Limerick Generating Station," a report prepared by ERTEC Rocky Mountain, Inc., gives an unrealistic impression that more is known about earthquakes in the eastern United States than really is known. For example, the report relies heavily on the concept of seismogeni~c zones "within which earthquakes are considered to be of similar tectonic origin." but fails to state explicitly that the
" tectonic origin" of all earthquakes along the entire eastern seaboard remains a mystery. Also, the following technical problems have been found with the report:
l '
- The conclusion derived from studies by McGuire (1977), McGuire (1979),
I
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r B-10 and McGuire and Barnhard (1981) that the rate of seismic activity in the eastern United States is well determined is, at least to some extent, overstated.
4 Earthquakes near Cape Ann. Massachusetts are assumed to be excluded from the " Piedmont" seismogenic zone, and there is no convincing geophysical evidence to support this assumption.
A subjective probability 'of ten percent was assigned to the credibility of any hypothesis that allows an earthquake the size of the 1886 Charleston event to occur in eastern Pennsylvania. This probability is rather low, and a twenty-five to thirty percent probability - suggested by at least one of the experts consulted in Appendix 3 - is not unreasonable.
There is no mention in the report of the fact that there may be a bias in the distribution of seismicity shown in Figure 1 due to incomplete reporting and/or recording of earthquakes.
Despite these significant problems, the results contained in the ERTEC report can still be of practical value. The peak ground motion curves (shown in Figure 9 of the report) for all seismogenic zonation models are of practical value since they illustrate the very wide range of hazard assessments that result from the lack of knowledge of the cause of earthquakas in the East. In assessing the seismic hazard it is useful to know how sensitive the resulting hazard evaluation is to changes in the geometry cf 4
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1 B-11 seismogenic zones. This is particularly true in cases like the East, where all zonation models are very speculative.
The results shown in Figure 9 for the " Decollement" hypothesis probably yield a conservative estimate of the seismic ground motion at the Limerick site. This conclusion is, ironic, since " Decollement" is possibly the most speculative of the four hypotheses considered. .Nonetheless, the practical application of " Deco 11ement" is ultimately useful, since its essential feature (as far as calculated seismic hazard is concerned) is that it treats the entire eastern seaboard as one seismogenic zone. This allows for the possibility that large earthquakes - such as the 1886 event near Charleston, SC -
could occur anywhere in that area, thus resulting in a rather conservative estimate of the seismic hazard at the Limerick generating station.
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B-12 REFERENCES Ando, C.J., Cook, F.A., Oliver, J.E., Brown, L.S., Kaufman, S.,
Klemperer, S., Czuchra,'B., and Walsh, T., COCORP seismic reflection profiling in the New England Appalachians and implications for crustal geometry of the Appalachian Orogen. EOS, Trans., Am. Geophysical Union, 62, no. 45, p. 1046, 1981.
Barstow, N.L., Brill, K.G., Nuttli, 0.W., and Pomeroy, P.7, An approach to seismic zonation for siting nuclear electric pcwer generating facilities in the eastern U.S., USNRC, NUREG/CR-1577, 1980.
Brown, L. , Ando, C. , Klemperer, S. , Oliver, J.E. , Kaufman, S. , Czuclira ,
B., Walsh, T., and Isachsen, Y., Adirondack-Appalachian crustal structure: The C0 CORP Northeast Traverse, EOS, Trans. , Am.
Geophys. Union, 63, no. 18, p. 433, 1982.
Chiburis, E. , Seismicity, recurrence rates, and regionalization of the northeastern United States and adjacent southeastern Canada, USNRC, NUREG/CR-2309, 76 pp., 1981.
Cook, F. A. , Albaugh, D. , Brown, L;,. Kaufman, S. , 0. liver, J. , and Hatcher, R., Thin-skinned tectonics in the crystalline southern App =1=ah4=a=: C0 CORP seismic reflection profiling of the Blue ..
Ridge and Piedmont, Geology, 7, p. 563-567,'1979.
Diment, W.H., Muller, 0.G., and Lavin,IP.M.,' Basement tectonics of New York and Pennsylvania as revealed by gravity ar.d magnetic studies, in Caledonides on the USA, published by Virginia Poly. Inst. and State Univ., Blacksburg, VA, 1979.
Diment, W.G. , Urban, T.C. , and Revetta, F.A. , Some geophysical anomalies in the eastern United States, in The Nature of the Solid Earth, Ed.
E.C. Robertson, p. 544-572, 1972. .
- Fletcher, J.B., Sbar, M.L., and Sykes, L.R., Seismic trends and travel-time residuals in eastern North' America and their tectonic implications, Geol. Soc. Am. Bull.,'89, p.'1656-1676, 1978.
' Hm=41 t on, R. , Geological crigin of eastern U.S. seismicity, in Earthquakes and Earthquake Engineering, Eastern United States, vol. 1 Ed. J.E. Beavers, p.3-24, 1981, ,-
McCann, W.R. , Nishenko, S.P. , Sykes, $.R. , ar.d Krause, J. , Seismic gaps and place tectonica: Seismic potential for major place boundaries.
Pure and Appl. Geophys. , vol. 117, p.1083-1147,1979.'
McGuire, R.K., Effects of uneartainty in seismicity on estimates of seismic hazard for the east coast of the United States, Bull. Sets.
Soc. Am. , vol. 67,' no.' 3, , p. 827-848, 1977.
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- B-13 McGuire, R.K., Adequacy of simple probability models for calculating ~
felt-shaking hazard using the Chinese earthquake catalog, Bull.
Seis. Soc. Am., vol. 69, p.877-892, 1979.
McGuire , P..K. , and Barnhard, T.P. , Effects of temporal variation in seismicity on seismic hazard, Bull. Seis. Soc. Am. vol. 71, p.
321-334, 1981.'
Sbar, M.L. . and Sykes, L.R. , Contemporary compressive stress and saiccicity in eastera North America: An example of intraplate tectonics, Geol. Soc. Am. Bull. , vol. 84, p.1861-1882, 1973.
Scaber, L. and Armbruster, J.G. , The 1886 Charleston, South Carolina earthquake and the Appalachian detach =ent, Journ. Geophys. Res. .vol.86.
no. B9, p. 7874-7894, 1981.
Sykes, L.R. , Intraplate scismicity, reactivation for preexisting zones of weakness, alkaline magmatism, and other tectonism postdating contintneal fragmentation, Rev. Geophys. & Space Phys., vol.16,
- p.621-688, 1978.
Yang, J.P. and Aggarval Y.P., Seismotectonics of the northeastern United States and adjacent Canada, Journ. Geophys. Res. , vol.86, njl. B6, p.4981-4988, 1981.
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