NUREG/CR-3493

BNL-NUREG-5171 1

A Review of the

Limerick Generating Station

Severe Accident Risk Assessment

Review of Core-Melt Frequency

Prepared by M. A. Azarm, R. A. Bari, J. L."Boccio, N. Hanan,

1. A. Papazoglou, C. Ruger, K.'Shiu,

J. Reed, M. McCann, A. Kafka

Brookhaven National Laboratory

Prepared for

U.S. Nuclear Regulatory Commission
NOTICE

This report was prepared as an account of work sponsored by an agency of the United States

Government. Neither the United States Government nor any agency thereof, or any of their

employees, makes any warranty, expressed or implied, or assumes any legal liability of responsibility for any third party's use, or the results of such use, of any information, apparatus,

product or process disclosed in this report, or represents that its use by such third party would

not infringe privately owned rights.

NOTICE

Availability of Reference Materials Cited in NRC Publications

Most documents cited in NRC publications will be available from one of the following sources:

1. The NRC Public Document Room, 1717 H Street, N.W.

Washington, DC 20555

2. The NRC/GPO Sales Program, U.S. Nuclear Regulatory Commission,

Washington, DC 20555

3. The National Technical Information Service, Springfield, VA 22161

Although the listing that follows represents the majority of documents cited in NRC publications,

it is not intended to be exhaustive.

Referenced documents available for inspection and copying for a fee from the NRC Public Document Room include NRC correspondence and internal NRC memoranda; NRC Office of Inspection

and Enforcement bulletins, circulars, information notices, inspection and investigation notices;

Licensee Event Reports; vendor reports and correspondence; Commission papers; and applicant and

licensee documents and correspondence.

The following documents in the NUREG series are available for purchase from the NRC/GPO Sales

Program: formal NRC staff and contractor reports, NRC-sponsored conference proceedings, and

NRC booklets and brochures. Also available are Regulatory Guides, NRC regulations in the Code of

Federal Regulations, and Nuclear Regulatory Commission Issuances.

Documents available from the National Technical Information Service include NUREG series

reports and technical reports prepared by other federal agencies and reports prepared by the Atomic

Energy Commission, forerunner agency to the Nuclear Regulatory Commission.

Documents available from public and special technical libraries include all open literature items,

such as books, journal and periodical artic 'les, and transactions. Federal Register notices, federal and

state legislation, and congressional reports can usually be obtained from these libraries.

Documen ts such as theses, dissertations, foreign reports and translations, and non-NRC conference

proceedings are available for purchase from the organization sponsoring the publication cited.

Single copies of NRC draft reports are available free, to the extent of supply, upon written request

to the Division of Technical information and 1

5

ocument Control, U.S. Nuclear Regulatory Commission, Washington, DC 20555.

Copies of industry codes and standards used in a substantive manner in the NRC regulatory process

are maintained at the NRC Library, 7920 Norfolk Avenue, Bethesda , Maryland, and are available

there for reference use by the public. Codes and standards are usually copyrighted and may be

purchased from the originating organization or, if they are American National Standards, from the

American National Standards Institute, 1430 Broadway, New York, NY 10018B.

GPO Printed'copy price:$ $6 -0 0.
NUREG/CR-3493

BNL-NUREG-5171 1

A Review of the

Limerick Generating Station

Severe Accident Risk Assessment

Review of Core-Melt Frequency

Manuscript Completed: July 1984

Date Published: July 1984

Prepared by

M. A. Azarm, R. A. Bari, J. L. Boccio, N. Hanan,

1. A. Papazoglou, C. Ruger, K. Shiu,

J. Reed*, M. McCann*, A. Kafka**

Department of Nuclear Energy Brookhaven National Laboratory

Upton, NY 11973

*Jack R. Benjamin Associates, Inc.

"Boston College

Prepared for

Division of Safety Technology Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission

Washington, D.C. 20555

NRC FIN A3393
,I,
ABSTRACT

A limited review is performed of the Severe Accident Risk Assessment for

the Limerick Generatinq 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

impacts on the plant components and structures. Particul ar attention is qiven

to uncertainties and critical assumptions. Limited requantification is performed for selected core-melt accident sequences in order to illustrate sensitivities of the results to the underlying assumptions.
-ivACKNOWLEDGEMENT

The authors wish to thank their colleagues' *in the Department of Nuclear

Energy at Brookhaven National Laboratory for many enlightening discussions and

comments throughout this-project.

The work was performed for the Reliability and Risk Assessment Branch

(RRAB) of the U. S. Nuclear.Regulatory Commisison. Mr.,Erulappa Chelliah of

RRAB was the technical monitor':of the project,. The authors wish to:

acknowledge Ashok Thadani,,Chief, RRAB, and Erulappa Chelliah (RRAB) for many

constructive comments on the preliminary and the final drafts of this report.

Finally, we would like to express our appreciation to Denise Miesell,

Susan Monteleone, and Nancy Nelson for an 'excellent job in typing this

document.
_v_

TABLE OF CONTENTS

Page

ABSTRACT .............................................................. 111i

ACKNOWLEDGEMENT ........................................................ iv

LIST OF FIGURES ....................................................... viii

LIST OF TABLES......................................................... ix

SUMMARY ................................................... x

1.0 INTRODUCTION..... ......................................................... 1-

1.1 Background ...................... . . . . ..-

1.2 Objective, Scope, and Approach to Review .................... 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 Fragi ity Anlss . . .2-1

2.1.1 Introduction ..................... . .... 2-1

2.1.1.1 Sensitivity Analys;isfor Sismic Efet ..... 2-4

2.1.1.2 Seismic Section Organization ................ *2-7

2.1.2 Seismic Hazard ........................... .... 2-7

2.1.2.1 Review Approach ............. . . .... 2-7

2.1.2.2 -Seismic Hazard Methodology .................... 2-8

2.1.2.3 Sei smogeni c Zones ....................... 2-10

2.1.2.4 Seismicity Parameters ......................... 2-15

2.1.2.5 Ground Motion Attenuation ................. 2-17

2.1.2.6 Comparison of the LGS Hazard Analysis with

Historic Seismicity........................... 2-18

2.1.2.7 Summary ...................................... 2-21

2.1.3 Sei smic Fragility ............ .. .. .. .. ................ 2-23

2. 1.3.1 Damage Factor . .. ... .. .. .. .. .. .. .. .. .. .. .. .. ... 2-24

2.1.3.2 Upper Bound Accelerations ................ 2-29

2.1.3.3 Reactor Enclosure and Control Structure ......... 2-30

2.1.3.4 Reactor Pressure Vessel Capacities ....... 3-31

2.1.3.5 Potential Impact Between Reactor Building and

C ontai nment . .. . ... .. ................... . .. 2-32

2.1.3.6 Electric and Control Equipment............... 2-34

2.1.3.7 Review of Significant Components ........ 2-36

2.1.3.8 General Fragility-Related Comments .......... 2-43

2.1.3.9 Closure.................-..... . ..... .. ... .... 2-46

2.1.4 References to Section 2.1............... ........ 2-47

2.2 Fire .................................................... 2-55

2.2.1 Deterministic Fire Growth Modeling ............... 2-55

2.2.1.1 Introduction..................................-2-55

2.2.1.2 Summary Evaluation of Deterministic Fire

Growth Model ing ............ .. .. ............ . 2-57

2.2..1.3 Detailed Evaluation of Deterministic Fire

Growth Modeling............................... 2-59
-vi -

TABLE OF CONTENTS (Cont.)

Page

2.2.1.3.1 Fuel Burning Rate................ 2-59

2.2.1.3.2 Fuel Element Ignition............. 2-62

2.2.1.3.3 Fire Near Enclosure Walls or

Corners ........................... 2-66

2.2.1.3.4 Stratified Ceiling Layer ........... 2-66

2.2.1.4 Recommendations for Improving Fire Growth

Modeling............................... 2-67

2.2.2 Probabilistic Fire Analysis Review..................... 2-68

2.2.2.1 Evaluation of Significant Fire Frequencies in

General Locations ....................... 2-70

2.2.2.1.1 Self-Ignited Cable Fires ........... 2-70

2.2.2.1.2 Transient-Combustible Fires ........ 2-71

2.2.2.1.3 Power Distribution Panel Fires ..... 2-71

2.2.2.2 Screening Analysis* ........................ 2-72

2.2.2.3 Probabilistic Modeling-of Detection and

Suppression .. .................... - -*.*.- 2-72

2.2.2.4 Probabilistic Modeling of Plant Damage State .. 2-74

2.2.2.4.1 Zone-Specific Comments............. 2-77

2.2.3 References to Section 2.2 ......................... 2-78

3.0 ACCIDENT SEQUENCE ANALYSIS . .. ... .. .. .. .. . .. .. .. .. ... .. .. . ... . .. . .... 3-1

3.1.1 Plant Frontline Systems ...... ~..................... .... 3-1

3.1.1.1 Overview of the LGS-SARA Approach in Frontline

System Modeling............. . . . ......... ... 3-1

3.1.1.2 BNL Revision and Review of Frontline System

F aul t Trees .............. .. .. .. . ...... . ... 3-3

3.1.2 Accident Sequence Analysis........................ 3-13

3.1.2.1 Overview of LGS-SARA Accident Sequence

3.1.2.2 BNL Review of Accident Sequence Quantificat ion ................................... 3-15

3.2.1 Overview ofthe LGS-SARA Accident Sequence Quantification ........................................... 3-34

3.2.2 BNL Revisions in Quantification of Accident Sequences .. 3-35

3.2.2.1 Fi~re Zone 2: 13 kV Switchgear Room .......... 3-36

3.2.2.2 Fire Zone 25: Auxiliary Equipment Room ....... 3-39

3.2.2.3 Fire Growth Event Trees for Fire Zones 20, 22,

24, 44, 45, and 47 ........................... 3-41

3.2.3 Review Results............................ . .. . .. . ... 3-42

3.3 References to Section 3 ... ... ... . .... . ... ............ .. .. .... 3-44
-viiTABLE OF CONTENTS (Cont.)

4.0 SOME GENERAL ISSUES AND SPECIFIC RECOMMENDATIONS .................. 4.1 Seismic Hazard and Fragility Recommendations ................. 4.1.1 Introduction ......................................... 4.1.2 Seismic Hazard .. ..............................

4.1.3 Seismic Fragility ..................................... 4.2 Fire.......................................................

4.3 References to Section 4.2....................................

Page

4-1

4-1

4-1

4-1

4-2

4-7

4-10

APPENDIX A:

APPENDIX B:

Detailed Review of the Quantification of the Fire Growth

Event Trees .....................................

Report of Professor Alan L. Kafka: A Critique of "Seismic

Ground Motion at Limerick Generating Station," by ERTEC

Rocky Mountain, Inc ......................................

A-1

B-1
-viiiLIST OF FIGURES

Figure No. Title Pg

2.1.1 Crust Block source zone used in the LGS-SARA

(taken from Appendix A, Figure 5 of

LGS-SARA)........................................ 2-50

2.1.2 Comparison of various historical seismicity

curves and the LGS-SARA seismicity curves

from Appendix A for sustained-based peak

acceleration for the Decollement and

Crustal Block, M=5.5 seismogenic zones........... 2-51

2.2.1 Piloted ignition of.EPR/Hypalon cable under

various external heat flux....................... 2-82

2.2.2 Fire suppression model.......................... 2-83

3.1.1 Reduced fault tree for the HPCI system................ 3-24

3.1.2 Reduced fault tree for failure to scram............... 3-25

3.1.3 The seismic event t r e e ...................... 3-26

3.1.4 Seismic event tree for loss of offsite power

with reactor scram ........................... 3-27

3.1.5 Seismic event tree for los of offsite power

without scram. .. .. .............. ............... 3-28

3.2.1 Fire-growth event tree for fire zone 2 ........... 3-45

A.1 Fire-growth event tree for fire zone 20....;....... A-17

A.2 Fire-growth event tree for fire zone 22 ....... A-18

A.3 Fire-growth event tree for fire zone 24...........I.....A-19

A.4 Fire;-growth event tree for fire zone 44 .......... A-20

A.5 Fire-growth event tree for fire zone 455....... A-21

A.6 Fire-growth event tree for fire zone 477....... A-22
-ixLIST OF TABLES

Table No. Title. .Page

2.1.1 Comparison of Mean Frequency of Core Melt Values ..... 2-52

2.1.2 Hazard Curve Contribution to Mean Frequency of

C ore Melt .......................................... 2-53

2.1.3 Hypothetical Mean Frequency of Core Melt (Based

on Individual Hazard Curves) ........................ 2-54

Suppression Data and Calculations Performed for

Suppression Success Probability Self-Ignited

2.2.1 Cable-Raceway Fires ............................... 2-84

3.1.1 Significant Earthquake-Induced Failures ............ 3-29

3.1.2 Mean Values for Random System or Function

Failures Used in Transient Events.................... 3-30

3.1.3 ATWS Mean Random Failure Values.............. .... 3-31

3.1.4 Dominant Seismic Core Damage Sequen e nce........ 3-32

3.1.5 Dominant Seismic Sequences With BNL Changes .......... 3-33

3.2.1 Summary of Fire-Analysis Results.................. 3-46

3.2.2 Critical Locations of Transient Combustible

Materials in the Auxiliary Equipment

3.2.3 Evaluation of Sequence Frequencies of Oil

Fires (Transient Combustibles ...................... 3-48

3.2.4 Summary of Fire-Analysis Results BNL Review .......... 3-49
SUMMARY

Overall, the Severe Accident Risk Assessment (SARA) for the Limerick Generating Station appears to u~se, state-of-rthe-art methodologies for evaluation of

the core melt frequency'due to seismic- and.,fire-initiating events. These re-.

sults are useful in a relative sense and should not be viewed as absolute

numbers. The authors of SARA are well aware of the uncertainties associated

with analyses of these events and provide discussions of the major contributorsto uncertainties.

The procedure used to quantify seismic risk is based on simple probabilistic models which use some data, but which currently rely heavily on

engineering judgment., The analysis does not include a comprehensive

consideration of design and-construction errors and,_ hence, may be

(conservatively or nonconserv~atively) biased.

The method used for estimating the probability distribution on frequency

of exceedance for the seismic hazard 'is a well-established, straightforw ard

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 selection of seismogenic zones and to the assignment of seismicity parameters. It was judged that the various issues raised with regard to

the seismic hazard analysis would i'ndividually have a small impact (less than 'a

factor of 2) on the mean value of the seismic-induced core-melt frequency, but

that the total impact could be moderate (less than a factor of 10).

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 justification for the fragility, values of various components and structures..

Simple audit calculations were performed in 'an attempt to replicate the

results given in the SARA for the mean frequency of seismic-induced core melt

from dominant accident sequences. The simple calculations were generally in

good agreement with the SARA results.,
-xi -

In the analysis of the core-melt frequency due to plant fires, the SARA*

employs state of the art technology for the determination of fire growth, detection, and suppression. In addition, the impact of fires on plant systems is

within the current state of the art. It was found that the analysis was conservative in many aspects, but this is in keeping with current methodologies in this difficult area which is fraught with large uncertainties. Additionally, it was found that part of the analysis, in particular, the deterministic fire growth modeling, has nonphysical aspects which may be either

co~nservative or nonconservative. From the foregoing, the reviewers believe

that it would be difficult to quantify the effect of these uncertainties,

particularly as they relate to probabilistic analyses.

The approach taken on the fire analysis to the identification of critical

plant areas is sound and all these areas appear to have been identified. 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 some of the specific estimates appear to be incorrect.

The cumulative fire suppression distribution function generated in the SARA

does not seem to agree with available data. BNL obtained a distribution fit

(Weibull) to the appropriate data base and therebr-generated a cumulative distribution which, for any given time, yields a lower probability of fire suppression than the corresponding SARA results.

On the basis of the review of probabilistic aspects of fire initiation,

growth, and suppression, a limited requantification was performed of the fireinduced core-melt frequency. An estimated increase in the fire-induced coremelt frequency by overall factor of. 2 is attributed to differences 1) in the

probability of fire suppression ata any given time and 2) in the frequency of

self-ignited cable-raceway fires. A major contribution to the core-melt frequency comes from the stage of fire growth in which all safe-shutdown

*This document provides a review of the impact of fire risk, as analyzed by the

licensee in their April 1983 submittal. The fire analysis presented therein

reflect the fire protection measures described in Revision 1 of the LGS Fire

Protection Evaluation Report*(FPER) (PECo, 1981). Impact of current plant design changes (Revision 4 of FPER, PECo, 1983) which the licensee addressed by

letter, dated July 15, 1983, has not been assessed in this document.
-xiisystems are assumed to be damaged and faulted by the fire. Each of these

contributors was examined separately in sensitivity studies, and they were

found to be equally important. Sensitivity studies were performed with regard

to operator error and it was found that the fire-induced core-melt frequency

was not very sensitive to (one order of magnitude) changes in 1) the failure of

the operator to depressurize the reactor in a required, timely fashion or 2)

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

helpful in assessing these risks at the Limerick plant.
1-1

1.0 INTRODUCTION

1.1 Background

In February 1983, Brookhaven National Laboratory (BNL) issued a report(1) (NUREG/CR-3028) on its review of the probabilistic risk assessment(2) for the Limerick Generating Station (LGS-PRA). The LGS-PRA excluded

seismic events, fires, tornadoes, hurricanes, floods, and sabotage from the

set of initiating events (internal events) that it considered. In April 1983,

Philadelphia Electric Company (PECo) completed a study which included the

evaluation of risk due to seismic-initiating events and to fires that might be

initiated within the plant. This study, the Severe Accident Risk Assessment

for the Limerick Generating Station (LGS-SARA), also included a revised analysis of the offsite consequence analysis with the CRAC2 computer code.

In June 1983, NRC requested that BNL undertake a preliminary, short-tern

review of the LGS-SARA. Results for a portion of this review are given here.

The present document covers the review of seismic and fire methodologies as

they relate to the determination of the core-melt* frequency. At a later date,

results will be presented for the balance of the review, which will cover the

analysis of the core-melt phenomenology, fission product behavior, and offsite

consequences.

1.2 Objective, Scope, and Approach to Review

The 'objective of this work is to perform a preliminary review of the

LGS-SARA including consideration of the core-melt frequency. This includes an

evaluation of the appropriateness of the overall methodology used to identify

structures and components damaged and faulted as the result of 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

*The concept of co-re-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

places in the present report).
1-2

of significant fires and the conditional failure probabilities of mitigating

systEms damaged and faulted during the fire. Finally, a determination is made

of the influence of the findings of this review on the prediction of the

core-melt frequency as calculated in the LGS-SARA.

It is noted at this point, that the determination of the impact of the

findings on the core-melt frequency is qualitative in some places and, at best,

semiqualitative in others. In general-, major uncertainties in the analysis are

highlighted, subjective notions arIe identified, and limited recalculations are

done to focus concerns and indicate sensitivities. A more detailed,

quantitative reevaluation of the core-melt frequency due to seismic events and

to 'fires would be a more time-consuming, resource-intensive enterprise.

This preliminary review of the seismic portions of the report was

conducted over a two-month period by BNL with the assistance of Jack R.

Benjamin Associates, Inc. (JBA). The BNL reviewers included J. L. Boccio

(overall fire hazard and vulnerability review), M. A. Azarm (probabilistic f ire

modeling), C. Ruger (deterministic fire modeling), I. A. Papazoglou (over 'all

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 review) and M. McCann (seismic hazard review).

Finally, JBA subcontracted with Professor A. Kafka of Boston College for a

review of the seismic hazard a~nalysis from a seismologist's viewpoint. The

overall review contained in Volumes I and II was coordinated by R.,A. Bani of

BN L.

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, 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 analyses. 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
1-3

of the core-melt sequence analysis related to seismic events. Similarly, Section 3.2 contains a review of the core melt sequence analysis relating to fire

events. Sections 3.1 and 3.2 rely on information developed in Sections 2.1 and

2.2, respectively'. Section 4 contains a discussion of general issues and

specific recommendations based on this review.

Note that all references are provided locally in the corresponding

sections or subsections.

2-1

2.0 EXTERNAL INITIATING-EVENT CONTRIBUTORS

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

perform a preliminary review of the LGS-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 Motion Hazard at Limerick Generating

Station

Appendix B: Conditional Probabilities of Seismic-Induced Failure

for Structures and Components for the Limerick Generating 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)(1) and the Zion

Probabilistic Safety Study (ZpSS).(2 ) (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 significantly impact the results.

Based on the experience gained from the IPPSS and ZPSS reviews, a preliminary

review of the LGS-SARA was conducted in a short time period in order to

discover the critical issues and to make recommendations to address those

issues which remain unresolved. In contrast to the previous reviews which

consisted of an in-depth evaluation of each section and subsection of the PRA

reports, this review focused primarily on critical areas which may impact the

results. Since both the hazard and fragility calculations for the LGS-SARA

were performed by the same engineers and were based on the identical method-
2-2

ologies used for the IPPSS and ZPSS, many of the issues and concerns generic

to all sites and plants already have been discussed and evaluated.(3) This

review documents the important concerns applicable to the Limerick plant. The

reader is directed to Reference 3 which provides a general point-by-point

discussion of the seisnic risk methodologies used in PRA studies submitted to

the NRC to date. Differences between the current study for Limerick and the

IPPSS and ZPSS reports are discussed in this report.

In the review of the LGS-SARA, JBA assumed that the Boolean equations for

the sequencies leading to core melt are correct. The review performed by the

BNL reviewers addressed the adequacy of the event and fault trees, random

equipment failures, operator errors, and resulting Boolean equations. The

discussion concerning potential discrepancies for these issues is given in

Section 3.1.2.

As part of the review a meeting was held at the Structural Mechanics

Associates (SMA) office in Newport Beach, California, on 8 July 1983. Dr.

Robin McGuire of Dames and Moore, who performed the seismic hazard analysis

while employed by Ertec Rocky.Mountain, Inc.; SMA, who conducted the fragility

analysis; and NUS met with Dr. John W. Reed and Dr. Martin W. McCann of JBA

along with representatives from the NRC. The purpose of the meeting was to

discuss issues raised to date concerning the LGS-SARA and to focus the review

effort on the critical components and issues. Subsequent to this meeting a

,tour of the Limerick plant was conducted on 15 July 1983. Toward the end of

the review, responses from questions the NRC submitted to Philadelphia

Electric Company (PECo).were provided to JBA.(25, 26) In addition, a

meeting was held in Washington, D.C., on September 26, 1983, which included

representatives from NRC, BNL, and PECo (including their consultants). JB A

did not attend that meeting; however,' the transparancies prepared by PECo were

transmitted to JBA. Based on these events, review of the LGS-SARA, and

discussions with the NRC, Sections 2.1 and 4.1 of this report were prepared by

JBA. Note that the information received toward the end of the project (i.e.,

References 25 and 26 and transparencies from the September 26, 1983, meeting)
2-3

was incorporated where the responses clearly resolved the outstanding issues;

otherwise, the concerns raised during the course of the review are documented

in Sections 2.1'and 4.1.

In the review, an attempt is made to look for both conservative and

unconservative assumptions which could signficantly impact the results. In

order to help the reader, an effort is made to indicate, where possible, the

ultimate impact of the issues which halve been raised. Comments are primarily

directed to the mean frequency of core melt or to the individual sequences

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 been adopted to quantify comments made in the review of

the LGS-SARA:

Effect on Mean Frequency

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 appropriate

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 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 discrepancies and, hence, may be biased (note that discrepancies

may be either conservative or unconservative). Because of the newness of

these types of analyses and the limitations pointed out above, the results are

useful only in making relative comparisons. Although more sophisticated

analytical models exist, the limitation of available data dictates that the

simple models used in the LGS-SARA are in a practical sense at the level of

the state-of-the-art.

2-5

2.1.1.1 Sensitivity Analysis for Seismic Effects

The approach used by NUS to combine the hazard and fragility curves is

different fro m the method used by Pickard, Lowe, and Garrick (PLG) for the

IPPSS and ZPSS. 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 fragility 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 directly using numerical integration.

In contrast, the NUS approach differs from the PLG methodology in two

respects. First, NUS included the potential for random equipment 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 core melt, but may have a

moderate effect on the median frequency of core melt relative to the case

where only seismic contributions are included.

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 contributed by the

significant sequences. This exercise also provided a basis for determiining

the possible changes which differences of opinion could produce on the mean

frequency of core melt. -The procedure used was based on the component

fragility curves represented by their median values and combined variabilities

(i.e., the randomness and uncertainty logarithmic standard deviations were

combined). In addition, mean values for the random equipment failure and

operator error events were assumed. This approach is approximate, but gives

reasonable results for mean frequency values.
2-6

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 integrated with the

hazard curves. Table 2.1.1 gives the comparison between the approximate

values calculated as described above and the values reported in the LGS-SARA.

In general, the approximate results compare reasonably well with the values

given in the LGS-SARA. The calculated mean frequency of core melt is 5.3-6

(6.5-6=6.5x10-6) and is within 10 percent of the LGS-SARA value of 5.7-6.

The maximum ratio for individual sequences is a factor of 2.5, which i-s a

moderate effect. However, the difference for sequence TsRPV, which consists

of a single component (i.e., the RPV), is approximately 50 percent. It was

surprising 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 2.1.2 gives the breakd own of the mean frequency of core melt

contributed by the various hazard curves. Over 83 percent is contributed by

the Decollement and the Piedmont, Mmax = 6.3 hazard curves,,with the

Decollement contributing slightly less. The Northeast Tectonic hypotheses,

which is weighted by a probability of 0.3 in the LGS-SARA, contributes-only

about 5 percent.

Table 2.1.3 considers the hypothetical case that only one hazard curve

exists and gives the value for mean frequency of core melt assuming th at 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 mel t values are given in Table 2.1.3

along with the ratios of values compared to the case where the curves are

weighted as assumed in the LGS-SARA. It is interesting to note that if the

Decollement is the only hazard curve, the mean frequency of core melt will

only increase by a factor of 4.0, which is a moderate effect. On the other

hand, if the Crustal Block, Mmax = 5.5 is the o.n~ly hazard curve, the mean

frequency of core melt will decrease by a factor of about 50, which is a large

effect.
2-7

The comparisons given in Tables 2.1.1, 2.1.2 and 2.1.3 give an indication

of the potential sensitivity of the mean frequency of core melt to changes in

the contributions from the different sequences and hazard curves.

2.1.1.2 Seismic Section Organization

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.

Recommendations for actions to address the significant unresolved issues are

presented in Section 4.1 of Chapter 4.

2.1.2 Seismic Hazard

2.1.2.1 Review Approach

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 summarizes 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

appropriateness 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

to 5, in that order.

In Section 2.1.2.6, a preliminary assessment of overall reasonableness

and accuracy of the LGS-SARA hazard curves is made through a comparison with
2-8

results derived from the historic site intensity data. 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 the same

manner, the impact that comments on the seismic hazard analysis have on the

results are indicated where possible.

2.1.2.2 Seismic Hazard Methodology

The approach used in the LGS-SARA seismic hazard analysis is well

established and considered appropriate to estimate the frequency of ground

shaking levels.( 4 .

5), The analysis consists of two basic elem~ents. 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 (i.e., maximum

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 i~aprobability value that

expresses the degree-of-belief 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, maximum magnitudes, etc.).

The final product of this analysis is a family of seismicity curves, each

.having a di~screte probability value associated with it. The discrete

probability values sum to one, implying that a complete probability

distribution on the annual frequency of exceedance has been derived.
2-9

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 implicit decision was made

only to consider those hypotheses for seismogenic zones, source parameters ,

etc., that had a major influence on the estimate of the freq uency of

occurrence. That is, of the many reasonable hypotheses that could be

considered to estimate the ground 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 process being considered. However, the

result is 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 probability

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 frequency is adequately

defined. Although the influence of individual parameters can be reasonably

estimated prior to performing the analysis, it is generally not true that the

analyst can select a set of hypotheses that will adequately define the

probability distri~bution on frequency 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 representation of such a

wide distribution by 6-10 points is not adequate. Certainly, if the entire

distribut ion were known and the points were selected in a prudent manner, this

may be reasonable. However, in the LGS-SARA, six hypotheses and their

discrete probability values were selected beforehand without knowledge of

their counterpart result on the 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.
2-10

Specific examples where. this could be achieved in the LGS-SARA are discussed

in the sections which follow.

In regards to the importance of having an adequate representation of the

probability distribution on the frequency of exceedance, one point should be

considered. A reliable representation of the probability distribution on

sei smi c ri sk (i.e. , core melt) , i s determi ned f or 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 frequency 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

analysis 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 frequency 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 seismogenic zones is based in part on geologic,

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 assigned probability weights

of 0.30, and the Decollement hypothesis was assigned a 0.10 weight. Major

concerns with the zonation used in the LGS-SARA are discussed below.

As described in the LGS-SARA, the Crustal Block hypothesis attempts to

account for the occurrence of earthquakes.-in the northeast by the movement

along the boundaries of large blocks of the earth's crust. It is assumed that

earthquakes occur along block boundaries while the interior areas are

relatively quiet. In the LGS-SARA, eight zones make up the Crustal Block
2-11

hypothesis (see Figure 2.1). Of these, Zone 8 is the dominant contributor to

the hazard at the site. This hypothesis is questioned on two accounts.

First, while the principle that large blocks of the earth's crust may control

the seismicity in the region along their boundaries is reasonable, such a

theory should correlate reasonably well with historic and instrumentally

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 2.1.1 indicates that the

closest proximity of Zone 8 to the LGS-SARA site is a Ipproximately 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 2.1.1, which also shows the distribution of seismicity to

1980, that the northwest boundary of Zone 8 appears 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(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 2.1.1).

Secondly, it is not apparent that the boundaries of seismogenic zones

should be coincident with the perimeter of a large geologic structure. 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 done for Zone 8 towards the southeast.

A redefinition of Zone-8 in the Crustal Block hy pothesis that 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 moderate increase in the median core melt

frequency is considered possible.
2-12

To *consider the possibility that large magnitude events could occur in

the northeast, the Decollement 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 Decollement hypothesis is one of a number of theories being

considered by seismological experts to explain the possible occurrence of

large magnitude events in the eastern U.S. The physical basis of this

hypothesis is the identification of a shall~ow-dipping reflector beneath, and

alon~g the east coast that has been interpreted as, a seismically distinct block

.of the earth's crust. (7,8),

A major concern with the Decolleme'nt hypothesis is the fact that patterns

of instrumentally located seismicity do not correlate well with it. That is,

fault *plane solutions and source depths do not suggest that earthquakes in the

region of Charleston, South'Caro lina, or anywhere else along the eastern

seaboard occur on a decollement surface. In addition, since the evidence that

a major decollement may exi'st generally applies to the southern Appalaci~ans,

it is not clear that a'decollement seismogenic zone should extend to the

northeast in the vicinity of the Limerick site.

At the SMA meeting, discussions with Dr. McGuire revealed that the

Decollement hypothesis was not selected solely on the basis of physical

arguments that it explai ns the seismicity in the east. A principal motivation

was its use as an all-inclusive'*hypothesis', i'n a probabilit istic sense, in

that it allows the possible occurrence of events as large as M6.8. That i s,

an assumption is made in the LGS-SARA that all reasonabl e hypotheses which

would consider the possibility that large-magnitude events could occur in the

vicinity of the plant site are fully represented by the Decollement

hypothesis. Although such 'an approach may provide a best estimate of the

ground shaking hazard at the LGS site, it is not clear that it is appropriate

or adequate for use in the LG.S-SARA. No basis is provided to support the

belief that the 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 Decollement hypothesis

(i.e., b-values and Mmax). Neither is it clear that the Decollement source
2-13

,zone is the most appropriate way to model the occurrence of large magnitude

events in the eastern seaboard.

The use of decollement tectonics to explain the occurrence of large

magnitude events in the east is one of many theories based in part on

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 Decollement source

zone is certainly one that could be used. However, in the LGS-SARA the,

Decollement zone serves as a single physical characterization 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 largemagnitude events in the east. Amiong the possibilities is to allow the

occurrence of M6.8 events in the other source zones defined in the LGS-SARA.

That is, an Mmaxý6.8 would be considered as one hypothesis on maximum

magnitude for each source zone. The basis for this approach is

straightforward. 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 vary. In part a variety of

such theories are~the basis of the seismogenic zone and hypotheses in the

LGS-SARA (i.e., Piedmont, Northeast Tectonic, and Crustal Block). 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.

Furthermore, a preference was given in the hazard analysis to the Piedmont,

Northeast Tectonic, and Crustal Block hypotheses. A combined probability

weight of 0.90 was assigned to them. A 0.10 probability was given to the

Decollement hypothesis. Consistent with this degree-of-belief and the

'consensus in Appendix B that large earthquakes can be expected on pre-exi sting

zones of weakness, the possibility of large-magnitude events in source

hypotheses that define such zones, should be considered. This approach was

discussed at the SMA meeting with Dr. McGuire, and recognized by him to be a

reasonable alternative to model the occurrence of large magnitude earthquakes.
2-14

However, it is the opinion of Dr. McGuire (and-but not necessarily the

consensus of all, the consultants) that the total probability wei~ght 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 agreed universally that such

events could occur at the LGS. The~degree-of-belief assigned to such a

hypothesis varied from zero to twe nty-.five or thirty percent. Presumably the

value of zero is actually a very small number, otherwise there could not have

been the aforementioned universal agreement. At this point in the preliminary

review of the seismic hazard analysis, the value of 0.1.0 is not accepted by

JBA nor all the experts, retai-ned ,in the LGS-SARA. Qualitatively, this value

should be considered a lower bound.

The alternative approach suggested to model large-magnitude events would

produce at least one additional hazard curve for each source zone. By vi rtue

of, the arguments on maximum acceleration, these'additional hazard curves would

be unbounded as is the curve-for the Decollement 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 of 2 or less, compared to the

existing hazard Curves. At higher 'accelerations, these new curves will be

unbounded and thus have nonzero occ~urren ce frequencies, unlike the previous

hypotheses.

With respect to their impact, the fact that these additional curves are

unbounded means that they will have a greater contribution to the mean

frequency of core melt than their counterparts for' each, source zone.

Previously, the Piedmont, Mmax=6.3 and Decollement hypotheses contributed 83

percent of the mean core melt frequency, since they only allowed accelerations

greater than 0.,80g to occur. All zones will have some contribution to the
2-15

mean frequency of core melt. The overall influence of these additional curves

is judged to result in a small increase in the mean core melt frequency.

2.1.2.4 Seismicity Parameters

For a prescribed zone of seismicity, the random occurrence of earthquakes

is defined by the seismic activity rate, the Richter b-value, and the maximum

magnitude that can be generated by the source. 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 least easily misinterpreted. For a prescribed area in the

east, the catalog of earthquake occurrences is generally believed to be long

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 activity per unit area

(i.e., say 104 kin2) the variation can be large. From Table 2 in the

LGS-SARA hazard analysis, the rate of seismicity for the four source

hypotheses varies from 4.33 to 38.0x10-3 events per-year, per-104 km2.

This effect is taken into account in the LGS-SARPA,,. 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 esti mate,-and the true best estimate

for a given source zone. In fact, some experts indicated a preference for a

regional dependence for b-va'lues. Furthermore, there is~the contribution of

statistical variability in b-value estimates derived from the data, which
2-16

depends on the number of data point~s. Thus, as a minimum, two sources of

variability exist in the estimate of Richter b-values:(1) a possible bias in

the use of the 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 hypot~heses have been sampled in the LGS-SARA. It should be

noted that the LGS-SARA did not. directly estimate Richter b-values from the

catalog of earthquake occurrences. In considering the estimate. of b-values,

PECo should consider the results obtained using the historic-data.

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 wa~s considered. Here, the

matter of what the size of the largest events should be 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 Mmax, whi~le 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. However, the existence of such a trend does not correlate very

well with results of recent studies questioning the existence of such a

trend.( 12) Thus, no' definitive basis exists to s upport 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 arguments provided in the LGS-SARA that suggest the 1982 New

Brunswick, Canada, earthquake belongs in this seismic province.
2-17

If the 1755 Cape Ann earthquake is considered to be a 6.0 event,( 14 )

the distribution on Mmax would be modified to reflect the fact that the

largest observed event had a magnitude of 6.0 as opposed to 5.7. If it is

assumed that the two point distribution on Mmax was changed from 5.8 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 would be small.

The hypothesis that a large-magnitude event, the size of the 1886

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 Mmax? This question-and the uncertainty in

Mmax should be addressed by PECo.

2.1.2.5 Ground Motion Attenuation

To describe the attenuation of ground motion with magnitude and distance,

Nuttli'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 value corresponds to a

factor of 1.8 times the median value at the one standard deviation level.

The attenuation relationship was modified in the hazard analysis to

predict sustained-based peak acceleration and to account for the random

orientation of ground motion. This factor is magnitude dependent. Above

magnitude 6.0, sustained-based peak acceleration is 1.23 times the 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 easter U.S. is a difficult task

due to the limited strong motion data available for that region. However, a

number of relationships have been developed and used in probabilistic haza rd

analyses.(199) Results of sensitivity studies are avail able to compare the
2-18

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 generally on the conservative side

(i.e., it gives higher accelerat-ions at a given frequency of exceedance

level).( 9) It is noted however that there can be c .onsiderable variation in

the hazard analysis results for various attenuati~on .relationships. This

suggests that a more compre hensive sampling of attenuation functions is

appropriate, since it is generally believed that the capability to predi-ct

.ground motion in the eastern U.S. is not well established. The impact of

including alternative attenuation hypotheses on the mean, core melt frequency

is considered 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 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 occurred 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 provi~ded in the FSAR describes

event size in terms of Modified Mercalli Intensity. To establish a

distribution of the MM 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,

is 0I + 2.6 -1.39 lnR (2.1)

where: is site intensity

10 epicentral intensityR =distance (miles.)
2-19

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 attempt was made to verify the catalog

reported in the FSAR or to correct the record for inconsistencies. Also, no

uncertainty in the estimate of site intensities was considered. Intensiti es

above MMI equal to IV were considered.

To define the distribution of seismic intensities at the site, *the

Gutenburg-Richter relation that describes the number of events versus

intensity is given~as follows:

log10 N(15 ) = a + b15 (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 acceleration

can be obtained by a transformation of intensity to peak ground accelerati on

using an appropriate relation. To do this, the following equation was

used:(17)

logioA = 0.014 + 0.30IS (2.3)

where A is peak ground acceleration in cm/sec2. 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 i s

assumed, with a logarithmic standard deviation of 0.28 (base-10), which

corresponds to a factor of 1.9 at the one 'sigma level.
.2-20

The distribution on acceleration at the LGS is estimated according to

v(A>a) = uEf(I)"AP-P(A>ajI) (2.4)

where u(A>a) =annual frequency of peak acceleration A,

greater than the value a.

u=seismic activity rate for intensities greater

than or equal to IV.

I*AI doubly truncated exponential distribution on

intensity I with parameter b-ln 10 where &I

is the increment on intensity.

P(A>a I I ) probability of peak acceleration A greater

than a, given an i'nte nsity I. This is

described by a lognormal distribution whose

median is defined by equation 2.3 with a

logarithmic standard deviation, of 0.28 (base

10).

The resul t of thi s computation, using Imax of VI , is. shown i n Figure 2. 1. 2

with selected curves from the LGS hazard analysis. The historic seismicity

curve compares to accelerations around 0.10g from the results obtained from

the Deco~llement and Piedmont 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 the Crustal Block zone, M=6.0, to within a

factor of 2. Since the assumed maximum intensity felt at the site is MMI VI,

the historic frequency curve falls off sharply.

Equation 2.-4 can also be used as a prediction tool by

possibility of site intensities greater than VI to occur.

estimate of the maximum site intensity that can occur must

allowing the

To do this, an

be made. This is
2-21

the same step that was taken in the probabilistic seismic hazard analysis. A

maximum intensity of X was 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 u(A>a) for a maximum intensity of X, is al so

shown in Figure 2.1.2. This assumption allows 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,1 a seismicity curve is estimated assuming a Richter b-value of 0.45

which corresponds to the 0.90 value used for earthquake magnitude in the

LGS-SARA. Again, a maximum intensity MM.I X is assumed. The hazard curve for

this case is shown in Figure 2.2. The effect of assuming a b-value of 0.45

(equivalent to 0.90 for the magnitude scale) results in 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 Decollement 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 adequacy and appropriateness 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 is a well

established straightforward approach to estimate the ground shaking hazard.
2-22

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

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 approa ch 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 inappropriately 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 Decollement source was used as an all-inclusive 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 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-impact of not considering the

uncertainty in this parameter is considered to be small.

Particular concern was expressed with regard to the estimate of maximum

magnitudes. For the Piedmont source, evidence was presented that questioned
2-23

the basis for establishing the distribution on maximum magnitude.

Specifically, the Cape Ann events should be included in the Piedmont province

and considered in the estimate of Mmax. The 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

this estimate was considered, neither was there any physical 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

possibility-of an MM intensity X event are consistent with the hazard curves

estimated for the Piedmont, Decollement, and Northeast Tectonic zones at low

accelerations.

The recommendations given in Section 4.1.2 are 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, their total effect could

be moderate.

.2.1.3 Seismic Fragility

The preliminary review of the seismic 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 seismic risk analysis. As described

in Section 2.1.1, the results of the meeting with SMA and theý 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
2-24

consi~dered in relationship 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 2.569 and, thus, was not included in the sequences. This

component was inspected during the plant tour, and its capacity value is judged

to be reasonable.

The comments concerning the'seismic fragility analysis are organized 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 concerns. 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 overlooked, but

which are philosophical in nature (i.e., do not have an immediate resolution)

or which are unlikely to have a major i~mpact on the results. Finally, Section

2.1.3.9 gives final closing comments on the prel1iminary review of the seismic

fragility analysis in the LGS-SARA.

Throughout the discussion recommendations are made for additional

information. Section 4.1.3 summarizes the recommendations for additional

actions required to resolve the fragility-related issues which have been raised

but not answered or completely resolved;"

2.1.3.1 Damage Factor

Three adjustment factors are used in the LGS-SARA to estimate capacity to

resist earthquakes. The hazard analysis'documented in Appendix A of the

,LGS-SARA presents th~e frequency of exceedance for seismic hazard in-terms of a

sustained-based peak acceleration parameter. 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 effect ive peak accelerations to reflect the less damaging

characteristics of low magnitude earthqu'akes. This-adjustment is identical to

the adjustments made in the IPPSS and the ZPS S. 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
2-25

response and, hence, damage caused by lower magnitude events. It is implied in

Reference 18 that the adjustment factor should be 0.5 for magnitudes less than

M5 and distances less than. 20 km. For magnitudes greater than M7 and distances

greater than 40 kin, 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 capacity of

structures by a factor of 1.4. The justification for this factor as discussed

in Section 4.1.3 is very similar to 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 phenomena 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 curves. 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 (i.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-SARA. More realistic values of 3 to 4

-for the ductility ratio should have been,' used. The use of low ductility values

compensated 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 thi s

expl anation.

The justification for the duration factor of 1.4 was also reviewed. The

underlying basis for the duration factor is recent work reported in Reference

19. As documented in this report, a series of analyses were conducted to

investigate the response of single-degree-of-freedom (SOOF) nonlinear

oscillators to real earthquake motions. Earthquakes which varied in magnitude

from 114.3 to 117.7 were used. It was explained at the meeting at $MA that a

duration factor is required to correct the capacity of SDOF systems when!

subjected t~o earthquakes less than M6 to obtain the same level of damage.
2-26

The ductility factor based on the approach developed by Riddell and 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 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 nonlinear 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.,

In an effort to verify the earthquake duration factor used in the LGS-SARA

fragility analysis, the data contained in Reference 19 was reviewed. As described above, arguments which support the use of an earthquake duration factor

are based on the assumption that seismic events of magnitude smaller than M6

contribute less to the likelihood of failure than predicted by the Riddell/Newmark model. It was on this basis thlat the median value of 1.4 was derived for use in the LGS-SARA. As a check, the data as 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 group. From the histogram for

each group the median response factor and -logarithmic standard deviation were

derived. 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.

Response Factor

Data Group F a_

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 Oc

0.12 Or

0. 08B.
2-27

From this comparison, it appears that the median factor used in the

LGS-SARA is over estimated by 14 percent (i.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 negligible effect on the median.

A second look at the scale factor data was taken by dividing the data i n

short and long duration (TD) groups. The data were divided according to

whether durations were less or greater than 2.5 seconds, as defined in

Reference.19.

In this case the artificial time history is *in the TD> 2.5 second group.

Basically all the records in the M 6 group were in the TDOM. second data set

with one exception. The UCSB Goleta recording of the M5.1 (M55.6) 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.

Response Factor

Data Group F

TD<ý2.5 sec. 2.85 0.51

ID>2. sec. 2.05 0.26

FED = FTDý2s/FTD>2.5 1.39 0.57

LGS-SARA 1.40 0.20 Oc

0.12 Or

0.08 ou

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 would seem to

point out, that although Reference 19 provides a clear indication of the

duration effect of strong motion on structural damage, results reported are
2-28

limited. in their application because of the relatively small data base. As

discussed at the meeting with SMA, the use of the M6 cutoff to establish the

duration factor is a gross characterization of a process that is continuous

over magnitude and/or duration. Thus, a median duration factor should

preferably be a function of magnitude. Data to establish such a function are

not available. Furthermore, Reference 17 also suggests that the~duration

factor has a frequency dependence. This was not taken into account in. the

LGS -SARA.

The estimate of the logarithmic standard deviation of the duration factor

in the LGS.-SARA appears low. In particular, due to the uncertainty in

estimating FED and the limited data base, OuO0.O8 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 Oc 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 19 are consistent with engineering judgment and observed earthquake

damage, the approach used in the LGS-SARA is a simplification of a complicated

issue.

The arguments leading to the 1-.4 duration factor, when included with the

ductility adjustment factor based on Reference 20, are generall y reasonable for

earthquakes with magnitudes less than M6; however, as discussed above, the 1.4

factor may be slight-ly 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 19 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.
2-29

Based on Reference 21, the hazard curve for the Decollement seismogenic

zone is the only curve which has average magnitudes equal to or greater than

M6.O0. For sustained-based peak accelerations equal to or greater than 0.40g,

the average magnitudes equal or exceed M6.0. It is estimated that if the

duration factor is changed to unity for this region of the Decollement 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 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 the reviewers by PECo to

support the LGS-SARA. This is the same argument which is given in the IPPSS

and ZPSS reports(1,2) for limiting accelerations. The explanation for

limiting upper-bound accelerations consists of two steps. The first step i s

the assumption that there is a maximum intensity associated with each source

zone corresponding to the maximum magnitude for that zone. This is assumed to

be true by seismologist. The second step related 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 systematically described in the

MMI scale. If the accelerations are higher than predicted, then a higher MMI

value (corresponding to more damage) would occur. However, since the maximum

MMI values are limited by the seismologist, a higher acceleration i s not

possible. The' problem with limiting accelerations for the Decollement 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
2-30

higher MMI val ues al so incl ude fail.1ure of most (if not allI) masonry structures.

As explained at the meeting at SMA, it was conservatively decided not to

truncate the Decollement hazard curve.

It also foll'ows 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 the hazard curves produces the sanie

result. Thus if upper bounds exist for lower intensity values, similar limits

should apply for higher intensity values for engineered concrete structures.

However, it is difficult to quantify this belief at this time. In conclusion,

the assumption not to truncate the Decollement hazard curve is on the

conservative 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.Og (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 Decollement

hazard curve has a small effect on the results of the LGS-SARA.

2.1.3.3 -Reactor Enclosure and Control Structure

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, the diaphragm capacity is adequate to redistribute

forces as the various wall sections yield.
2-31

Based on a median capacity of 0.90g, 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 VesselCapacities

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 CRD guide tubes,

it was assumed that the containment structure had an effective damping value of

10 percent. Since the-original analysis of the combined containment/NSSS was

based on 5 percent damping for the concrete structure, a 1.3 factor, which

increased the capacity of the RPV components, was developed from the ground

spectral accelerations by SMA.

It is not obvious from the LGS-SARA or the calculati-ons that the 1.3

factor is appropriate since the stresses in the containment structure may not

be sufficiently high to warrant the assumed 10 percent damping value. The

median capacities of the three RPV components range between 0.67 and 1.37g,

while the limiting median capacities of the supporting containment structure

components are as follows:

Sacrificial shield wall 1.6g

Containment wall (shear failure) 3Mg

RPV pedestal (flexural failure) 2.8g

The upper portion of the RPV 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 major portion of

the resistance comes from the shield wall , then 10 percent damping is probably

appropriate. On the other handS if the input to the RPV is dominated by the

support at the top of the shield wall, 10 percent damping may be too large.
I .

2-32

If the 1.3 damping response factor is changed to unity, which is the most

conservative assumption for this factor, it is estimated that the mean

frequency of core melt wo uld increase by a factor of approximately 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

should be included for the RPV-related component capacities. Also, the model

used to develop the capacity of the RPV lateral support is approximate and,

hence, additional uncertainty is present. It is believed that due to the SRSS

operation for combining uncertainties, the effect of these additional

uncertainties would have a small effect on-the mean frequency of core mel t.

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 0.45g the two structures begin

to impact at elevationý 289 feet (it is believed that elevation 283 feet is the

correct level). It is also stated tha t since the reactor building shear walls

are expected to fail between 0.74g and 1.Og 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 cause high

frequency motions which could affect electrical and control equipment. Based

on inspection of the plant, the gap between the reactor building 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
2-33

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 the chance of failure of the electrical equipment

located in the reactor building will be increased by impact between the two

structures. 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 l earned 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 pieces will be small since the reinforcing

steel will tend to hold any fractured 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; however, small lines may

be damaged by falling concrete pieces.

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 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

lateral supports were found on the isometric plans in the plant engineering

office. It was confirmed that this line belongs to the RHR system and is a low

temperature line. The first critical support was located approximately 10 feet
2-34

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 flexibility to resist two-to-three inches of

relative movement. A stress of approximately 10,000 psi would be caused by a

three-inch relative displacement which, when added to other stresses, probably

would not significantly affect the core melt frequency distribution.

Several small lines (probably control -related) were attached to a valve

close to 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. This should be systematically confirmed for all

small lines.

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.

2.1.3.6 Electrical and Control Equipment

The mean frequency of c 'ore melt repo rted in the LGS-SARA is 5.7x10-6 per

year. About 60 percent of this value is contributed by sequence T5E5UX,

which includes the following five electrical or control components which are in

series:

440-V bus/SG breakers

*440-V bus transformer breaker

125/250-V dc bus

*4-KV bus/SG

Diesel-generator circuit breakers
2-35

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 the LGS-SARA for sequence T5E5 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 independent unit (e.g., if there

are two independent switchgear breakers instead of only one),.the mean

frequency of core melt will increase by approximately 0.4x10 6 per year.

Several issues should be considered in determining whether additional

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 i n

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?

The second consideration is the question of independence between

components. It can be argued that identical units have high capacity

dependence (i.e., if two units of the same component are subjected 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 also may have high response dependence. This 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

bujs/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 the frequency of failure of one unit. On the other hand, i f

the units are constructed differently and/or placed at different locations,

they may approach being independent which in the extreme case implies that the

frequency of failure is approximately equal to the sum of the individual

failure frequencies.
2-36

In order to evaluate the impact' of this concern PECo should determine the

number, location, and characteristics of the electrical and control equipment

which are part of sequence TsEsUX, 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 performed 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) The following comments are given for

the 17 significant components.

Offsite Power (500/230-Ky Switchyard) (Si) - 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 (S2) - This component is not. a major contributor

to the mean frequency of core melt. The capacity of the tank is based on the

weakest failure mode which is shell buckling. A small ducti lity value of 1.3

was assumed. This is probably reasonable but may not be conservative since a

buckle could cause 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 tank is on rock. It was not apparent from the tour of the Limerick site
2-37

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 summary, the fragility parameters for the condensate

storage tank appear to be reasonable.

Reactor Internals (S3) - The capacity of this component is limited by

the strength of the shroud support. The exact failure location was not given

in the calculations. The capacity factor was derived based on 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 effect of increasing the logarithmic standard deviati on

for time history variability is small.

As discussed in Section 2.1.3.4, the factor of 1.3 which increases the

capacity of the reactor internals 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

were 1.0, would be an increase in the mean frequency of core melt by a factor

of approximately 1.10.

Reactor Enclosure and Control Structure (S4) - The capacity of this

component is controlled by the failure of the lowest story shear walls and i s

based on adjusting the forces obtained from the original design analysis to

median-centered values. As discussed in Section 2.1.3.3, the median capacity

is better represented by 0.90g (as compared to 1.05g given in the LGS-SARA).

This change would increase the mean frequency of core melt by approximately 20

percent.

It was noted that the uncertainty value for modeling was only 0.10.

Because of the approximate nature of the analysis which was conducted, a val ue

of at least 0.20 is more appropriate. In comparison, a~modeling uncertainty
2-38

value of 0.17 was used for testing in developing ,the fragility for equipment,

which gives an indication of a value for this factor that is more reasonable.

As discussed in Section 2.1.3.1, a ductility value of 2.5 assumed for the

case of shear wall flex ural failure is low. However, the effect of this value

is balanced by the extra factor'.assumed for earthq uake size effects used to

adjust. the hazard curves from sustained-based peak accelerationto an effective

peak acceleration parameter.

CRD Guide Tube (SO) The capacity of a CRD guide tube is controlled by

functional binding of the control rod due to bending. The fragility parameters

are based on test results coupled with the response of the guide tube

,calculated during the plant design.* The test capacity was increased about 20

percent based on judgment since failure was not observed in the tests. This is

probably on the cons~ervative side.

Since the CRD guide tubes are attached to the reactor pressure vessel

(RPV) the comments above for the reactor internals, pertaining to use of a

one-time analysis history and containment damp in'g, also apply to the CRD guide

tube analysis.

Reactor Pressure Vessel (SO -Th capacity of the RPV is due t~o the

potential failure in the weld between the connections of the top supports for

the RPV and the top of the shield wall.. An approximate analysis was- used to

determine the median capacity factor, wherein the total capac ity was assumed to

be equal to the sum of the capacities from the support skirt and failure in the

weld at the top support. A 0.10 Uncertainty value was included for modeling,

whi ch, i n the opi ni on of the rev iewers, i s small1. Simil1ar to the comments made

for the reactor enclosure and control structu re above, a value of at least 0.20

is appropriate for this type of approximate analysis. The effect of this-size

of increase in variability would have a small effect on the mean frequency of

core melt.

The comments given for the reactor internals, pertaining to one-time

history and containment damping, also apply to the RPV capacity.
2-39

Hydraulic Control Unit (S7) - The components of the hydraulic control

unit consist of valves, tanks, piping, and electrical controls. The fragility

parameters are based on tests and fragility cal culations performed for the

Susquehanna nuclear power plant. In essence, the median capacity from

Susquehanna was scaled by the ratio of the two SSE peak ground acceleration

values (i.e., 0.10/0.15). It is not. apparent from the documentation in either

the LGS-SARA nor the supporting calculations for this component whether the SSE

scaling from Susquehanna is appropriate. The concerns include possible

differences in the foundation condition and, hence, the response of the reactor

enclosure, locations of the hydraulic control units in the two pl ants (i .e. , i s

one unit higher, therefore it has a higher response?) and, finally, con

struction and,. hence, similarity of the two units. These issues should be

addressed by PECo.

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 (TRS) and the required

response spectrum (RRS) at different frequencies. The total range of values

for different frequencies and for the two horizontal directions were used to

calculate the uncertainty value. 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

directi ons-and the maximum of the largest ratio. If this approach is used, the

uncertainty val ue is approximately one-third (i.e., 0.09 compared 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
2-40

value. The final median-Value is essentially equal to the average of all

ratios of the TRS to RRS values. Since there Iwas no failure, the median value

is on the conservative side.

It should be noted that the total- uncertainty logarithmic standard

deviation value, for the hydraulic control unit is 0.52 which is the 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 consi-dering other uncertai nti es d~ue,,to modeling which have not been

included.

.SLC Test Tank (S.8) -The capacity for the SLC test tank is based on

generic calculations for rigid equipment.' This tank is ýsupported 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 analysis performed for

this tank is not applicable to the actual, component.

The capaci ty of the anchor bol ts whi ch attach 'the- base of the f our ,col umns

to the concrete floor should be analyzed. The response factor should be

recalculated taking into account the flexbility of the t ank, and the actual'

charactertistic of the four columns.- Becaus e analyses assumed the tank to be.

rigid, the capacity may be overly conservative for this effect.,

If the tension force in the columns or anchor bolts control the capacity,

the earthquake component factor may be as low as 0.71l (as compared to 1.04

which was assumed in the generic component, analysis). Since the capacity may be

.controlled by a ductile element, a ductility value, greater than 1.0 may be

appropriate. In summary, *a component-specif'ic analysis should be conducted for

the SLC test tank.

Nitrogen Accumulator (S') -The nitrogen accumulator is 'described in the

calculations as an 18-inch diamet~er ,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 component is based on extrapolating an analysis from
2-41

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 buckling 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 failure mode apparently

was not checked. Th 'ere are no stiffening elements in the vicinity of the

anchor bolts, which may mean that tearing of the base plate flange is the

.weakest capacity. The possibility that this potential failure mode was

overlooked in the original design calculations should be checked.

The uncertainty value for modeling error was assumed to be 0.10 which is'

small. A value equal to 0.20 would be more appropriate; however, this change

would have a small effect on the frequency of core melt.

440-V Bus/SG Breakers (S11) - The capacity of this component was

developed in a similar manner to the capacity for the hydraulic control uni t,

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 minimum 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 doubli ng

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 summary, the fragility parameter values for this component appear

reasonable, but it was not possible to check all the calculations. Since this

component is a significant contributor to the mean frequency of core melt, a

specific analysis should be conducted for this component.
2-42

40VBus Transformer Breaker (S12), 125/250-V DC Bus P1l3), 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 component's 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. Comments concerning these three components are the same as given

below for the diesel generator circuit breakers.

Because these three components contribute signficiantly to the mean

frequency of core melt, a specific component analysis should be conducted for

each.

Diesel Generator Circuit Breakers S5)-The capacity of the diesel

generator circuit breakers is based on an analysis 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 comments

above). The same issues for that component also apply to the diesel generator

circuit breakers (and also the three components above, i.e., S12, S13, and

Since this component is a significant contributor to the mean frequency of

core 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. However, the actual

fragility parameters are based on generic passive 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 .
2-43

RHR Heat Exchangers (S17) - The capacity of the RHR heat exchanger was

obtained by scaling the 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 controlli ng

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 will1 be

somewhere between 0.71 and 0.93.

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 Comments

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 th e review

are documented for completeness. The reader is directed to Reference 3 which

gives a more detailed discussion of some of these general issue s.

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 very sensitive to the relationship

between the equipment fundamental frequency and the frequency correspondi ng to

the peak of the floor response spectrum, it is easy to visualize cases where a

sl ight shi ft i n f requency coul d mean a factor of 2 or 3 (or even more) i n the

value of the spectral ordinate. Thus the logarithmic standard deviation for

response should be developed on a case-by-case basis.
2-44

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 i n the

LGS-SARA. On the other hand, the-median capacity values are probably on the

low side. These two effects likely are selfe-compensating.

No uncertainty was assigned to the ground response spectrum factor used in

the analysis. By definition this implies that this is the 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 information which were used. The use of

subjective or data-based information (either analysis or testing) should be

specifically noted to inform the reader. In addition, sensitivity analyses

should be performed to indicate the robustness of the assumptions. This i s

particularly applicable to Chapter 3 where the fragility, hazard, and systems

information is combined 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

TsEsUX, it appears that any additional capacity or response-related

dependency effects would not have a significant impact on the mean frequency of

core melt. For the case of TsEsUX, Section 2.1.3.6 discusses the

implications if additional components were added to t-he series expression . For

the current Boolean expression for the TSESUX sequence, if any additional

dependency exists, the frequency of- core melt would decrease. As discussed in
.2-45

Reference 3, there are potential dependency effects which could effect the

fragility values for cable trays and piping systems, although it is likely that

the current capacity values account for these effects.(3)

Another important issue is the use of ductility factors for one degree of

freedom (SOOF) models to represent multidegree of freedom (MOOF) structures or

equipment.(3) Research is required to resolve this issue. At the present ,

not enough uncertainty is generally assigned for this situation.

As discussed'Section 2.1.1, design and construction discrepancies -are 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 best, the results

of a seismic PRA can only be used to make relative comparisons.

One concern which was raised is potential leakage through internal

components caused by seismic motion, thus bypassing a closed valve barrier.

This probably is not a major problem but should be formally verified by PECo.

The MSIV and purge and vent valves are'important examples. Also, the type of

SRV used at Limerick has a history of sticking randomly in the opened position

(i.e., failing to close after the signal is received). The possibility that

seismic motions could increase the likelihood of this type of failure should be

addressed.

The potential for secondary components failing, falling, and impacti ng

primary safety-related components apparently has not been systematically

addressed since the plant is still under construction. The potential effects

of block walls failing has been considered. Other components could also. be a

potential hazard. At the completion of construction, secondary components

should be reviewed and their capacities incorporated into the LGS-SARA if they

are weaker than the primary 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.
2-46

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 factor 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).(24)

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 effectiveness of.

evacuation was quantified for the various accident classes. The argument for

limiting 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 acceleration limits for the hazard curves, the rationale was

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.

.Because of the concern for potential failure of the control room ceiling

at the Indian Point Nuclear Power Plant (Ref.3), the control room ceiling at

Limerick was inspected during the plant tour. The ceiling -at Limerick was

found to consist of a light weight "egg-crate" structure which is supported by

wires and braced between walls. There is no tr ansite, reflector panels located

above the ceiling as found at the Indian Point Power Plant. Therefore, it is

concluded that the ce iling at Limerick does, not pose an undue hazard during a

seismic event.

2.1.3.9 Closure

The LGS-SARA differs from the IPPSS and ZPSS in that the mean frequency of

core melt is dominated primarily by five electrical components in series, which

have nearly the same median capacities. In contrast, nonelectrical. components

and structures controlled the results of the IPPSS and ZPSS.
2-47

The capacities for the LGS-SARA electrical components are based on generic

tests and are not component specific. This approach is reasonable as long as

the components do not control the final results. Based on the response given

by PECo at the September 26, 1983, meeting, it appears that scaling the

capacity values by the ratio of the SSE accelerations for the Susquehanna and

Limerick (i.e., 0.10/0.15) may be overly conservative by a factor of 2 for the

electrical components. 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 Cornmission, 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 Seismi-c 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 Safety

Analysis Report, 1983.

7. Cook, F. A., D. Albaugh, L. Brown, S. Kaufman, J. Oliver, and R. Hatcher,

"Thin-Skinned Tectonics in the Crystalline Southern Appalachians: COCORP

Seismic Reflection Profiling of the Blue Ridge and Piedmont," Geology,

Vol. 7,. pp. 563-567, 1979.
2-48

8. Seeber, L., and J.G. Armbruster, "The 1886 Charleston, South Carolina

Earthquake and the Appalachian Detachment," Journ. Geophys. 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 Anomalies

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. Fl etcher, J. B. , M. L. Sbar, and L. R. Sykes, "Sei smi c Trends. and Travel -

Time Residuals in Eastern North America and Their Tectonic Implications,"

Go.Soc. Am. Bull., Vol. 89, pp. 1656-1976, 1978

13. Yang, J. P. and Y. P. Aggarwal, "Seismotectonics of the Northeastern United States and Adjacent Can ada," Journ'. Geophys. Res., Vol. 86, No. B6,

pp. 4981-4988, 1981.

14. Street, R. L., A. Lacroix, "An Empirical Study of New England Seismicity:

1727-1977,".Bull. Seis. S .oc. Am., Vol. 69, pp. 159-175, 1979.

15. Nuttli, 0. W., "The Relation of Sustained. Maximum Ground Acceleration and

Velocity to Earthquake Intensity and Magnitude," Report 16, Misc. Paper

S-7-1, U.S. Army Engineerý Waterwiays Experiment Station, Vick .sburg, Miss. ,

1979.

16. Cornell, C. A. and H. Merz, "Seismic Risk Analysis of Boston," Journal of

the Structural Division, American Society of Civil Engi~neers, ST1O, 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.
2-49

18. Kennedy, R. P. , "Comments on Effective Ground Acceleration Estimates," SMA

Report 12901.04R, February, 1981.

19. R. P. Kennedy, et al., "Engineering Characterization of Ground Motion Effects of Charateristics of Free-Field Motion on Structural 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," Department of Civil Engineering, Report UILU 79-2016, Urbanaý, 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 Regulatory Commission,

NUREG/CR-2405, October, 1981.

23. Bohn, M. P., "Interim Recommendations for a-Sim plified Seismic

Probabilistic Risk Assessment Based on the Results of the Seismic Safety

Margins Research Program," Lawrence Livermore National Laboratory, Draft,

April 15, 1983.

24. Letter from R. D. Campbell t o J. W. Reed, dated July 11, 1983.

25. Letter from J. S. Kaufman to A. Schwencer, dated August 24, 1983.

26. Letter from J. S. Kaufman to A. Schwencer, dated August 29, 1983.
CD,

Figure 2.1..1 Figu? Crust 21.1 Block source zone~ used in the IGS-SARA (taken from Appendix A, Figure 5 of LGS-SARA).
2-51

10-2. 4

Historic Seis-icity Cur-ves

I. Imin = IV, IM = X, b = 0.45

2. Imin= IV, 4M = X, b=O0.72

LS1-SPR4 Seisricity Curves

4. Decolletrut

1 01-11 U

(1)

r_

U

LA.J

'4-

0

>1

Q;

10- 4-

i n-5...

i U~. Cu~IL~I..K

1

5

4.

2

3

10-7 4

0 0.20 0.40 0.60 I

0.80

Peak Ground Acceleration (g)

Figure 2.1.2 Comparison of various historical seism~icity curves and

the LGS-SARA seismicity curves fromi Appendix A for

sustained-based peak acceleration for the Decollernent

and Crustal Block, M=5.5 sei~smogernic: zones.
2-52

Table 2.1.1 Comparison of Mean Frequency of Core Melt Values

Sequence

TSESUX

TSRPV

TSR BCM

TSE 5 W

Total

*4.0-6=4. 0x1- 6

Contribution to Mean Frequency of Core-Melt

Approximate Analysis LGS-SARA Value-s

2.8-6* 3.1-6

9.5-7 9.6-7

4.4-7- 8.0-7

6.0-7 5.4-7

3.5-7 1.4-7.

1.1-7 1.1-7

5.3-6, 5.7-6
2-53

Table 2.1.2 Hazard Curve Contribution to Mean Frequency of Core Melt

Hazard Curve

Deco 1 1ement

Piedmont, Mmax=6 -

3

Piedmont, MmaxS.S8

Northeast Tectonic

Crustal Block, Mmax=6 .O

Crustal Block, Mmax=5.5

Total

Contribution to

Mean Frequency of Core Melt

2.1-6

2.3-6

5.4-7

2.4-7

6.2-8

1.5-8

5.3-6

Percentage

39.9

43.7

10.3

4.6

1.2

0.3

100.0
2-54

Table 2.1.3 Hypothetical Mean Frequency of Core Melt

(Based on Individual Hazard Curves)

Individual, Mean Frequency Ratio to

Hazard Curve of Core Melt 5.3-6 Value

Decoll1ement 2.1-5 4.0

Piedmont, Mmax:=6.3 1.5-5 2.9

Piedmont, Mmax'5.8 3.6-6 0.68

Northeast Tectonic 8.0-7 0.15

Crustal Block, MmX6 .0 4.1-7 0.08

Crustal Block, Mma X=5.5 1.0-7 0.02
2-55

2.2 FIRE

2.2.1 Deterministic Fire Growth Modeling

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 determined,

given an initial size fire. The deterministic 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 deterministic fire growth model. Briefly, this code is a synthesis of simplified,

quasi-steady unit models resulting in what is commonly 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 computer codes(3-7)

which use the unit-model approach to model compartment fire development. Of

parti cular 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 from 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 general thoughts are deemed warranted on the complexity

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, however, we cannot predict the initiating item's growth rate because basic

combustion mechanisms are not well understood. There are even questions and

doubts regarding the ability to predict the burning rate of a non-spreading,

hazardous scale fire in terms of basic measurable fuel properties., However,

until meaningful standard flammability tests and/or more sound scientific
2-56

predictions are developed, realistic "standardized" fire test procedures

should continue to be formulated for empirical measurements of the rate of

growth of isolated initiating items, the attendant fire plume, its development

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 COMPBRN) is

about the best that can be taken at the present time.

However, because fire modeling is still in a state of infancy, many

judgmental assumptions must be made in both modeling and physical data in

order to model fire development in the complex enclosures existing in nuclear

power plants. Additional complexity is introduced when one considers

electrical cable insulation as the fuel rather than the more commonly

considered. fuels such as wood or plastic slabs, which may have a more uniform

composition than cable insulation.

In fact, as discussed later, some of the models used in COMPBRN are nonphysical. That is, although these models usually lead to highly conservative

results, they do not adequately reflect the dependence on the physical

parameters which are evidenced in experimental data. Other models,

assumptions, and omissions in the application of COMPBRN to the Limerick SARA

are either conservative or nonconservative.

This combination of nonphysical models and conservative as well as nonconservative assumptions leads to very large uncertainties in the deterministic modeling process. It is therefore also difficult to quantify the effects

of these uncertainties on the probabilistic analysis, since the latter uses

the results of the deterministic analysi s as input. Indeed, as a general comment, one wonders whether more is gained by making gross judgmental assumptions, using them in an uncertain deterministic methodology and "cranking" the

results through a probabilistic analysis, than would be gained by making

direct j~udgments 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-the following sections. Section 2.2.1.2 briefly summarizes

our concerns with the deterministic modeling, whi'le Section 2.2.1.3 gives a
2-57

more detailed discussion of each item. Some suggestions for reducing the

uncertainties are given in Section 2.2.1.4.

2.2.1.2 Summary Evaluation of Deterministic Fire Growth Modeling

The deterministic methodology contained in the computer code

COMPBRN(1,2) is used in the Limerick SARA to evaluate the thermal hazards of

postulated fires in terms of heat flux, temperature, and fire growth. Thi s

code employs a unit-model approach which is acceptable given the current state

of the art in enclosure fire modeling as discussed in the previous section .

However, we find some of the submodels contained in the code to be nonphys ical

and some assumptions overconservative, while other assumptions and

applications yield nonconservative results. The uncertainties arising from

the combination of these counterbalancing models and assumptions are diffi cult

to quantify, but if forced to draw a conclusion we feel the deterministic

analysis 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 of a nonphysical, nonconservative model or assumption with

another non-physical model or assumption, no matter how conservative, leads to

a quantitatively useful result.

On the basis of our initial review of the deterministic fire modeling in

the Limerick SARA, we have identified the following items of concern, which

will be discussed in more detail in the next section.

The burning rate model is probably the most important source of uncertainty in the COMPBRN code. The methodology employed is not realis tic and can

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 fuel

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 ) It is

difficult to determine if the cable insulation burning rates obtained by this

method are conservative or nonconservative. For the postulated transientcombustible oil fire, the burning rate considered appears overconservative

with respect to that reported in the literature.(11)
2-58

Another example of nonphysical modeling is the fuel element ignition time

relationship. This model yields a finite fuel ignition time even if the incident heat flux is considerably below the critical value of 20 kW/m 2 found

necessary to initiate cable insulation damage in experiments.(12) The model

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 decrease as the target

fuel heats up. Cable damageability criteria based on a critical heat flux and

an accumulated energy, as discussed later and in Ref. 12, would be more

appropriate. The model used in COMPBRN leads to highly conservative cable

ignition times.

The model used to calculate the radiat ive heat transfer from the flame to

a target object is also overly conservative. The radiative heat flux obtained

from this model-is much greater than that o btained from a classical StefanBoltzmann model, wherein the heat flux is a function of the flame gas temperature to the fourth power. The COMPBRN model also neglected the attenuation of

the heat flux with distance due to intervening hot gas or smoke. The model

neglects, too, the partial reflection of the impinging radiative heat flux

from a target fuel element, as well as reradiation, convection, and other

1losses.

Additional conservatism is introduced by assum ptions 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 that adjacent cable raceways are separated from the initial fire by

the minimum-separation criteria specified for redundant safety-related cable

raceways (5 feet vertically and 3 feet horizontally). In other wordsý, 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 from the initial fire by more

than 20 feet were assumed to be damaged in a time interval equivalent to thp
2-59

damage time of a fire barrier taken as a 1-inch-thick ceramic-fiber blanket.

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,

those closer to the ceil~ing failing earlier than those below. Intermediate

growth stages between stages two and three might be appropriate.

Anothe r area of uncertainty concerns the quantity and size of the assumed

transient-combustible fires. The Limerick SARA assumes three possible

transient-combustible configurations; 2 pounds of paper 1 foot in diameter, 1

quart of solvent 0.5 foot 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 combinations 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, a 1-foot-diameter pool represents the most

severe hazard. A larger-diameter pool will give a larger heat release, although 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 imposition.

Some considerations omitted from the Limerick SARA would tend to make the

analysis nonconservative. These include the ef~fect~s 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 stratified

ceiling layer.

2.2.1.3 Detailed Evaluation of Deterministic Fire Growth Modeling

2.2.1.3.1 Fuel Burning Rate

The COMPBRN code~i) models the specific burning rate, m"', of the fuel,

which is equivalent to the mass loss rate in combustion, for fuel surface controlled fires as
2-60

The term rn" is defined as 'a speci~fic burning rate constant, and. the

second term 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".. adsraerrda "f Ir'.adsraerrda

tion, q os

in" (.q" iffI - (2.2) 0 f1 ,r

where.1. is the heat required to generate a unit mass of Vapor. Note that the

use of HF, the heat o -f combustion of the fuel, in Eq. (4.,4) of 'Ref.1, is

incorrect.' The correct formulation, is given by Eq.1(3) of Ref.13.

Note that if the externally applied heat flux', q*"ext' is zero, the

object will burn at a constant rate given by 4n"4"0. The consideration of

0as a 'constant for an element of fuel burning during the early growth

stages of a fire is questionable. For noncharri~ng combustibles, such as PMMA

or Plexigl-as, experimental data'in~dicate that mil"0 i~s 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 q offl ,r and ý"lossý* The 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 Eq. (2.1) with rn"0 and q*"ext constant by definition.

In COMPBRN, Eq. (2.1) is applied to ,each small squa~re "fuel element" into

which the individual cable trays (.super modules) have been discretized. The

fire is assumed to initiate in one element and spread'to, adjacent elements

when their ignition criteria are reached owing to the' incident -radiation from

the initial fire. A constant value of 11 =.0.002 kg/m 2 _sec is chosen for

each element. This methodology has a nonphysical result when the complete

ca~ble tray is considered, since the specific burning rate becomes a function

of the arbitrary number of elements into which the tray is divided.
2-61

For instance, if a fuel element was burning in infinite space with no

.externally applied heat flux, then according to Eq. (2.1) its burning rate

would be rntot-rno. However, if this fuel element is -divided into two

contiguous subelelements (1) and (2) with equal areas A/2 and with the flame

of subelement (1) supplying the external heat flux to subelement (2) and vice

versa, then, according to Eq. (2.1),

M"tot m rn"0 = [m"10 + Cs q".ext] ,(2.3)

where we have tacitly assumed that

11'

q ext,j = q ext,2 = q ext

Likewise, if the element were divided into n subelements with each j-th

element supplying an external heat flux to every other element, by defini tion

the progressive total burning rate when each of the j-subelements become

involved will not be equivalent to.the total burning rate if all' the

subelements had been involved initially. This indicates that care must be

exercised in using Eq. (2.1) to predict the ensuing development of a fire

along an individual cable tray.

Intermediate 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/mmn was

reached after about 37 minutes. This translates into a specific steady state

burning rate of 0.008 kg/m 2-sec. Use of such data and those of Ref.10 could

remove some of the uncertainty of the present model.

For transient combustibles, the fuel is not discretized and the speci fic

burning rate is assumed to be the const ,ant steady state value, rn"o Tabl e

D-4 of the Limerick SARA gives M"o value for paper and oil of about 0.061

kg/m2-sec'. It is believed that 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/m 2-sec.
2-62

2.2.1.3.2 Fuel Element Ignition

In the COMPBRN code, a' fuel element is considered ignited simply if its

surface temperature exceeds a critical ignition temperature, T*. Additionally, the fuel elements are modeled as semi- ,infinite slabs and the loss'es from

the fuel to the environment due to reradiation and convection are neglected.

An expression for the ignition time, t*, is obtained by solving the heat

conduction equat~ion, following page 75, Ref. 14, for the condition of a

constant imposed surface heat flux, q of

t* =(,r/4a)[k(T*-T0)/q"0J? (2.4)

This expression is physically incorrect since it implies that an ignition time will be reached no matter, how small a value of heat flux is

applied. Cable flammability test data(12) show that cables are generally

not damaged unless the heat flux is above a critical value of about 20 kW/m2

owing to heat losses at the surface..

Also, -the assumption of constant, imposed'heat flux i~s overly conservative since the heat flux received by an object is a function of the object

surface temperature, T5, which increases with time as the object is exposed

to the external flux.

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 h[Tp1 Ts] ,(2.5)

where Tpis the plume temperat .ureat the -cable height, TS i-s the cable

surface temperature, and h is the surface heat transfer coefficient.

Therefore, the surface heat flux will decrease-substantially as the
2-63

temperature of the cable surface approaches the plume temperature. The

COMPBRN code assumes that 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

estimated the plume temperature at 10 feet above the fire using three methods.

These include two correlations of convective heat flux by Alpert,(15416)

[one of which was used in COMPBRN(1] and a more recent plume correlation by

Stavrianidis.(17) The plume temperatures thus obtained range between 370 0 K

and 450'K. These low values indicate that cables within the convective plume

and located 10 feet above the fire would never reach their designated critical

ignition temperature of 840'K. This indicates the overconservativeness of

Limerick SARA which predicts cable ignition in 4 minutes for this target/fire

source configuration.

Of course, one must also consider the radiative heat transfer from the

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, yield

a radiative heat flux, 1" r, of 0.42 kW/m2. This is based upon use of the

following equation:

qr= (aTfl4/lr) (A P/02) ,(2.6)

where a is the Stefan-Boltzmann constant; Tf1 is the flame temperature

(1255-K)(17 ); x 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: x is equal to 5 ft;

and A p is the flames projected surface area. The emissivity, ,was assumed

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 Eq. 2.4) markedly higher than the 4 minutes stated in the Limerick SARA.
2-64

Even the radiative heat flux model ,as described in COMPBRN, yields a

value-of radiative heat flux lower than that required to achieve the critical

ignition temperature of 840*K within 4 minutes. In COMPBRN, the radiative

flux is given by

q F1 041l Qr/Afl ,(2.7)

where F0..fl i s the shape factor between the object and the fl ame, Afl i s

the flame surface. area, and Qr is the heat radiated by the fire which is

expressed as

'Y (2.8)

In the above expression, -Yreflects the radiant output fraction (-r=0.4 as

assumed in Ref. 1) and Q represents the total heat release rate of the fire.

To reconcile this wide disparity between ignition times reported and those

calculated by the methods' described above, "aback"l calculations were made using

Eq. 2.4 which indijcated that an imposed surface heat flux, q It0 9 of

approximately 12 kW/m2 is required to achieve a t* of roughly 4 minutes.

This value is obtainable using the COMPBRN model, if Afl in Eq. 2.7

represents the projected flame area (or pool area in this case) and not the

flame surface area. This is clearly inconsistent with the methodology used to

derive Eq. 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.

Even in 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, Eq.
2-65

(2.5), is used, once the surface temperature increases above the plume

temperature, heat will be convected away from the target reducing the effects

of radiation.

The selection of 840*K as the spontaneous ignition temperature for

EPR/Hypalon cable is also somewhat conservative since Table 3-1 of Ref. 9 presents experimental data showing that the critical temperature at or below

which ignition cannot be achieved is 893 0K for piloted ignition and is considerably higher for spontaneous ignition. Actually, as stated by Siu,(1)

the concept of a threshold ignition temperature is somewhat imprecise. Experimental data generally exhibit significant variations with further uncertainties arising if ill-defined cable insulation composi tions are involved.

The crucial issue is not whether the fuel surface reaches a certain temperature level , but whether the heat gains by the pyrolyzing gases are great

enough to overcome the losses and trigger the combustion reactions, and the

resulting heat of gaseous combustion is great enough to. sustain the reacti on.

Lee(12) has developed a set of cable damageability criteria along these

lines. For an applied heat flux, the time for spontaneous ignition is defined

in terms of a critical heat flux, q" cr, at or below which ignition cannot be

initiated and an accumulated energy, E, required for sustaining Ignition.

t = E/(q"ext - q cr) .(2.9)

Figure 2.2.1 (attached) shows test data(12) for the inverse of time to

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 ,

Eq. (2.4), using a critical spontaneous ignition temperature of 840*K. The

COMPBRN model is more conservative than even the piloted ignition data,

especially for low levels of external heat flux, i.e., a given external heat

flux will give an earlier time to ignition-than the data. Also, while the

data show no ignition below a heat flux of about 20 kW/m 2, the model predicts an ignition time for all values of heat flux. The 10-minute ignition

time for stage-two self-ignited cable raceway fires is indicated for

reference.
2-66

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 presence of walls will increase the gas

temperature at an elevation above the fire by a magnitude that can be

theoretical ly estimated by considering initiating fires having "equivalent"

heat release rates 2 and 4 times the actual heat release rate for walls and

corners, respectively. The neglect of this effect will have a nonconservative

effect on fire growth calculations,* especially in Fire Zone 2 where cabl e

trays are stacked against the "J" wall.

Evidence of the increased gas temperatures at a given elevation above a

fire is available in the literature. In Ref..16, Eqs. (3) and (4) illustrate

the concept of equivalent heat release rates mentioned above. Figure 6 of the

same reference shows test data of the fire positioni-ng effect s on ceiling

temperature. On page 119 of Ref. 19, the average plume temperature rise is

found to increase by factors of 1.75 and 2.5 for f ires adjacent to walls or

corners, respectively. Finally, Table A-i of Ref. 20 shows the upper-layer

gas temperature is likewise affected by burner locations near walls and

corners.

The increased gas temperatures in the presence of walls are due to the

effects of reduced cool air entrainment, which results in higher flames 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 consi ders complete combustion, which maximizes

the heat release rate and the temperatures near the fire, the wall effect

causes local temperature increases which must be considered to yield a

conservative result.

2.2.1.3.4 Stratified Ceiling Layer

The application of the COMPBRN code in the Limerick SARA failed to consider the stratified hot gas layer near the ceiling of enclosures even though
2-67

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

the size of the enclosure. However, in small fire zones, such as the static

inverter room, the hot gas layer near the ceiling could preheat the nonburning

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 affect the definition of fire growth stages in the Limerick SARA. It is conceivable that

unprotected cables near the ceiling., although horizontally separated by more

than 20 feet from an initiating fire, could ignite more quickly than a cable

closer than 20 feet but considerably below the ceiling. This-would tend to

have portions of fire growth stage 3 ahead of fire growth stage 2.

The ceiling gas layer model in COMPBRN is based on a simplified steady

gross heat balance. A uniform gas temperature is assumed throughout the upper

hot layer. Alpert(15) indicates that the ceiling gas temperature decreases

with distance from the ceiling, as well as with radial distance from the plume

axis. More recently, Newman and I-fill(21) have developed a transient correlation for the heat flux below the ceiling of an enclosure containing.a pool

fire, which incl-udes 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 Recommendations for Improving Fire Growth Modeling.

The previous sections have detailed some of our concerns. regarding the

sometimes nonphysical, usually overconservative, deterministic fire growth

modeling in the Limerick SARA. There are four major areas where we feel the

modeling can be made more realistic: the cable burning rate model, the fuel
2-68

element ignition time model, the flame radiant heat transfer model, and the

surface temperature dependence of the convective heat transfer model.

'Incorporation of recent test data(9,1O) 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 cable

ignition/damageabilit'y criteria,(12) based on a critical Iheat flux and an

accumulated energy, would yield cable ignition times more consistent wi th test

data. Improvement of the model for calculating the radiated heat flux received by a fuel element, by using ~an appropriate flame area and by considering attenuation due to hot gases and soot, will result in more realistic

fire growth scenarios and establish a more accurate 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. This will reduce the convective heat absorbed as the

object heats up and will1 allIow f or convecti ve cool ing if its temperature

exceeds that of the local fire plume.

2.2.2 Probabilistic Fire Analysis Review

*For the Limerick Generating Station (LGS) Unit 1, the Severe Acc ident Risk

Assessment (SARA) study reports that fire accident sequences constitute a significant portion of the overall1 public risk. In our review of the document,

we found no evidence contradicting this conclusion. However, our

understanding of the state of the art in fire-PRA, as well as the existing

inadequacies in both physical and probabil-istic modeling in this area,

precludes any judgment based on the quantitative results pr esented in the LGS

report. Further, the expected large uncer tainties associated with the

quantitative results would suggest that less importance be given to the

numbers. Hence, the scope of our review is twofold: first, to identify the

existing inadequacies in physical and probabilistic modeling in fire PRAs in

general; and, second, to review and comment on the existing LGS report for the

fire risk assessment.
2-69

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

sections which 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:

1. The approach taken for systematic identification of critical plant

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 estimating 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 nonconservative.

4. The LGS analysis appears to have identified all important safety components and cabling which are located in the crit ical fire areas, except for Zones 44 and 47.

5. The event trees for panel fi res generated by the LGS analysis should

be modi fied to take into account the layout of the panels with respect

to the critical portion of the zone.

6. The cumulative suppression distribution function generated in the LGS

report does not seem to agree with available data.

7. Suppression probabilistic modeling seems to be very conservative and

is not representative of the actual case.

8. The LGS analysis does not quantify the uncertainty of the final results. The uncertainty bounds generated are merely judgmental.
2-70

Consistent with these conclusions, the following section discusses each item

in detailI.

2.2.2.1 Evaluation of Significant Fire Frequencies in General Locations

In this part of the LGS analysis, the estimated frequencies of fires in

general locations were based on historical' fire occurrence data in NPPs. The

'gener~al locatio~ns 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 NPP~s. The point estimate frequencies

calculated for the general-locations seem to be reasonable, but the

uncertainty bounds were not determi~ned. The frequency of fires for the

individual fire zones was then Icalculated"!using the ratio of the weight of

combustible material contained within a zo ne to the total weight of

combustible material in the general location.- There is no justification for

using this ratio for estimating the specific zone fire frequency. However,

the results of these estimations were used, for the systematic identification

of critical fire zones through screeni~ng analysis, rather than the detailed

fire risk assessment.

For the detailed fire risk assessment,'the estimated fire occurrence

frequency within each zone was based on three different mechanisms of fi re

initiation: self-ignited cable fires, tra nsient combustible fires, and

distribution panel fires. Following are c omm ents regarding each type of fire

occurrence frequency estimation.

2.2.2.1.1 Self-Ignited Cable Fires

Three incidents of cable-raceway fires, have been reported in the data base

for NPPs. Two of them spread beyond one cable tray and were estimated to burn

for 30 minutes before being extinguished. The 'LGS. report i ndi cates that all

these cable fires were attributable to bad cable spl ices and underrated

ýcables. A review of the LGS data given in, Tables D-1 and D-2 of their
2-71

submittal suggest that incident 43 (Table D-1) was not caused by underrated

cables or bad splices. Hence, we cannot agree with the fivefold reduction of

self-ignited cable-raceway fire frequencies as indicated in the LGS report

based on the Limerick protection measures and flame retardant cables. It

appears to us that a threefold 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 the individual fire

zones, the frequency per reactor year was weighted according to the 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 weighting factor. In our view, the number of

conductors and splices, the. voltage/ power ratings, the geometric factors, etc.

may be more suitable for weighting the frequency of fire in each fire zone,

rather than simply the insulation weight. This indicates that large

uncertainties are present in the fire frequency estimates of various zones.

2.2.2.1.2 Transient-Combustible Fires

Three types of trans ient-combusti bl e fires were included in the analysis.

The quantity and the area of each type of transient combustible were conýsidered to be fixed. The state of the art for fire risk analysis is to consider 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 that no data are available, the

fr equency of fires for transient combustibles estimated in the LGS report

seems to be reasonable.

2.2.2.1.3 Power Distribution Panel Fires

The estimated frequency of fires occurring in power distribution panels

was based on five reported fires tha t occurred during 564 years of reviewed

U.S. LWR experience. The point estimate of fire frequency within a power

distribution panel was derived from these data and seems reasonable.
2-72

2.2..2.2 Screening Analysis

A systematic approach is used in the LGS report to identify 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 that zone will be disabled. Thecore-melt probabi~lity was then recalculated and multiplied by the frequency of

fire occurrence in that zone to provide a measure for screening analysis.

With this approach, the LGS fire analysis appears to have identified all1 the

critical areas in the plant. The quantitative reassessment of their results

are beyond the scope of this review. From our review of the FPER and the use

of engineering judgment, the critical fire areas identified by the LGS report

seem to be reasonable..

2.2.2.3. Probabilistic Modeling of Detection and Suppression

The probabilistic suppress ion/detect ion model used in the LGS study in the

form of a cumulative probability distribution to predict the probability of

failure to extinguish the fire within a speci fied time interval is'based on

actual plant data for automatic detection and manual suppression. It is

indicated that the data base for cable i nsulation fires reported by Fleming et

al.( 22) was used to construct the suppression probability distribution.

This document was reviewed and the cumulative suppression/detection was

reconstructed according to 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 2.2.2. Table 2.2.1 presents the data used by BNL. It is our

understanding that in the LGS estimate of the suppression success probability,

the self-extinguished cabinet fire incidents were included. In our opinion,

the LGS report should not take 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 distribution with t he assumption that the longest

suppression period is 1.3 hr (based on the longest suppression period observed

in the data base). We feel it is more appropriate to obtain a distribution
2-73

fit to the data rather than the "eyeball fitting" used by the LGS report. In

our analysis, the lognormal, 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 parametric Weibull dis tribution is the best

choice. A cumulative Weibull distribution F(x) can be defined by two

parameters, 7i and a , and is given by

F~x) = 1 -exp (-x/a)'1 . (2.10)

The estimated (a,r7) val'ues 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

modified LGS and the BNL curves is given in Figure 2.2. In the time interval

of 30 to 75 minutes, Curve I obtained by the Weibull fit to the LGS data i s

essentially the same as 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,

comparison of 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.

As in other conventional probabilistic risk assessments, the LGS report

assumes that fire growth and suppression are two independent processes, and

they are treated separately. This is one of the most important deficienci es

of existing fire risk analyses which usually results in very conservative

values for fir e-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

based on the data base which indicates that all five reported distribution

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 potential to propagate. In addition, a fivefold reduction was considered, based on
2-74

engineering judgment, to give credit to the IEEE 383 qualified flameretardant cable insulations. This reduction may not be justified. The combustibility of cable insulation can best be described through the sensitivity

of the cables to various thermal environme .nts, expressed as the change in

generation rate of combustible vapor per unit change in the flux received by

the combustible. This value, usually denoted by "S", is 0.17(g/kj) for EPR/Hypalon and 0.22 (g/kj) for PE/PVC cable insulation.(23,10) Hence, a

maximum factor of 2 may be credited because of flame-retardant cable

insulations.

Additionally, during a visit-to the plant, it was noted that some of the

panels are airtight. For these panels, we. feel the probability of fire

propagation is negligible and, 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 nonconservative. In general, we do not expect the impact of

panel fires to change appreciably if more detailed analyses were performed.

2.2.2.4 Probabilistic Modeling of Plant Damage State

Generally, three stages of fire growth and corresponding states of shutdown equipment damage were evaluated in the~analysis. The first stage considered is damage to components in the immediate vicinity of the source of

fire. The second stage is fire growth to adjacent unprotected cable raceways

separated from the initial fire by minimum separation criteria (5 ft vertically and 3 ft horizontally). The third stage of fire growth represents fire of

sufficient severity and duration to damage the mutually redundant shutdown

methods which may have cabling with a separation distance of at least- 20 feet

or protected by fire barriers. Certain inherent assumptions in the analysis

are as follows:

1. The rate of fire growth is not dependent on the suppression.

2. A 20-ft separation is considered t~o be equivalent to a 1/2-hour fire

barrier (1-in- thick cerami-c blanket).
2-75

3. Cable raceways separated from the fire source by 40 ft or more were

considered undamaged by the fire.

4. It was assumed that long-term heat removal systems not required until

20 hours into the fire-induced transient could be recovered by operati ng val ves .manuall1y and operati ng pumps l ocal ly. -The probabiIi ty of

failure by the operator to perform these recovery actions was considered 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 owing to lack of information and t ime

limitations. However, on the basis of a limited identification of various

critical components and systems in different fire zones by means of the

information gathered from LGS-FPER and the plant visit, we concluded that i n

most cases the LGS report identified the components properly. There are two

exceptions as follows:

1. In Zone 44, BNL has identified seven distribution panels and motor

control centers. These are distribution panels 10D201, 100202, 10D203

and motor control centers 10B211, 10B212, 10BV215 and 10B216. We have

also concluded that a fire in distribution panels 10D202 and 10D203

would affect the operation of the HPCIS, and a fire in distribution

panel 10D201 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 and only two of them are critical

(10D201 and 100203).

2. During the plant visit, a booster fuel pool cooling pump was noted in

Zone 47, General Equip-ment Area, pump in the vicinity of the

northeast corner, which is the critical area in this zone. This pump
2-76

was not identified in the LGS report. -Therefore, its potential for

intitiation and progression 'of fire'adversely affecting the cables in

this area was not considered.

Before presenting our comments on each' critical fire. zone, a. further

discussion of the inherent assumptions used in the LGS report mentioned

earlier in this section is appropriate, mo re s pecifically, the nature of the

interaction between fire growth and suppression activities. In the LGS report,

it was assumed that a fire can progress regardless of suppression initiation,

but terminates with some probability after an expected time which is required

for successful suppression. The lack of physical modeling for -the suppression

phase of a fire scenario appears to be one of. the weakest links in the

analysis. We are aware of this deficiency in other fire PRAs and it seems to

be a conventional practice, usually resultfing in very conservative estimates

for fire impact on equipment and cabling. Wh il e reevaluation of the results

given in the LGS report,' taking into account proper detection and suppression

modeling, is beyond the scope of this review,,.it seems necessary to discuss

the basis for such analysis.

In the analysis of a fire scenario, initiation time for detection and suppression 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

response, some design charts are available which graphically, or through some

equations, determine the response time vs the spacing, ceiling height, and

heat, release rate.(24 -

26) If detailed fire growth m odeling, with the

associated uncertainties of various fire parameters, is available for a

specific scenario, the detection and suppression re~sponse 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) as defined in Ref. (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
2-77

may be estimated by means of available data on response time during fire

drills and some engineering judgment. 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 combination of simplistic models,

coupled with empirical correlations, may be used. For example, the effect of

sprinkler systems on fire growth may simply be modeled in the form of global

energy balance.(27)

In conclusion, the time in which fire can reach 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 LGS report, it is conservatively assumed that probabilities of various stages of

growth can be determined using the time period for the completion of successful suppression, rather than the initiation of suppression. 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-Specific Commients.

In addition to the generic comments made in previous sections, there are

additional zone-specific comments that may affect the results of the fire PRAs

given in the LGS report. These comments, mostly concerning the layout of

different components in various critical zones, are based on the review of the

FPER and the plant visit.

a. Zone 44, Safeguard Access Area (CH=36 ft, A=8930 ft2,3 ASD=357.2,

S=M).* In this zone, there are a total of seven motor control centers

(MCC) and distribution panels. Four of these panels are located close

to the critical corners. These are distribution panels 10D202 in SW,.

10D203 in NE, 10D201 in SW, and MCC-10B21i in SW (Drawing M118, Rev.).

The event tree associated with the panel fires should be modified.

*CH is the ceiling height, A is the floor area, and ASD is the area per smoke

detector. The "S=M" represents manual suppression, where "S=A" represents

automatic suppression.
2-78

b. Zone 45, CRD Hydraulic Equipment Area (CH=25 ft, A=12860 ft2.

ASD=676.-8 ft, S=M/A). The only critical panel which is located in the

NE corner is the MCC-10B224. The other panels are not located in the

vicinity of the NE corner (Drawing M119, Rev. 19). The event tree associated with the panel fires should be modified.

c. Zone 47, General Equipment Area (CH~not available, A=9800 ft2,3

ASD=490 ft, S=M/A). According to the drawing M120, Rev. 18, none of

the distribution panels, load centers, or motor control centers are

located in the vicinity of the critical NE corner. Therefore, the event tree associated with panel fires in this zone should be modified.

The only component located in the NE corner of this zone that may result in a fire hazard is a booster fuel pool cooling pump.

2.2.3 References to Section 2.2

1. Si u, N.0. , "Probabi1i sti c Model s f or the Behavi or of Compartment Fi res,"

School of Engineering and Applied Science, University of California, Los

Angeles, Ca., NUREG/CR-2269, August 1981.

2. Siu, N.O., "COMPBRN - A Computer Code for Modeling Compartment Fires,"

School of Engineering and Applied Science, University of California, Los

Angeles, Ca., UCLA-ENG-8257, August 1982.

3. Mitler, Henri E. and Emmons, Howard W., "Documentation for CFCV, the

Fifth Harvard Computer Fire Code," Harvard Universityý Cambridge, Ma.,

October 1981.

4. Quintiere, J.G., "Growth of Fire in Building Compartments," ASTM Special

Technical Pub. 614, 1977.

5. Tatem, P.A., et al, "Liquid Pool Fires in a Complete Enclosure," 1982

Technical Meeting., the Eastern Section of the Combustion Institute,

Atlantic City, N.J., December 14-16, 1982.
2-79

6. Zukowski, E.E. and Kubota, T., "Two Layer Modeling of Smoke Movement i n

Building Fires," Fire and Material, 4, 17, 1980.

7. Delichatsios, M.A., et al., "Computer Modeling of Aircraft Cabin Fire

Phenomena," FMRC J.1. OGON1.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, Intermediate-Scale

Fire Tests of Cable Tray Installations," EPRI NP-1881, Electric Power Research Institute, Palo Alto, Ca., August 1982.

10. Tewarson, A., Lee, J.L, and Pion, R.F., "Categorization of CableFlammability, Part 1: Laboratory Evaluation of Cable Flammability Parameters," EPRI* NP-1200, Electric Power Research Institute, Palo Alto, Ca.,

October 1979.

11. Tewarson, A., "Fire Behavior of Transformer Dielectric Insulating

Fluids," DOT-TSG-1703, prepared for U.S. Dept. of Transportation;- Transportation Systems Center, by Factory Mutual Research Corp., Norwood, 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. Tewarson, A., "Physico-Chemical and Combustion/Pyrolysis of Polymeric

Materials," NBS-GCR-80-295, prepared for U.S. Dept. of Commerce, National

Bureau of Standards, Center for Fire Research by Factory Mutual Research

Corp., Norwood, Ma., November 1980.

14. Carslaw, H.S. and Jaeger, J.C., "Conduction of Heat in Solids, 2nd Ed.,

Oxford Clarendon Press, 1959.
2-80

15. Al pert, R. L. , "Cal cul ati on of Response Time of CeiIi ng-Mounted Fi re Detectors," Fire Technology, Vol. 8, 197.2,,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.,

August 1982.

17. Stavrianidis, P., "The Behavior of Plumes Above Pool Fires," a thesis

presented to the Faculty of the Department of Mechanical Engineeri ng of

Northeastern University, Boston, Ma., August 1980.

18. Orloff, L., "Simplified Radiation Modeling of Pool Fires," FMRC J. I. No.

OE1E0.BU-1, Factory Mutual Research Corp., Norwood, Ma., April 1980.

19. Zukoski , E.E. , Kubata, T. , and Cetegen, B. , "Entrainment in Fire Pl umes,"

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,. NBSIR 82-2520,

September 1982.

21. Newman, J.S. and Hill , J.P. , "Assessment of Exposure Fire Hazards to

Cable Trays," EPRI-NP-1675, Electric Power Research Institute, Pal o Alto,

Ca., January 1981.

22. Fleming, K., Houghton, W.J., and Scaletta, F.P., "'A Methodology for Risk

Assessment of M ,ajor F .ires and its Application-to an HTGR'Plant,,"

GA-A15402, General Atomic Company, San Diego, Ca., 1979.

23. Tewarson, A., "Damageability and Combustibility of Electrical Cables,"

paper presented at FMRC/EPRI Seminar, Factory Mutual Conference Center,

Norwood, Ma., December 1981.

24. Benjamin, I., et al,, "An Analysis of the Report on Environments of Fire

Detectors," Ad Hoc Committee of the Fire Detection Institute, 1979.
25. Newman, J.S. , "Fire Tests in Ventilated Rooms - Detection of Cable Tray

and Exposure Fires," EPRI 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.
2-82

0 10 20 30 40 50 60 70 80

EXTERNAL HEAT FLUX (kW/m2

Piloted ig~nition of EPR/Hypalon cable under various. external

heat flux.

Figure 2.2.1
2-83

Ld

U-I

z .0

U)

D

LL

0

x

-J

Z-)0 15 30 45 80 75 9

TIME AVAILABLE TO SUPPRESS FIRE (minutes)

0

Figure 2.2.2 Fire suppression model.
2-84

Table 2.2.1 Suppression Data and Calculations Performed

for Suppression Success Probability

Self-Ignited Cable-Raceway Fires

Index* Plant Name Time to Bring Fire

Under Control (hr)

Type of

Detection

Type of

Suppressi on

58**

23

25

24

8

28

37

42

46

27

Browns Ferry

Zion .2

San Onofre 1

San Onofre 1

Kewa unee

Three Mile Is. 2

Vermont Yankee

Nine Mile Pt. 1

Oyster Creek

Trojan

7.0

1..3

0.7

0.5

0.5

0.5

0.5

0.05

0.05

0.05

Automatic.

Manual/Automatic

Manual

Manual

Automati c/Manual

Manual

Automatic

Manual

Manual

Manual

Manual

Manual /Automatic

Manual

Manual

Automat i c/ Manual

Manual

Manual

Manual

Manual

Manual

*Indices are the same as those in Fleming's. report.(22 )

**The fire occurrences during the construction phase or those that were selfextinguished and confined to a cabinet were not included. In addition, the

Browns Ferry fire indicated above is not included in our analysis.
3-1

3.0 ACCIDENT SEQUENCE ANALYSIS

3.1 Seismic

The objectives of this section are to provide (i) a brief description of

the methodology and assumptions adopted in the LGS-SARA(1 report in the

quantification of seismic accident sequences, and (ii) a BNL review comments

of particular 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 the modifications. This section is divided into

two parts. Section 3.1.1 addresses those plant frontline systems which are

identified in the LGS-SARA report and the method of quantification by which

system unavailabilities, including the seismic contributions, are evaluated.

Section 3.1.2 summnarizes the sei~smic event tree approach and the seismic

accident sequence analysis.

3.1.1 Plant Frontline Systems.

This section comprises two subsections. Subsection 3.1.1.1 presents 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 unavailabil-ity. 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 is also included.

3.1.1.1 Overview of the SARA Approach in Frontline System Modeling

The system analysis part of the LGS-SARA effort is based extensively on

the structure and contents of the LGS-PRA.(2) This includes use of 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

that appear in the LGS-PRA system fault trees constitute, in part, the group

of components for which fragility evaluations were conducted.
3-2

LGS-SARA purported to have examined the fragility of two groups of components:.those contained in the LGS-PRA system fault trees and those

identified as having the potential of significantly influencing the likelihood

of core damage from seismic events, such as the reactor vessel and other

related structures. A detailed discussion of component fragility is presented

in Chapter 2.1. These components are then ranked according to the

acceleration capacity of each item; those with a median ground acceleration

capacity greater than 1.56 g were not considered~since they are deemed to have

a far higher ground acceleration capacity than those predicted for the reactor

site. On the basis of this criterion, a fi 'nal list of 17 components are

selected for use in the LGS-SARA evaluation, Table 3.1.1.

Each seismic frontline system fault tree developed in the LGS-SARA analysis is made up of two parts: the first part, which leads to the failure of

the system, consists of the random independent failures evaluated in the

LGS-PRA; the second part includes all the pertinent seismic-related failures

as determined using a specified criterion. This criterion for inclusion as a

seismic-relate'd failure requires that the component appears in Table 3.1.1.

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 system, failure of the condensate storage tank (CST) necessitates 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 the 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 removal (RHR), s tandby liquid con 'trol (SLC) ,'automatic depressurization system

(ADS), and emergency power. An example of HPCI seismic system fault tree is

given in Figure 3.1.1.
3-3

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 focusing solely on seismicrelated failure events. In some circumstances, these random independent

failures when coupled with a substantial. reduction in the operator's abil ity

to follow procedures due to high stress conditions resulting from an

earthquake may contribute significantly to core damage.

BNL reviewed the LGS-SARA modularized system fault trees developed for

the seismic analysis. These fault trees were based on the list of 17

components identified as more susceptible to seismic-event-related failure. On

p. 3-1 of the LGS-SARA, it is stated th at the internal system fault trees

provide, in part, the list of components for which fragility functions we re

developed and that additional items were included when they were deemed to

have the potential for significantly influencing the likelihood of core damage

from seismic events. BNL agrees that consideration of only those components

identified in the internal event system fault trees does not ensure inclu sion

of all important seismic-sensitive components, since in the construction of

the internal event system fault trees, depending on the level of detail in the

development of the trees, approximations may hav .e 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, when it is used in the seismic assessment, dependence on pi ping

failure would not have been properly evaluated. It is not clear from the

report that a systematic search was conducted to identify and select

components for fragility evaluation to ensure that all components sensitve to

seismic events are included in the analysis.

In the modularized system fault tree approach, intersystem and support

system dependences are not explicitly modeled. LGS-SARA did include the common mode failure of the diesel generators' as a means of failing the systems.
3-4

Preliminary review of the fault trees appea~rs to i~ndicate that common mode

diesel failu~re is one of the dominant scenarios leading to core damage. The

BNL review of the Limerick internal event r~eport(7) assessed that

contributions from inclusion of 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 insi gnificant.

In addition to those dependences discussed earlier, one dependence

involves failure to transfer water source f rom CST to suppression pool. This

operation is required for both 'the, HPCI andý the RCIC systems whenever there is

a low CST level. An operator failure to transfer, given a low CST level, is

likely to affect both high pressure systems!. This dependence should be

included to properly reflect its impact on the final results.

Electric Power

The failure of the electric 'power systiem is modeled* with the failures of

seven components; namely, two faults leading to the loss of the 440-V power

supply, three faults resulting in t~he loss 'of the diesel generators, one leading to losing the 4-ky bus, and one to loss of dc power.

For'both the HPCI and RCIC systems, l oss of control power due to failure

of the dc bus is assumed to disable the systems. 'In principle, it is possible to operate the two high pressure systems in a total blackout condition'

for an extended period of time with the operator manually providing the control s necessary. Nevertheless, in the evenlt of 'an earthquake, BNL concurs

that the LGA-SARA assumption may be more re~alistic.

In the LGS-SARA A .ppendix B, it is esti'mated that at 1.0 g no significant

damage to the diesel generator fuel oil tanks is expected, and at accelerations somewhat in excess of 1.0 g, failure of attachments Would btlikely.

It Iwas 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
3-5

core-damage frequency. The data reported in the Ind'ian Point Safety Study 4

for the diesel generator oil tank are 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 the diesel generator fuel oil tank to show that they have the

capacity much greater than 1.0 g to justify their exclusion from the system

fault trees.

Human Error

In the seismic part of LGS-SARA, it was reported that in estimating the

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 limited to a maximum probability of 1.0. The factor of 10 is based on

the fact that an earthquake sufficiently intense to damage reactor systems

will initially disturb the performance of the operators and raise doubts in_

their minds about'the performance of instrumentation and controls. The

earthquake may also lead to component failures not normally encountered in

plant operations and, therefore, may require innovative actions on the part of

the operators.

It is BNL's judgment'that during and subsequent to an earthquake, the

operators' ability to follow procedures, to diagnose-problems, or to take corrective actions depends on the intensity of the earthquake. Given the limnited

information available in this area, it is often difficult to quantify the

likelihood of failure under these unusual circumstances. However, one would

expect that an increase in the human failure probability is warranted.

Moreover, there are three factors which are also important in determi fling

the human failure probability. One of them is the avail-ablity of reliable

instrumentation. Subsequent to an earthquake, with alarms and annunciators sounding, it may be difficult for an operator to adequately assess

the plants true condition, since some instrumentation may give erroneous

information. This is a much more challenging situat ion which significantl y
3-6

increases the complexities confronting the operator. Two types of human

failure may result, in addition to those normally considered in the LGS-PRA:

1) the operator may be misled by false instrument readings, and follow the

wrong procedure in securing the plant; or 2) the operator may be misled by

wrong and confusing information into an error of commission.

The second contribution to human'failure that was not addressed in

LGS-SARA is the effect of aftershock upon the ability of the operator to

discharge his responsibility. On the basis of seismic data, the probability

of occurrence of aftershock decreases following an exponential 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 within the time frame when

operator action is critical , it may further impair his ability to respond to'

the-.demands of the plant.

The third area entails the subject of display instrumentation, which is

intended to provide the operator-with pertinent information to help him to

understand the status of the plant. Display instrumentation could be in the

form of lights, chart recorder, annunciators, alarms, etc. In the event of an

earthquake or an aftershock, the failure modes- of the display instrumentation

could be: 1) display. information inconsistent with other indicators, 2) loss

of display function. LGS-SARA should furnish. a discussion on this subject to

ensure that failure of display instrumentation has been investigated and i s

deemed to have no significant impact on final results.

BNL concludes that the increase in the human failure probability by a

factor of 10 may be reasonable in some instances, whereas it may be

conservative or nonconservative in Others depending on the situat~ion. An

example of how the absence of readily available and reliable information can

affect the operator's ability to pursue the proper actions i~s given for the

CST. The CST is calculated to have a median ground acceleration capacity of

0.24 g, which is comparatively low in light of the other component values. It

constitutes one of the two water sources from which the HPCI and the RCIC take
3-7

suction. Failure of the CST would necessitate a transfer of the suction from

the CST to the suppression pool. As for the HPCI, this transfer process is

automatic, i.e., given that there is a low CST tank level, an automatic

switchover will be initiated; however, for the RCIC, this transfer is a manual

operation. The failure mode of the CST water level sensors, given that t he

CST is failed, was not addressed in LGS-SARA. Nevertheless, one could

postulate the following: 1) that despite the failure of the CST, whether it

be ruptured or toppled over, the level sensors give a low level reading; 2)

that in the failure of the CST, the level sensors-are damaged and erroneous or

misleading information results. Preclusion of one or the other would require

a more detailed investigation of the failure modes of both the CST and the

level sensors.

The occurrance of scenario 2 implies that the information given to the

operator is misleading, and hence the failure probability for the operator to

respond properly should be close to unity rather than based on an arbitrary

rule of thumb - a factor of 10. It so happens that when this factor of 10 is

applied to the HPCI and the RCIC transfer from CST to suppression pool, t he

human failure probability is unity. But there are other human operations

within these system fault 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 restart system, fail ure

to transfer service water, etc. A detailed discussion of this impact upon

system unavailabilities is deferred to the next section.

Finally, it is important to note that LGS-SARA did not convey to the reviewers that the increase in human error was applied consistently to all the

pertinent basic human events. BNL reviewed the LGS-PRA system fault tree s and

-identified a number of manual operations which are omitted in the seismic systemn 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 is needed, and pertinent findings on

manual errors should be included in the modularized system fault trees.
3-8

Relay Chatter

It is reported in LGS-SARA that low accelerations cause a momentary interruption of control circuits and power supplies (typically from relay-contact

chatter); however, relay chatter i's dismissed as a means of leading to system

failure since the operator can intervene and reset the c ircuit, and hence

restore the system to its initial state.

It appears that the question here is not whether the relays will chatter

or at what acceleration they will1 begin to chatter, but what credit should begiven to the operator to 'reset them if relay chatter occurs. If in' one part

of LGS-SARA, it is maintained that in the event of an earthquake, human error

should be modified by a factor of 10 to reflect increased stress, it seems

only consistent that these human response s to reset relays be treated

similarly in assessing their failures. BN L is of the opinion that if there is

relay chatter, failure on the part of the operator to reset would result in

the equivalent of a relay failure.

If one wants to quantify the impact of relay chatter upon the system

failure, then one would have to ascertain relay fragility information for the

various kinds of relays., The Indian Point study (4) states that relay

chatter occurs at 1.2 g *and presents no major difficulty.. The SSMRP data(5)

show that chatter occurs at as low as 0.75 g (spectral acceleration).

Moreover, for certain relay chatter which results in a breaker trip, reset of

the systent may be readily possible at the control room; however, some relay

trips may require rese tting at local panels which substantially increases 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 human action required to reset relay, which leads

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
3-9

been trained and for which no procedure has been written, what is the probability 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

operator can indeed reset the relays in a reasonable time and restore the

system.

An example to illustrate these points can be found in the SLC f ault tree.

There are two relays per SLC pump, for example, K4A and K5A for train A, K4Band K5B for train B, etc. If chatter causes these relays to terminate the

operation of the three SLC pumps, this will lead to a direct failure of the

SLC system. Furthermore, in the redundant reactivity control system, relay

chatter may cause all APRM channels to fail, which in turn will result in

failure to initiate the SLC explosive valves and the SLC pumps. In an ATWS

accident event, the time available to an operator to respond to these

challenges is also significantly reduced to the order of minutes. In light of

this information, the impact of relay chatter upon the SLC system should be

evaluated in more detail.

Transients

A list of the LGS-SARA mean random failure values and the nomenclature is

given in Table 3.1.2; the first column of values are those given in the LGSSARA report; The second column tabulates the values used in the internal event 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 should be used in the LGS -SARA

study. Differences between the first and second columns are quite obviou s.

Despite the fact that few explanations are furnished in LGS-SARA to address

the differences for both high pressure systems, these differences are

miniscule. But for the low pressure system (V) and the manual

depressurization (X) function, a more detailed discussion is warranted.

As stated repeatedly in LGS-SARA, the seismic evaluation was based

extensively on the LGS-PRA; therefore, it is reasonable to assume, unless
3-10

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

Automatic Depressurization System (ADS). The low pressure injection function

(V) represents either the failure of the ADS hardware or a simultaneous

failure of the LPCI and the LPCS systems. In the LGS-PRA, the X function

unavailability is calculated to be 2 x 10-3. the V function value estimated

by BNL is based on the LGS-PRA unavailability of the LPCI and LPCS systems

given that there is a loss of offsite power and a failure to recover offsite

power to be 2.65 x 10- 4. According to the information provided on p. C-15

of Table C-6 of the LGS-SARA, it appears that the V function defined in the

report consists only of the [PCI and the LPCS systems; this notion is further

confirmed in the Boolean expression of X = XR+A shown on p. C-14 of Table

C-5. XR 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 depressurize the reactor does not require electric control or motive power, it i s

possible to argue that the hardware failure of the ADS is lumped with the X

function without much impact on the function unavailability. However, this is

not consistent with what has been presented in the [65-PRA, and may result in

misleading conclusions of dominant sequences. The impact of properly

including -the ADS hardware failure within the V function for various accidentsequences 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 NUREG/CR-3028. The most substantial increase between the

LGS-SARA and. NUREG/CR-3028 internal event values occurs with the V functiona factor of 3.7 followed by a factor of 3.0 increase for the X function.

The common mode diesel generator failure probability of 1.88x10-3 was

reported in earlier revisions of the LGS-PRA, and that this value was used in

the NUREG/CR-3028. Subsequent revisions to LGS-PRA modified the'

unavailability to 1.08x10-3 . claiming that the earlier version was a

typographical error. In LGS-SARA, a diesel generator common mode failure mean
3-11

value of 1.25x10-3 was reported. BNL agreed that the 1.88x10-3 value is

overly conservative. Recently, studies(6) to better evaluate the diesel

common mode una vailability have suggested values below 1.0x10-3. In this

review, BNL will use the 1.25x10-3 for comparison purposes.

The increase in numerical values for the HINIA and RIN3 is due to the

following: HINIA and RIN3 represent fail'ure to provide flow from the

suppression pool , given that the CST water is unavailable. The major

.difference between the two events lies in the manual action required to

perform the operation for the RCIC system, FSAR, p. 7.48. Consequently, if a

factor of 10 increase is assumed, the manual error for failure to transfer

becomes unity and dominates the failure of the RCIC system. Because of the

automatic transfer function in the HPCI,.a similar increase in the manual

error results only in minimal increase in the system unavailability. If,

instead, a human factor of 7.5 is used, the RIN3 will be 0.75, whereas, HINIA

would remain unchanged.

Another major change that is evident if the factor of .10 increase is used

in the manual depressurization function. This increase results merely from

applying the human error factor of 10 to the NUREG/CR-3028 value of 6x10-3 .

It appears that for HINIA, RIN3 and X, the increase due to the human

error factor was not included in their values as it'should be.

Anticipated Transients Without Scram

In the event of an earthquake resulting in an ATWS, LGS-SARA analyzed the

sequence using a loss of offsite power ATWS event tree. A set of mean failure

values that was used in the LGS-SARA analysis is shown on Table 3.1-3. The

first three columns in the table present values used in the LGS-SARA study ,

the LGS-PRA, and the BNL internal' 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 pointed out that these values are

representative numbers; one shoul~d refer to the report-s indicated for more

detailed information.
3-12

LGS-SARA values for both the HPCI and RCIC failure values are in general

lower than those of the.LGS-PRA and NUREG/CR-3028. The increase for RCIC,

RR, is about a factor of 6.6 times. As for the ADS inhibit function,' the

LGS-SARA value is 8.0x10 3 vs 2.0x10-2 from NUREG/CR-3028; another factor

of 10 increase due to the intense stress level for the operator brings the

final value.(last column) to 2.0x10-1. There is~no disagreement on the

values-of UH as assessed by BNL and LGS-SA.RA. Little increase is noted

between the LGS-SARA and BNL values for the SLC system; however, the LGS-SARA

value is about a factor of 10 larger than the LGS-PRA value. W2 ,was

reported in LGS-SARA to be 0.1 rather than the 0.14 used. in the LGS-PRA. 'The

diesel generator common mode failure, HINIA, and RIN3 fail~ure values are

described in the previous paragraphs..

The value selected for the mechanical failure of the scram system

,increased to 1.5x1Q-5. The variabl'e'PCR is defined in the text of LGSSARA to have a value of 0.2, but no description of PCR is provided. Fail1ure

to scram is defined in LGS-SARA as

C (1CR)ý CR.+ PCR (S3 + S5 +:S7),

A telephone conversation with PECo revealed that' the PCR is a

judgmental factor applied to the seismic failure of the reactor internal s and

CRD guide tubes.,(see Figure 3.1.2,). PECo stated that failure of the CR0 guide

tubes or the reactor internals due to an.elarthquake would cause a failure to

scram only 20 % of the time. Since information on how this PCR value is

obtained is incomplete, it is difficult for BNL to judge-its validity.

Another area of concern is in the treatment' of random failure to scram;

BNL believes that, if there is a challenge to the. scram, system, the failure to

scram proba~bility should not be we ighted by a factor of (1-PCR)= 0.8. Also

in the telephone conversation with PECo, it was explained that they attempted

to preserve the scram failure Probability from the.0.8 reduction by increasing

the scram failure probability from 1.0x1075 -to 1.5x10-5. It is suggested

that a more detailed documentation of Ithese points by PECo be provided i n

LGS-SARA.
3-13

For the purposes of sequence quantification to be presented in the next

section, failure to scram is defined by BNL as follows:

CM = CR + S3 + S5 + S7.

Finally, it is suggested that a detailed discussion be provided in

LGS-SARA to identify and reconcile differences in the random failure values

used in LGS-SARA and LGS-PRA.

3.1.2 Accident Sequence Analysis

This section addresses the definition of accident sequences and the

quantification of core damage probability in the event of an earthquake.

Section 3.1.2.1 briefly describes the approach and methodology used in

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 Sequence Analysis

LGS-SARA examined various fragility estimates (provided in Appendix B)

and concluded that the offsite power system was most susceptible to an

earthquake which, when failed, would result in an initiating event. Failure

of pipes and valves causing an initiating event is dismissed as highly improbable in light of the significantly greater capacities of these components.

For this reason LGS-SARA maintains that the frequency of a seismically induced

LOCA. (large, medium, or small) is insignificant. The simultaneous occurrence

of an earthquake and a random LOCA event is also estimated to be smaller by a

few orders of magnitude than the loss-of-off site 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 success or failure of a number of critical functions whose operation or inopera-
3-14

tion greatly affects the analysis to be followed '(see Figure 3.1.3). Thi s

tree is made up of five functions, namely, the seismic-event-initiating

f requency, reactor' pressure vessel , reactor, and control buil1di ng, .and reactor

scram. Failure of the reactor pressure vessel. due to an earthquake leads

directly to core damage., The failure was identified to be initially the

failure of the vessel supports which, in turn, results in of al~l four steam

pipes being severed. To mitigate such a breach of the reactor-coolant

boundary is far beyond the, capability of the EGGS.

If the reactor pressure vessel stays intact, failure of the reactor and

control building will result in core da mage 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 Figures

3.1.4 and 3.1.5, respectively.

The event tree presented in Figure 3.1.4 is identical in structure to

that of the internal loss-of-offsite power event. Systems which are required to mitigate the event are assessed and accident sequences are defined.

In Figure 3.1.5, the mitigation of an ATWS event is presented. Its

structure is again identical to the one given i~n the LGS-PRA for loss of

off~site power.

Inputs to these event trees for individual systems are based upon the

modul ari zed system f aul t 'trees,, and- a di scuss'ion of these trees i s provi ded i n

Section 3.1.1. Quantification of these event trees was performed using the

computer code SEISMIC. ,the'Mon-te* 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 on the basis of the Boolean

expression inputed for that par ticular sequence. Median and mean values, and

confidence levels of the sequences are also evaluated and those for the

dominant sequences are reported.
3-15

3.1.2.2 BNL Review of Accident Sequence Quantification

BNL reviewed the event trees and assumptions which enter into the development 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 revisions made to the

modularized system fault trees, estimates of their impacts on respective

accident sequence core-damage frequencies are described.

Methodol ogy

The event tree - fault tree methodology employed in the LGS-SARA

represents a widely practiced approach used within the nuclear industry today

to assess accident sequences and core-damage frequencies. BNL agrees that it

is adequate in evaluating risk indices within the context and requirements of

today's risk assessment studies.

The LGS-SARA analysis is based extensively on the approach and results

of the LGS-PRA. Two event trees from the LGS-PRA were adopted to analyze the

seismic-initiating event. They are the transient and ATWS loss-of-offsite

power trees. While BNL agrees that these trees will model the loss-of-o -ffsite

power event adequately if caution is exercised in tiddressing the dependent

fail'ure of components due to an earthquake, additional information should be

included in LGS-SARA to establish why the seismic event evaluation can be

based extensively on the internal event analysis. In other words, it shoul d

be shown that external event accidents do not warrant separate event trees to

model the different scenarios. The rationale on why the LGS-PRA event trees

were used should reflect these concerns.

Initiating Events

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 cerami c

insulators (median ground acceleration capacity of 0.20 g) was identified to
3-16

be the major initiating event contributor. Failure of the reactor and control

building and of the 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

members. In particular, it-is not obvious that effort was devoted to examining the non-safety-related equipment or equipment not important for a safe

shutdown of the plant to determine if they could become initiating-event

contributors in an earthquake. These two types of equipment .are not subjected

to the same rigorous seismic qualification standards as other seismically

qualified components. Depending on the capacities of these non-safety

components, an earthquake with low ground accelerations might cause a reactor

trip-without failing the switchyard ceramic insulators. Such an event will

initiate a transient which should be evaluated by event trees similar to those

pr esented in Figures 3.1.4 and 3..1.5. The difference between the event trees

is that there is offsite power in this case. In the event that a transient

does not occur given an earthquake, then the sequence is a success eve nt.

For example, the feedwater system is not required' for a safe shutdown of

the plant nor is it safety related; however, if an earthquake occurs, control,

relays, and other components of the feedwater system may generate a trip of

the system which will result in a reactor transient.

In Figure 3.1.3, the event TS, sequence n 'umber 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

adequately accounted for in the turbine-trip-initiating event.

If, in an earthquake, 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 the turbine-trip
3-17

event tree. However, if there is failure of non-safety equipment or trippi ng

of the equipment offline which results in a plant transient, such as the loss

of feedwater, then the event tree should 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 will be similar to the one in

Figure 3.1.4 with certain random failure values modified to reflect the

availability of offsite power. BNL estimated that by transferring TS to

this new event tree, the only sequence which 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.

If,' in the reactor transient, there is a failure to scram, an event tree

similar to Figure 3.1.5 should be developed. It is conceivable that the

contribution to risk due to Class V sequences may not be negligible. It is

recommended that these considerations of additional accident sequences shoul d

be addressed in the LGS-SARA.

Not Event Quantification

LGS-SARA stated that non-failure states are included in the Boolean

expression of the accident sequences and therefore in the quantification

process. BNL performed some preliminary estimates of the core-damage

probability for the 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. In an earlier

draft of this report (August 15, 1983), a question was raised as to the

appropriateness of the LGS-SARA NOT event quantification. As a result of

discussions with PECo and its consultants, NUS, and of a re-examination of

BNL's preliminary estimates, it appears that there is a reasonable agreement

(see Tables 3.1.4 and 2.1.1).
3-18

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 is inferred that the failure of the ADS

hardware is included in the definition of X which is the manual depressurization function,

X = A + XR.

XR represents the random failure of the manual depressurization function; A

comprises seven different types of electric failures, including the loss of

the 440-V power supply, the 4 kV supply, the diesel generators, and the dc~

power. LGS-SARA conservatively ass 'umed that the failure of all these events

would lead to a failure of the ADS hardware. In essence, only the failure of

the dc power supply would lead directly to an ADS 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 440-V bus does not

result in failure of the ADS. It is for this reason that LGS-SARA is

conservative when it assumed that X = A + XR.

Since NOT events are considered important in sequence quantification ,

they should be included in the sequence evaluatioA..- However, a conservative

definition of X, may lead to nonconservatism in other sequences, which

although not necessarily manifesting itself in the change of the core-damage

frequency, may s ubstantially affect-the risk evaluation.

For instance, if accident sequence T SESUX (sequence No. 6 in.Figure

3.1.4), is calculated by assuming either the 440-V, the 4 kV, 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

nonconservative departure from the system modeling, since the operation of ADS

can only im ply that dc power is available.. This will1 tend to underestimate

sequences TSESU, TSESUW, and TSESUV. The impact may res~ide in

underestimating the contribution to accident Class IS, whereas the change in
3-19

core-damage probability may be inconsequential. Other risk indices, such as

latent and acute fatalities, may be affected differently.

One of the approaches to .addressi ng thi s concern i s to i ntegrate the ADS

hardware with the low pressure injection function, V, consistent with the

LGS-PRA definitions.

Sequence Quantification

The 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 quantification, 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: TSESCMC2

and TSRBCM. The value in parentheses 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 requaritified. The

value in parentheses denotes the revised system unavailability. The last

column is the core-damage frequency as a result of the requantification. The

sequence TSESUX is calculated in LGS-SARA to have a core-damage frequency

of 3.1x10-6 and if the manual depressurization function random failure is

modified to the new BNL value-of 6.0x10-2, BNL estimated that the core

damage will increase to about 4.0x10-6. It is assumed in the calculation

that beside 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 3.8x10-6 is observed. Increases in the failure to transfer from the CST

to the suppression pool produce similar results, 3.8x10-6. If all these

modifications are integrated into the accident sequence TSESUX, the
.3-20

total core-damage frequency is about 5.2 x 10-6, an increase by

approximately a factor of 1.7.

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-damage frequency

,for the accident sequence TSEsUX becomes 4.0 x 10-6. In other words, if

there is a total failure to transfer from the CST to the suppression pool,

because of 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 affected sequences affected-are the ATWS sequences. BNL

revised the definition of Cm 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,

CM =CR + S3 + S5 + S7

leads to an increase of about a-factor of 5 for-both the TSES CMC2 and

the TSRBCM accident sequences.

There- is no impact for the remaining three dominant sequences as a result of the modifications in Table 3.1.2. The total 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) TSESUV Accident Sequence

The core-damage frequency of the accident sequence T5E5UV is

calculated in LGS-SARA to be 5.9 x 10- 9. The Boolean expression of this sequence can be written as follows:

TSESUV =TSRPVRBES7ýUXV

=TS37 6S4SlVMXUV.
3-21

If one uses the definitions of V and X provided in LGS-SARA, the following

expression will result:

TSESUV =TSS 6WJ4TM XRJSlS2Sl7 + XRJS1VRHRRR(31

+ XRJS1S17HRRR + XR)NSlS 2vRG

where Si, i =1,2,...17 are the seismic-induced component failures; a detailed listing is given in Table 3.1.1. The bar above each variable denotes a

NOT event. CM is the mechanical failure to scram; A is seismic failure of

the electric power system; the subscript R denotes random failures. H and R

represent the HPCI and the RCIC systems, respectively. G denotes the combination of transfer and high pressure system failures, and is defined as follows:

G=HINIA * RIN3 + HINIA * RR + RIN3 * HR

However, if one uses the BNL definitions of X and V, namely, X = XR and V

LPCI *LPCS + ADS, where LPCS and LPCI are the same as those defined in LGSSARA, and where the added term ADS is the sum of the ADS hardware random

failure AR and the electric power A, the following Boolean expression is obtained:

TSESUV =TS5 6S4_CM XRSlA + 7RSlS 2Sl7

+ XRS1RRHRVR + 7RS1RRHRS17 (32

+ XRS1RRHRAR + 3CRS1S2 G VR

+ XRS1S2 G AR

VR represents the random failure of the LPCS and LPCI systems. Comparison

of the two expressions in Eqs. 3.1 and 3.2 indicates that except for NOT-A,

Eq. 3.2 contains all the terms of Eq. 3.1 and three more terms besides. 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 i s

suggested that a more detailed analysis be provided in LGS-SARA to better

identify the contribution of TSESUV to core damage and to the final risk.
3-22

(III) Other ATWS Sequences

BNL reviewed the LGS-SARA ATWS event tree and found that, in addition to

those two dominant ATWS sequences, TSE SCMC2 and TSR BCM, the contribution to

core damage from other ATWS sequences defined in Figure 3.1.5 is relatively

small.* However, with the BNL definition of CM, there will be about a factor

of 5 increase for all the ATWS sequences in Figure 3.1.5. The total ATWS

core-damage frequency reported in LGS-SARA is 8.1 x 10-7; by eliminating

the PCR and using the BNL CM 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 it is conservative in assuming that those failure modes defined

for the reactor internals and the CRD guide tubes will1 directly cause a

failure to scram. BNL tends to agree that the definition of failure modes of

these components may be conservate and would encourage additional analysis to

support the LGS-SARA assumptions. A refined analysis in this area is needed

since it will have signifi cant impact on the acute and latent fatalities.

Examination of ATWS function unavailabi'lities 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 W2

functions. BNL did not reassess those accident sequences affected by these

modifications; however, because of the change in magnitudes of some of these

functions, and the fact that significant contribution to risks comes from the

Class IV events, it will be prudent to evaluate the effects of these changes

upon the results on core damage as well as the final risks. Sensitivity

analysis would also provide helpful 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 require more detail ed
3-23

analysis, whereas others need additional information to substantiate the assumptions. Requantification of some changes was made wherever possible, and

results are discussed earlier in this section. It appears that for these

modifications investigated, at most a factor of 2 changes to the core-damage

frequency is observed. In view of the large uncertainty associated with the

seismic accident sequences, these changes in magnitude do not constitute any

significant impact on the core-damage frequency, but their effects on the

acute and latent fatalities may be significant.
3-24

4

FAILURE TO

PROVIDE ADEQUATE

FLOWJ FROM HPCI

RANDOM

LNDEPENOENT

FAILURES

m

0

I'LIXTURE OF

RANJDOM AND

SEISMIC

FAILURES

SEISMIC

FAILURES

A

Figure 3.1.1 Reduced fault tree for the HPCI system.
3-25

Figure 3.1.2 Figue Reduced 31.2 fault tree for failure to-scram.
Seismic

frequency

Reactor

pressure

vessel

Reactor

and

control

building

*Off~ite

power

available

Ile~ it: tor

SC1,dill

[No.

Se-quence

code

Meian

ainnual

sequence

frequency

or

trans fer

Degradud.-

core

claiss

SRPV R sJ C. 4 1

2

3

4_

5

6

TE SCM

TSR ECM

TSR PV

OK

Fig. 3-8

Fig. 3-9

9.60-7

1.40-7

8.0-7

Ws.

is

Is

St~III

T wVvas put in a special daI&s

'A seismiic evient that doet not load to the loss Of olfsite p)ovva it c"IagIeif 0'I tUu 11 uillap

and is adequately accounted for in turbine-trip initiating event.

Figlure 3.1.3 The seismic event tree.
Tiaifr rciiS/V.SlV.HPCI T-,ely -~ M a nru*jal O-j.%u1.

F-Wier 3-7 CNanfCIC hs SeQ.

4 .. 1.bla ~ wI u p w No. CW0I C: a

TSES M. P U x V W

1 TSES OK

2 TSESW 1,07-7 U/Is

3 TSESU OK

4 T5ESUW

5 TSESUV $59is

6 TSESIJ X 3 15 6 1

7 1i5EP OK~

8 TE I 107-9 tIIS

9 SSP OK

10 TSEsPIJW

I I TSESPUV

__________12 TSESPUX 3.15-8I

13 TSESM

Figure 3.1.4 Seismic event tree for loss of offsite power with reactor scram.
12 7 S E CMP 0OK

13 TO SrUVW2 I 10A2 COU

01 1 SrýrCM PU

16 T E V I (A) a

IB I1C1 OK

ID l~~~CG 2 ' 2 OK

20 lLICMCA 2

N2

21 1.11 CC 2 0

22 TO1 CIC I LU

24 lEIýIC, ~I P P

r4 T E CAA I

Figure'3.1.5 Seismic event tree for loss of offsite power without scram.
3-29

Table 3.1.1 Significant Earthquake-induced Failures

Medi an

ground

Failure cause acceleration

No. Component or mode capacity 0 R a U

g Ceramic insulator

S1 Offsite power (500/230-ky

switchyard)

S2 Condensate storage tank

S3 Reactor internals

S4 Reactor enclosure and

control structure

S5 CR0 guide tube

S6 Reactor pressure vessel

S7 Hydraulic control unit

S8 SLC test tank

Sg Nitrogen accumulator (SLC)

S1 SLC tank

S11 440-V Lius/SG breakers

S12 440-V bus transformer

breaker

S13 125/250-V dc bus

S14 4-ky bus/SG

S15 Diesel-generator circuit

S16 Diesel-generator heat and

vent

S17 RHR heat exchangers

breakage

Tank-wall rupture

Loss of shroud support

Shear-wall1 collapse

Excess bending

Loss of upper support

bracket

Loss of function

Loss of support

Anchor-bolt shearing

Wall buckle

Power circuit

Loss of function

Loss of function

Breaker trip

Loss of function

Structural

Loss of lower support

(anchor bolts)

0.20

0.24

0.67

1.05

1.37

1.25

1.24

0.71

0.80

1.33

1.46

1.49

1.49

1.49

1.56

1.55

1.09

0.20 0.25

0.23

0.28

0.31

0.28

0.28

0.36

0.27

0.27

0.27

0.38

0.36

0.36

0.36

0.32

0.28

0.31

0.32

0.25

0.35

0.22

0.52

0.37

0.20

0.19

0.44

0.43

0.43

0.43

0.41

0.43

0.32 0.34
3-30

Table 3.1.2 Mean Values for Random System or Function Failures

Used in Transient Events

t

LGS-SARA LGS -PRA NUREG/CR-3028

H PCI

RC IC

V

w

H IN IA

RI N3

DGC

x

8.8X10-2

7.6x10-2

1.0xl10 4***

2.6x10-4

1.0X10-2

7. 0x10-3

1.25x10-3

2. 0X10-3

0.07

0.07

2.7x10-4

1.0x10-2

7.0OX10 3

1 .08x10-3

2. 0x10-3

0. 1157

0.07

3. 7x1- 4 *

1.Ox10-2

7.OXlO-3

1.88x10'3

6.0x10-3

BNL

(this review)

0. 1157

0.07

3. 7x1- 4 *

2.6x10-4**

1.0x10-2

1.0

1.25x10-3

6.0x10-2

*Wi t

**Wi t

***Wi t

HPC I

RCIC

V

w

HINIA

RI N3

DGC

x

h

h

h

ADS hardware and

no offsite power.

LPCI and LPCS oni

no offsite power.

and only RHR.

- High Pressure Coolant Injection

- Reactor Core Isolation Cooling

- Low Pressure Injection Function

- Containment Heat Removal Function

- Failure to Transfer From CST to Suppression Pool in HPCI

- Failure to Transfer From CST to Suppression Pool in RCIC

- Diesel Generator Common Mode Failure

- Manual Depressurization.
3-31

Table 3.1.3, ATWS Mean Random Failure-Values

SARA LGS -PRA -NUR EG/CR-3 028 BNL (this review)

HR

RR

D

UH

C12

CM

HINIA

RI N3

W2

PCR

DGC

8.8x10-2

7.6x10-2

8.Ox 10-3

2.OXlO-3

1.6x10-2

1.5xl105

1.0x10-2

7.0x10-3

0. 1

0.2

1.2 5xI10 3

0.1

0.5

2.OX0 x-4

20x10-4

1.5X10-3

1.0X10-5

1.0x10-2

7.Ox1lO 3

0.14

1 .08x10-3

0.14

0.5

2x40-2

2. Ok 0 -

1.4 X10-2

1.0X10-5

1.0xI10 2

7.0OX10-3

0.14

1.88x10-3

0.14

01.5

2x10 1I

2.Ox10-3

1 .4X10-2

j.OxlO-5

1.0x10-2

1.0

0.14

1.0

1'.25xI10 3

HR

RR

D

UH

C12

CM

W2

PCR

HPCI random failure

RCIC random failure

ADS inhibit failure

Failure to control reactor vessel level 8

Failure of two or three SLC pumps

Scram fail ure-mechanical

Failure of both RHR

Fraction of events that lead to scram fail ure*

*Definition not given in LGS-SARA, inferred from modularized system fault tree.
3-32

Table 3.1.4

-Sequence

TSESUX.

TSRB

TSRPV

TSESCMC2

TSRBCM

TSESW

Cl-ass

/II

is

sI/Ill

Dominant Seismic,

LGS-SARA

3. 1x1-6

A.6x10-7

8. 0x10.7.

5.4x10 7

4

1.4x10_7

.1.1x10-7

5.7x10-6

Core Damage Sequences

BNL Estimates

2.8x10-6

9.5x10 7 .

4.4x10-7

6.,Oxj7

3. 5x10-7

1. 1x10-7

5.3xj10 6 Total
3-33

Table 3.1.5 Dominant Seismic Sequences With BNL Changes

Sequence (Core

Damage Probability)

TSESUX (3.1x10-6)

Sy stem Modi f ied

(Unavailability)

X (6x10-2 )

U (8.1x10-3 )

HINIA, RIN3 (1xI10 2 ,1.O)

All combined

TSESCMC2 (5.4xl10 7) CM

TSRBCM (1.4xl10 7) CM

TSPRV (8.0x10-7)

T5E5W (1.1x1,0-7

TSRB (9.6x10-7 ) -

Total 5.7x10-6

*Based on LGS-SARA sequence values.

**Sum total of TSESUX (combined) and the other 5 sequences.

Core Damage*

Frequency

4. 0x10-6

3.8xlcr6

3.8x10-6

5.2x10-6

3. 0x10-6

1.8XI10 6

8.0x10-7

1.1x10-7

9,6x1O-7

1.2xl10 5**
3-34

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 zone~s along with the corresponding results, to describe the BNL modifications to the quant ification, and to present the revised results. This section is organized as follows.

Section 3.2.1 summarizes, the LGS-SARA approach to quantification of accident sequences and presents the mean values for the frequency of core damage

for the di Ifferent fire zones. Section 3.2.2 presents the detailed BNL review

of the different fire types for two fire zones: Fire Zone 2, whose fire

growth event tree is similar in s tructure to all other fire zones except for

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 summary 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 i~dentified the following

steps used in the quantification of accident sequences:

1. Identification of potential initiating fires within the fire zone; the

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 intermediate stages

between ignition and damage to all safe shutdown systems served by cabling or components located within the fire zone.
3-35

4. Eval uati on of each f ire growth stage i n terms of (a) the probabi1i ty

of failing to suppress the fire before reaching each stage, and (b)

the shutdown systems that remain undamaged at each stage.

5. Evaluation of the conditional probability of core melt at each stage

of fire growth, taking credit only for the reliability of systems not

already damaged by the fire. This was achieved by modifying the fault

and event trees developed in'the LGS-'PRA.( 2)

6. Evaluation of the core-damage frequency associated with individual

fire growth stages by combining the frequency of failure to suppress

the fire at each stage of growth and the associated probabilities of

core damage from random failures of the undamaged systems.

7. Summation of the core-damage frequencies associated with each damage

stage for all types of fires to obtain the overall fire-induced coredamage frequency for the fire zone.

Following 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 contribution 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).

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 main disagreements found in this review are

summarized below.

a. A reduction factor of 5 in the frequency of self-ignited cable raceway

fires was used in the LGS-SARA report. As described in Subsection

2.2.2. 1.1, the BNL review indicates that a reduction factor of 3 is

more appropriate, if we use the existing data base.

b. It is the BNL judgment that the probability of fire suppression success is overestimated in the LGS-SARA report. On the basis of the

discussions in Section 2.2.2.3, the following probabilities of failure

to extingui~sh a fire in t minutes, P(t), are used in the BNL review:
3-36

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 BNL review of the, LGS-PRA(7)

(NUREG/CR-3028). It is noted that'a- computer reevaluation of-system

unavailabil-ity or core-damage fault ..trees was not made; only hand calculations

were performed.

The approach 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 seq'uences-for each fire growth stage.

3.2.2.1 Fire Zone 2: 13-kV 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 probabi'liti~es is discussed below.

Event A: Frequency of Cabl'e-R'aceway 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 cab le in the reactor enclosure and

control structure, (172,799 pounds) And (2) the frequency of cable fires per

reactor year: -
3-37

8,736(lb) x(5.3x10-3 8gj-/

172,799(lb) 3 ) 8.x0yr

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 Accident Given Fire Growth Stage 1

Since most of the cabling in this fire zone is associated with balance-ofplant (BOP) equipment, loss of the power-conversion system for inventory

makeup and long-term heat removal was assumed. At this stage all

safety-related equipment is undamaged, and the dominant a ccident 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.OxlO5

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

This stage represents damage to all safety-related -equipment except that

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:
3-38

*Class I

QUX = 4.9x10-5

QUV = 6.6x10-5 '

*Class IL,

QW = 4.5x10-3

PQW =.4.5xl10 5

*Total 4.7x10-3

Event E: Fire Suppressed Before Damaging Protected Raceways

This event is concerned with the probability of fail'ing to suppress this

fire before protected cables serving shutdown methods A and B are damaged.

This is equivalent to failure to suppress the fire within one hour after the

fire. This probability is equal to 8.0x10-2, using the BNL curve given in

Section 2.2.2.3; the 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-shutdown 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 melt is thus 1.0.

b.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., 1.8x10n3/yr.

Event B: Undamaged 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-conversi~on system.and the quantification of this event is

identical with that described for Event B in Section a.
3-39

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 qualitative comments about how this value was

obtained. However, the BNL review does not change this value.

Events D, E, and F

Given that a fire has propagated from the panel in which it originated to

adjacent cable raceways, the quantificati on of the conditional probabilities

associated with events D, E, and F is identical with that described in

Section a.

c. Quantification of the Fire Growth Event Tree for Transient-Combustible

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 trans ient-combusti bl e fires

given in the LGS-SARA report seems to be reasonable; this probability is equal

to 1.3x10-5/yr.

Events B, C, D, E, and F

The evaluation of the conditional probabilities associated with Events B

to F is identical with that described in Section a.

3.2.2.2 Fire Zone 25: Auxiliary Equipment Room

a. Self-Ignited Cable Fires.

The frequency of sel f-ignited cable f ires i n the raceways of -the auxil1iary

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
3-40

4 400(lb) x5.3x10-3 4.5x10-5/yr

172,799(lb) 3

In the LGS-SARA report it is argued that on the basis of 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 maximum fire damage

that could result is the loss of one division of safe-shutdown equipment, and

assuming the most demanding transient, MSIV closure, the dominant accident

sequences and their conditinal probabilities are:

*Class I

QUX = 5.6x10-4

QUV = 9.2x10Q5

*Class II

QW = 7.8xlcr.7

*Total = 6.5x10-4

Using these conditional probabilities, the resulting frequency of core

melt is (4.5x40 5) x (6.5x40 4) = 2.9x10-8/yr.

b. Self-Ignited Cable Fires.

The f requency of cabi net f ires i n the auxiIi ary equi pment room i s

estimated a's 1.75x0o4/cabinet-year. This auxiliary equipment room has' four

cabinets where fires may cause significant damage to safe-shutdown systems.

Assuming that a fire in any of those cabinets would destroy the contents of

the cabinet, the following equipment would still remain undamaged:

1. The RCIC or HPCI System

2. Means of Reactor Depressurization

3. The LPCI System (Two Trains)

4. The Core Spray System (One Train)

5. The RHR System
3-41

Assuming that the initiating event is a transient with isolation from the

power. conversion system (LGS-SARA assumption), the following are the dominant

accident sequences with their conditional core-melt probabilities:

*Class I

QUX = 5.6x10-4

QUV = 2.9x10-5

*Total .5.9x10-4

The core-melt frequency resulting from self-ignited panel fires is

therefore: 4 x (1.75x10- 4 ) x (,5.9x10-4 )- 4.1x10-7/yr.

c. Transient-Combustible Fires.

The frequency of transient-combustible fires were estimated as follows:

Trash-can Fire =3.4x10- 4/yr

Solvent-can Fire =3.4x10- 41yr

Oil Fires - 3.4x10-5/yr

Heat transfer analysis was used to evaluate cable temperatures resulting

from external-exposure fires, and on the basis of this analysis, locations

within the fire zone where 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 4.4 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 equipment room associated

with Unit 1, excluding the area taken up by cabinets), the core-melt frequency

is calculated and given in Table,3.2.3. It should be pointed out that the

dominant sequences for each fire location are QUX and QUV.

3.2.2.3 Fire Growth 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 coredamge frequency are presented.
.3-42

3.2.3 Review Results

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 resuits are:

1. The total core damage frequency resulting from fire-induced transients

obtained in the BNL re'view is 5.2x1 O-5/yr, as compared to

2.3x10-51yr reported in the LGS-SARA report.

2. The differe *nce between the BNL review'and the LGS-SARA core damage

frequency can be attributed to two factors: (a) the probability of

fi~re 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 fre~quency comes from the fire growth stage 3

(about 85% the both BNL review, and about 81% in LGS-SARA). At, thi s

fire growth stage, in almost all zones, all safe-shutdown '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 LGS-PRA quantification) have a small impact upon the

total fire-i'nduced core-damage frequency.

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 I

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).
3-43

Thus, calculations were performed to show the impact of these two modifications, and the results are as follows:

a. If the LGS-SARA probability of failing to suppress? the fire within 60

min. (0.04) is used instead of the BNL value (0.08), the total fireinduced core-damage frequency would be equal to 3.6x10-5 /reactor

year.

b. If the LGS-SARA reduction .factor (RF=5), used in the calculation of

self-ignited cable-raceway fires is used, instead of the BNL value

(RF=3), the total fire-induced core-damage frequency would be equal to

3.8xl0-5/reactor year.'

Another area where some se nsitivity study is warranted is in the evaluation-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 val ue of X given

in the BNL review of the LGS-PRA(7); i.e., X=6.0x10-3. If this value is

increased 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

*is used, the total fire-induce~d core-damage frequency is increased. to

5.6x10-5 (an increase of 7.7%).
,344

3.3 References to Section 3

1. Philadelphia Electric.Company, "Limerick Generating Station', Severe Ac-

.cident Risk Ass~essment," April 1983.

2. Philadelphia-Electric Company, "Limerick Gener ating Station, ProbabilI i sti c

Risk Assessment," March 1981.

3. Kolb, D. L., et al., "Review and Evaluation of the Indian Point

Probabilistic Safety Study," NUREG/CR-2934, SAND82-2929, December 1982.

4. Pickard, Lowe, and Garrick, "Indian Point Probabilistic Safety Study."

Prepared for Consolidated Edison Company of New York, Inc., andPower

Authority of the 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 Systems at Nuclear Power Plants," NUREG/CR-2989, ORNL/TM-8545,.July 1983.

7. Papazoglou, I. A., et al., "Review of the Limerick Generating Station

Probabilistic Risk Assessment.," NUREG/CR-3028, BNL-.NUREG-51600, February

1983.
A

Fire In

cable/TC/pa nel

FGS 1

B

Undamaged

Systems

mitigate

accident given

FGS1

C

F ire suppressed

before damaging

unprotected

iaceway

lFailure gives FGS-2)

D

Undam aged

systems

mitigate

accidenit given

FGS2.

E

Fire suppressed

before ddiamaing

prutected

race ia Vs

(FIulue ijives FGS3)

F

Undamaged

sysitems

niitigate

accidtent given

F GS3

Core

1Stj'. Itiis I

Annual $equeCncO

flequiuiicy

I

- 4

ý1.0/1.0/4.0-2

F_

8.0-2

8.9-5/1.3-5

1ý.8-3

/ 1.0

OKOK

OK

CM 7.1-6 1 .0-6 5.8-6

(see not~e) (see

note)

CM 4. 2-7 6.1-8 3.4-7

CM 5.3-9 17 .8-l 11.1 -7

cii

U,

4.7-3

6.0-5

- I V - I -

1 7 . 5-6 1 .1-616.2-6

FGS = Fire growth stage

TC =Transient comTbustible

Note: Because of the evaluation of event E, the probability of eve~nt C

is not included in twe evaluation of the sequence freueluncy.

Total annual core-melt frequency 1 . 5-5

F igure 3. 2, 1 F ire-growth event tree f or f ire zone 2
2

20

22

24

25

44

45

47

Table 3.2.1

Fire Zone Ra

1,2-ky switchgear

room 2

Static inverter

room 5

Cable-spreading

room 6

Control room N

Auxiliary equipment

room N

Safeguard access

area 4

CRD hydraulic

equipment area .4

General equipment

area1

3-46

Summa ry of Fire-Analys is Results

Annual Contribution to Core-Melt Frequencya

lf-Ignited TransientCable Self-Ignited Combustible

ceway Fire Panel Fire Fire Total

.1-8

egl igible

egl~igi ble

.2-6

.7.-6-

.2-6

3--:5.

3.2-6

3.5-8

NAc

1.6-7

1.0-7

1.5-6

1.0-6

5.0-7

6.5-6

5.9-7

1.5-8

1.9-7

1.0-7

2.6-7

4..1-7

6.6-7

1.8-7

2.4-6

6. 2-6

1.0-7

2.5-7

2.6-7

3.6-7

6.1-6

6.4-6

1.9-6

2.2-5

1.0-6

2.3-5

Contribution from all

other fire zones

Total annual core-melt

frequency from fires

apoint estimates

b2 .4 -6 = 2.4x10-6

cNot applicable
3-47

Table 3.2.2 Critical Locations of Transient Combustible

Materials in the Auxiliary Equipment Room*

Area of location (in2 )

Sol vent-mCan

Fire

Oil

Fi re

Systems assumed to be undamaged and capable of RPV

Fire Location inventory makeup

Intersection of floor areas

10U792 (a) and 10U791

0 2.4

Intersection of floor

10U791 and 10U793

Floor area 10U795 (c)

Floor area 10U789 (d)

(b) areas 7.7

0.6

0.6

12

LPCI train D, means of

depressurization

LPCI train D, means of

depressuri zati on

LPCI trains B and C, means

of depressuri zation

LPCI trains C and D, means

of depressurization

2.3

2.3

*Table 4.4 of LGS-SARA.
3-48

Table 3.2.3 Evaluation of Sequence Frequencies of

Oil Fires (Transient Combustibles)

Fire Locationa.

Annuala a

Frequency

Criticala

Locati on

Probability

Probabil1ity b

of Random.

.Equi pment

Failure

Core-Mel. tb

Frequency

OIL FIRE

Location

Locati on

Locati on

Locati on

a

b

c

d

3.4-5

3.4-5

3.4-5

3.4-5

0.01

0.05

0.01

0.01

0.022

-0.022

0.014

.0.014

7.5-9

3.7-8

4.8-9

4.8-9

Location

Location

Locati on

Locati on

a

b

C

d

SOLVENT FIRE

3.4-40

3.4-4 . 0.03

3.4-4. 0.1003

3.4-4 0.003.

0

0. 022

0.014

0.014

0

ý2. 2-7

1.4-9

1.4-9

Total 2. 8- 7c

aFrom LGS-SARA Table 4.5

bBNL Review

cThe corresponding LGS-SARA value is 2.6-7.
2

20

22

24

25

44

45

47

Table 3.2.4

Sel

Fire Zone Rac

12-ky switchgear

room

Static inverter

room

Cable-spreading

room

Control room

Auxiliary equipment

room

Safeguard access

area

*CRD hydraulic

equipment area

General equi pment

area

3-49

Summary of Fire-Analysis Results

BNL Review

Annual Contribution to Core-Mel

f-Ignited Trans

Cable- Self-Ignited Combus

eway Fire Panel Fire Fi

t Frequencya

ienttible

re Toi Lal

7.5-6

2.4-7

3.7-7

NOb

2.9-8

1.3-5

9.6-6

3.9-6

3.5-5

6.2-6

7.5-8

NAb

4.8-7

4.1-7

.3.3-6

1.8-6

1.7-7

1.2-5

1.1-6

4.3-8

7.4-7

2.2-7

2.8-7

7.8-7

8.6-7

3.7-7

4.4-6

1.5-5

3.6-7

1.1-6

7.0-7

7. 2-7

1.7-5

1.2-5

4.4-6

5. 1-5

1.0-6

5.2-5

Contribution from all

other fire zones

Total annual core-melt

frequency from fires

apoint estimates

bNot applicable

4-1

4.0 SOME GENERAL ISSUES AND SPECIFIC RECOMMENDATIONS*

4.1 Seismic Hazard and Fragility Recommendations

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 given in this section. These recommendations are primarily directed to

PECo and are based on discussions already presented in Section 2.1. Rathe r

than repeating the background, each recommendation is presented and followed by

the applicable subsection in Section 2.1 which can be referred to for additional information. Also, recommendations are made to the NRC to perform additional review tasks to complete the review of the LGS-SARA.

Section 4.1.2 gives the recommendations for the hazard analysis and Section 4.1.3 gives the recommendations for the fragility and associated syst em

analysis concerns.

4.1.2 Seismic Hazard

The following recommendations should be addressed by PECo. The numbers in

parentheses at the end of each recommendation 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. Specific~ally, a redefinition of Zone 8 i s recommended that is better correlated to the pattern of seismicity in

the vicinity of Limerick and the geologic structure of the Triass ic

Basin (see Section 2.1.2.3).

2. The possible occurrence of large-magnitude events (i.e., M7.0) should

be considered as an alIternative 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).

*These~recommendations to PECo and NRC on how to improve the PRA are provi ded

as a result of our short-tern review. They are not intended as necessary conditions to be fulfilled in order to assure safety in licensing considerations.

Such considerations are beyond the scope of this study.
4-2

3. The uncertainty in Richter b-values should be considered in the

seismic hazard analysis. Consideration 'should be given to the distribution of earthquake magnitud 'es 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 largemagnitude (i.e., M = 6.8) events considered in the hazard analysis.

Specifically, the basis for assumming that the magnitude estimated for

the 1886 Charleston, South Carolina earthquake is the largest event

that can occur should be provided. 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 Piedmont

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 Brunswick Canada event is included in the

Piedmont province (see Section 2.1.2.4).

The following recommendation is addressed to the NRC.

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 performed.

4.1.3 Seismic Fragility

The following recommendations should be addressed by PECo. The numbers i n

parentheses at the end of each recommendation refer to the subsection of Section 2.1 which gives background information.

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 provided. Specifically, the concern 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).
4-3

2. Justification should be provided for the median duration factor.

Specifically, the median value of 1.4 and the variability 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. A median capacity value greater than 0.90g for the reactor enclosure

and control structure should be justified (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, CRD guide tube, and reactor

pressure vessel should be justified. 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). Our

understanding is that a weighted model damping analysis was performed

and a value between 9 and 10 percent was obtained;(26) however, we

have not reviewed that analysis. Note that this recommendation has a

lower priority since the mean frequency of core melt would increase by

only 10 percent for this effect.

5. The implications of impact between the containment building 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 two

buildings due to relative displacements. In particular, the various lines between the two structures should be systematically reviewed to verify that relative displacements will not decrease the

structural capacities.
4-4

In addition, it should be verified that no safety-related components

will be damaged by spalled concrete caused by impact of the two

structures (see Section 2.1.3.5).

Finally, it should be systematically verified that failure of small

lines (due to falling concrete) attached to the safety-related piping.

near the junction of the two structures and anchored to the reactor

enclosure will not contribute to t he frequency of core melt.

6. In regard to the safety-relat ed electrical components which

significantly affect the frequency of core melt including, but not

limited to:

* 440-V bus/SG breakers (S11)

* 440-V bus transformer breaker (S12)

* 125/250-V dc bus (S13)

* 4-Ky bus/SG (S14)

* Diesel-generator circuit breakers (S15)

identify the number of actual components, their locations, and their

characteristics relative to the generic tests at Susquehanna which

were used to derive their capacities. Justification should be

provided for the number of each c omponent type which should be

included in the Boolean equation for sequence TsEsUX. Consideration should be given to the possible effects of capacity and response dependencies which exist (see Section 2.1.3.6).

7. Justification shouldb 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 c apacity'values for the following Limerick components:

Hydraulic control unit (07)
4-5

* Nitrogen Accumulator (59)

* 440-V bus/SG breakers (S11)

* 440-V bus transformer breaker (S12)

* 125/250-V dc bus (S13)

* 4-KV bus/SG (S14)

* Diesel-generator circuit breakers (S15)

Consideration should be given to the location of the components~in the

two plants, foundation conditions, and construction similarities.

Based on the response given by PECo at the September 26, 1983, meeting, it appears that the equipment is located higher in Susquehanna

than L'imerick and that a factor of 2 conservatism may exist in the

median fragility values used for Limerick. It is recommended that

fragility parameter values specifically calculated for each of the

above components at Limerick be developed (see Sections 2.1.3.6 and

2.1.3.71).

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 c onfiguration. The capacity of

the anchor bolts should be checked and the earthquake component

factors derived based on the actual response and capacity

characteristics (see Section 2.1.3.7, Component S8).

9. The similarity bet ween the nitrogen accumulators at the Limerick and

Susquehanna plants should be verified since the analysis from Susqeuhanna was used as the basis for the capacity of the nitrogen accumulator at Limerick (see Section 2.1.3.7, Component S9).

10. The possible failure of the SLC tank due to tearing of the base pl ate

flange near the anchor bolts should be checked to verify that it i s

not the weakest capacity (see Section 2.1.3.7, Component SID).
4-6

11. A specific analysis should be conducted for the diesel generator heat

and vent which is based specifically on the characteristics of this

component (see Section 2.1.3.7 Component S16)-

12. After construction of the plant is completed, a systematic review of

the plant, including walkthroughs, should be conducted to locate

secondary components which could fail , fall , and-impact primary

s~afety-related components. Analyses of potential fail-ures should be

conducted to determine whether the secondary components are weaker

than the primary components already considered (.see Section 2.1.3.8).

13. The percentages of occurrences when evacuation would be affected by

earthquakes should be recalculated using realistic relationships between damage to civil structures an d 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 engineers should be

performed since these components are major contributors to the mean

.frequency of core melt.

2. Other significant nonelectrical components are based. on generic

capacities. Independent,ý specific calculations should be performed

for the following components since they are important to the final

risk.

Hydraulic Control Unit (S7)

Nitrogen Accumulator (S9)

Diesel Generator Heat and Vent (S16)
4-7

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

judgment rightly so, to use a hybrid of all three. A hybrid approach is

indeed warranted. Physical models suffer from the complexity of the large

number of variables and relationships required to calculate a fire history.

Point probability models suffer from small and inadequate data bases. While a

completely probabilistic approach also suffers the inadequacy, a more serious

deficiency is its inability to accurately model certain phases of fire

devel opment.

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 combustibles), and inert

surfaces (walls and ceilings). As Friedman(l) points out .1rather simplisti -

cally, 20 interaction vectors involving heat and, ",terial flux exist between

these 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- principle, to permit description of the growth rate of the fire. However, in

order to make safety assessments, it is also mandatory to have additional i nformation, such as carbon monoxide and smoke content, for its impact on plant

personnel safety. More important, 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 process. 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
4-8

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.

Furthermore, the very conservative assumptions used in the Limerick SARA

fire analysis (in most respects) may, if taken out of context, lead to a

distorted perspective of fire risk relative to other risks at the plant.

In some respects, assumptions and submodels that are touted to be conservative are tantamount to gross violations in physical realities. Several

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 which have been

addressed directly.

The Limerick SARA on fire analysis has considered only intrazone fire

propagation. A true assessment of fire risk must consider interzone fire

propagation and all aspects pertaining thereto, including the debilitating

effect of smoke migration, which has no immediate bearing on component

reliability, but which should have immediate implications with regard to

manual suppression effectiveness. Hence, smoke propagation should have been

considered even if its level of sophistication is 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 deta~il

commensurate with the pro~babilistic/deterministic analysis that is applied to

assess fire risk.

The foundation on which the fire propagation model, basically a one-room

fire model, rests is sound. Various compartment fire models(2) have been

developed and COMPBRN can -be considered as one which lies within their
4-9

This zonal approach has several important advantages: (1) computational

simplicity, (2) ease of decoupling zones for independent investigation, ( 3)

simpler comparison 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 model s

(1) are limited by computer capacity, (2) do not yet properly treat actionat-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 judgment, be pursued with the field approach used as

a basis for "fine-tuning" the unit models that are built into the zone-model

approach.

Zone models, like COMPBRN, represent a nearer-tern engineering approach

which is closely tied to experimental observations. However, a basic phil osophical limitation in zone-model structure is. its emphasis on predicting room

flashover. For an assessment of nuclear power plant risk, a prediction of the

onset of flashover is not as crucial as a prediction of the effects of

in-place component vulnerability during the earlier fire-growth stages. Thus,

for completeness a larger spectrum of initiating fire sizes must be

incorporated into the analysis.

Accord~ingly, several of the unit-models employed 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 turbulence

induced buoyancy on plume development, (4) intrazone mass and energy exchange,

and (5) implementation of existing knowledge and correlation of

fuel -fl ammabil ity characteristics, specifically, current cable flammability

and damageability indices.
4-10

Another keypoint regarding the practical use of a zone model in general,

and COMPBRN in particular, is that the structure of the numerical code i s not

IIuser friendly." Before one can use a code employing a series of unit models,

one must be aware of the assumptions built into the analysis, the key physical

parameters and their sensitivities, and finally a working knowledge of the

state of the art in fire phenomena and model~ing.

4.3 References to Section 4.2

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,

April 1983.
A-1

APPENDIX A: Detailed Review of 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 growth event tree for all types of fires in Zone 20 is shown in

Figure A.1.

A.1.a Quantification of the Fire Growth Event Tree for Self-Ig nited CableRaceway 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

x () 9.8x10-5/yr. 172,799(lb) 3

Event B: Undamaged Systems Mitigate Accident Given Fire Growth Stage 1.

On the basis of the locality of the initial fire a reactor-trip transient

is ass umed, with the loss 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
A-2

*Class II

PW =4.5x10- 5

QW =5.4x10- 6

*Total = 7.OxlO-5

Event C: Fire Suppressed Before Damaging Unprotected Raceways.

The probability 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 Growth Stage 2.

Fire growth stage 2 represents damage to all safety-related equipment

except that associated with shutdown method A which it served by cable

raceways protected with ceramic-fiber fire blankets; a~lso unaffected is

equipment associated with the power-conversion system.. The dominant sequences

and their conditional probabilities are as follows:

*Class I

QUX = 8.4x10-6

ýQUV =2.6x10- 5

*Class II

PW = 6.6x10-5

QW = 7.9x10-6

*Total = 1.1x10O4

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 hour (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-con'version system would remain

undamaged to mitigate the accident. The dominant sequences with their
A- 3

conditional probabilities are:

" Class.I

QUV = 2.OxI10 2

" Class II

pW = 1.0x10-2

Total 3.OxlO-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.

On the basis of the panels located in this zone, a reactor-trip transient

is assumed, and the following equipment is assumed to have failed: HPCI, RHR

Trains B and D and Train B of LPCS. The dominant accident sequences and their

conditional probabilities are:

*Class I

QUX = 8.4x10-6

QUV = 1.1X10-5

*Class II

PW = 4.5x10-5

QW = 5.4x10-6

*Total = 7.0x1O05

Event C: Fire Suppressed Before Damaging Unprotected Raceways.

The quantification of this event is identical with that for Event C i n

Panel Fires for Fire Zone 2 (see Section 3.2.2.1.b).
A-4

Events D, E, and F

The quantification of the conditional probabilities associated with those

events is identical with that described for self -ignited cabl e-raceway fires

in Section A.1.a.

A.1.c _Quantification of the Fire Growth Event Tree for Tran~sient-Combustible

Fires.

Event A: Frequency of Transient-Combustible Fires.

BNL agrees with the freque~ncy of transient-combustible-ýfires as calculated

in LGS-SARA, i.e., 1.7x10-5/yr.

Events B, C, D, E, and F

The evaluation of the conditional probability associated with Events B, C,

D, E, and F is identical with that described in Section A.1.a.

A.2 Fire Zone 22: Cable-Spreading 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 Cabl eRaceway Fires.

Event A: Frequency of Self-Ignited Cable-Raceway Fires.

This frequency is calculated in the same way as fo .r Event A in Section

3.2.2.1.a, i.e.,

35,526(0b) 5.3x10-3

x ( ) = 3.6xl104Iyr. 172,799(lb) 3

Events B and C

Since all fires are capable of damaging adjacent cable raceways, except

those protected by a ceramic-fiber blanket, Event B is effectively omitted and

Event C is assigned a probability of 1.0.
A-5

Event D: Undamaged Systems Mitigate Accident Given Fire Growth Stage 2.

The initiating event is a transient with isolat~ion from the

power-conversion system, and the only equipment potentially operable is that

associated with shutdown methods A and B. The dominant accident sequences and

their conditional probabilities are:

*Class I

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-in, thick ceramic-fiber blanket which is equivalent to a

1/2-hr fire rating. Thus, Event E is assigned a probability of 1.95x10-1 ,

which is the probability of failing to suppress a fire within 1/2 hr.

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 the 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.OxlO4

PQW =6.6x10- 5

.Total = 3.1xI10 3
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 B, C, D, 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 Zone 24: The Control Room

Since there is no exposed cable insulation in the control room, the only

types of fires analyzed i n this section are: Sel~f-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.

This stage represents damage that is confined to the cabinet in which the

fire starts. There are 17 separate'.cabinets in the control room. However,

only fires in .3 cabinets can cause significant damage. Fires in one of those

cabinets may disable all systems required for reactor s~hutdown except for

equipment controlled from the remote shutdown panel. Fires in the other two

cabinets will only disable the power-conversion system.
A-7

The transient resulting from any of these fires is Loss of Feedwater or

MSIV Closure and the dominant'accident sequences, with their conditional

probabilities, are:

*Class I

QUX = 3.1x10-5

QUV = 1.3x10-4

*Class 11

QW =2.5x10- 5

PQW 3.9x10-6

*Total = .9x10-4

Event C: Fire Suppressed.Before Propagating Beyond the Confinement of the

Cabinet.

BNL agrees with the evaluation of the probab~ility of a cabinet fire

propag'ating beyond the confinement of the cabinet as given in LGS-SARA, i. e.,

2.5x10-2.

Event D: Undamaged Systems Mitigate Accident Gi'ven Fire Growth Stage 2.

In this stage, only the equipment which can be operated from the remote

shutdown panel is considered potentially operable. The dominant accident

sequences and their conditional probabilities are identical with those

calculated for Event F in Section A.2.a, i.e., the total conditional coredamage probability is equal to 3.1x1Q-3.

A.3.b Quantification of Core-Damage Probability for Transient-Combustible

Fires.

BNL agrees with the quantification of the frequency of transientcombustible fires which can damage safe-shutdown equipment in the control

room. This frequency is equal to 7.2x10-5/yr.

Given the occurrence of a transient-combustible fire, it is assumed that

only the equipment that can be operated from the remote shutdown panel is

potentially operable. In this case the dominant accident sequences and their
A-8

and their conditional probabilities are identical with those calculated for

Event F in Section.-A.2.a; i.e.., the total conditional core damage frequency is

equ al to 3.1x10O_3 . So,. the, to Ital, co~ntribution of transient-combustible

fires to the core-damage frequency is given by-:

(7.2x,10- 5 ) x (3.1x1O03) =2.2x10- 7/yr.

A.4 Fire Zone 44: Safeguard Access Area

The fire growth event tree for all types of fire s in Zone 44 is shown inFigure 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 way as for Event A in Section

3.2.2.1. 1, i.e.,

28,290(lb) 5.3x10-3

172,799(LB) 3 )=2'x0/r

Event B: Undamaged Systems Mitigate Accident Given Fire Growth Stage 1.

The a-ccident-initiati'ng event wa's taken to be a transient withMSIV

closure, and at this stage of the fire the following systems would remain

potentially operable: RCIC or'HPC.I system, the ADS, the R.HR, system (three

trains), and the LPCS (one train). The dominant accident sequences and their

conditional probabilities are:

Class I

QUX*= 566x1W-4

QUV =9.2x10- 5

Class II

QW =7.8x10- 7

Total =6.5x10- 4
A-9

Event C: Fire Suppressed Before Damaging Unprotected Raceways.

The probability of this event is given by the probability of failure to

suppress the fire within 10 minutes (estimated time before damage to

unprotected raceways). The BNL value for the probability of this. event i s

0.43.

Event D: Undamaged Systems Mitigate Accident Given Fire Growth Stage 2.

Given fire growth stage 2, the following equipment would remain

potentially operable: the ADS and the RHR sy stem (2 trains). The domi nant

accident sequences and their conditional probabilities are as fol~lows:

*Class I

QUX = 6.Ox10-3

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 failure to

suppress the fire within 1 hour (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 the Fire Within 1 hr.

= 0.5 x8.0 x10-2 =4.Ox,10-2

Event F: Undamaged Systems Mitigate Accident Fire Growth Stage 3.

In this zone, fire growth stage 3 represents 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 PowerDistribution Panels.

Event A: Frequency of Fires.

The frequency of panel fires is determined from the number of panels
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 ) =1.5xl10 3/yr..

Event B: Undamaged Systems Mitigate. Accident 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

dominant accident sequences are:

Class I

QUX =5.2x1O0 6

Class II

QW = 1.1X10-8

PW =9.4'x1- 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.

Events D, E, and F

Once the fire has pro pagated to cable raceways, the quantification of

Events D, E, and F is identi-cal with that given in Section A..4.a for selfignited cable-raceway fires.

A.4.c Quantification of the, Fire Growth Event Tree for Trans ient-Combusti bl e

Fires

Event A: Frequency of Fires.

BNL agrees with the frequency of fires. calculated in LGS-SARA, i.e.,,

1. 7x10 5/yr.
A-11

Events B, C, D, 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 Equipment Area

The fire-growth event tree for all types of fires in Zone 45 is shown i n

Figure A.5.

A.5.a Quantification of the Fire Growth Event Tree for Self-Ignited CableRaceway 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. ,

18,637(lb) 5.3x10-3 l.x10-4 y -x (- ) = 19 /r

172,799(lb) 3

Events B and C

The quantification of these events is identical with that for the same

events in Section A.4.a.

Event D: 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 equipment potentially operable is tha"'

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

*Total = 3.OxlO-3
A-i12

Event E: Fire Suppressed Before Damaging Protected Raceways.

The probability of this event would be given by the probability of failure

to suppress the f ire withi n 30 mi nutes (time to damage to protected raceways).

However, only fires in one quadrant (northeastern) are capable of damagi ng

equipment associated with both shutdown methods. So, the 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 with the event is 1.0.

A.5.b Quantification of the Fire Growth Event Tree for S elf-Ignited Panel

Fires.

Event A: Frequency of Panel Fires.

BNL agrees with LGS-SARA evaluation of panel fires in this zone, i. e.,

3 x (2.2x10-4) = 6.6xl10 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 t wo of the three panels can. cause a turbine-trip

transient and at the-same time disable one high pressure-injection system

(HPCI or.RCIC) and one RHR train. The dominant accident sequences and thei.r

conditional probabilities are as follows:*

Class I

QUX = 1.1x10-5

QUV = 1.8x10-6

Class II

PW =1.3x10-7

QW = 1.6x10-8

Total = 1.3x10-5
A-i13

Since only two of the three panels can contribute to the accident

sequences, the probability of Event B is: 1.3x10-5 x 2/3 =9.0xj0-6.

Event C: Fire Suppressed Before Damaging Unprotected Raceways.

The probability of this event is identical with that for the, same event in

Section 3.2.2.1.b.

Event D: Undamaged Systems Mitigate Accident Given Fire Growth Stage 2.

The quantification of this event is identical with that of Event D in

.Section A.5.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

- x Probability of failing to suppress the fire within 30 minutes

3 (time to damage to protected raceways)

1 x 1.95x101 

,= 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.5.c Quantification of the Fire Growth Event Tree for Transient-Combusti bl'e

Fires.

Event A: Frequency of Transient-Combustible Fires.

BNL agrees with the frequency given in LGS-SARA, i.e., 1.7x10-5/yr.

Events B, CD, E, and F

Given a transient -combustible fire that causes the ignition of cable
A-14

trays, the evaluation of all these events-is identical with that for the same

events in Section A.5.c.

A.6 Fire Zone 47: General Equipment Area

The fire growth event tree for all types of fires in Zone 47 is shown in

Figure A.6.

A.6.a Quantification of the Fire.Growth Event Tree for Self-Ignited Cable

Raceways.

Event A: Frequency ofSelf-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.,

17,791(lb) 5.3xj10 3

172,799(lb) 3 ) 1.x0/r

Events B, C, and D

The quantification of these events is ide~ntical with that-for the same

events in Section A.5.a.

Events E: Fire Suppressed Before Damaging Protected Raceways.

The probability of this event would be given by the probability of fai ling

to suppress.the fire within 1 hour (time to damage to protected raceways).

However, only fires in one quadrant (NE) are capable of damaging equipment

associated with both shutdown methods., So, the probability of Event E is

given by: 8.0x10-2/4 = 2.Ox1lO 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 Quantification of the Fire Growth Event Tree for Self-Ignited Panel

Fires.

Event A: Frequency of Self-.Ignited Panel Fires.
A- 15

,BNL 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 the 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 probabilities are

*Class I

QUX = 4.9x10-5

QUV = 8.1x10-6

*Total = 5.7x10-5

Since only fires in three of the five panels are contributors to those

sequences, the probability of Event B is given by

3 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 Mitigate 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 D in

Section A.6.a.

Events E and F

It is BNL judgment 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.
A- 16

A.b.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, D, E, and F

Given a transient-combustible fire that ign ites cable trays, the

.quantification of these events is identical with that in Section A.6.a.
A

Fire In

cablefTC/pa nel

FGSI

8

Undamaged

systems

mitigate

accident given

FGSI

C

F ire suppressed

beinre damaging unprotected

raceway

(Failure gives FGS-2)

D

Undamaged

systems

mitigate

accident given

FGS2

.E

Fire suppressed

before damaging

protected

raceways

(Failurfe gives FGS3)

F

U ndamaged

systems

mitigate

accident given

F GS3

Core

St. Annual suquence fre tiucncy tits

- 1

I 8.0-2

cabl TC pae

OK

OK

ýCM 2,.3- 7 4.1-8 4.2-8

(see (see

note) note'

CM 4.6-9 8.0-D1.j9-(.

CM. 7.0-9 1.2-9.3.1-8

1.8-5/1 .7-5/-

0.43/0.43/4.0-2 13.0-2

1.1-4

7.0-5

-4 4-I1 2- 7 4. 3-81 7.5-8 FGS = Fire growth stage

TC = Transient combustible

Note: Because of the evaluation of event E, the Probability of eve~nt C

is not included in the evaluation of the sequonce freqluency.

Total annual core-melt frequency =3.6-7

Figure A.I1 Fire-growth event tree for fire zone 20
A B

Fire In

c-ablef'TC/panel

FGS 1

Undamaged

systems

mitigate

accident given

FGS1

C

F ire suppressed

before damaging

unprotected

raceway

(Failure gives FGS,2)

ABD E F

Undamaged

systems

mitigate

accident given

F GS2

Fire suppressed

before damaging

protected

raceways

(Failure gives FGS31

Undamaged

systems

mitigate

accident given

FGS3

Core Annual sequence frequency

-4

1.95-1 I

OK -

OK

OK

CM 2.2-7 4.3-7

CM 1.5-7 3.1-7

CM1

0-.

3.6-4/7.2-4 1.0

4. 3-4

NA

-,

3,.7-7 17 .4-7

FGS = Fire growth stage

TC = Transient combustible

Note: Because of the evaluation of event E, the probability of event C

is not included in the evaluation of the sequence frnequency.

Total annual core-melt frequency -1 .1-6

Figure ý. 2 . Fire-growth event tree far fire zone 22
A

Fire in Panel

FGSI

a

Undamaged

sys terns

mitigate

accident given

FGSI

C

Fire suppressed

before spreading

beyond confinement

of cabinet

(failure gives FGS2)

0

Undamuaiycd

systemsn

Un111it edt

aciXCUfiit IJIVCUI

FGS2

Core

I-.

FGS zFire growvt) stagec

TC =Transient combustible

Contribution to CM from TIC fires 2 .2-7

Total core-melt frequency 7.0-7.

per year

Figure A.3 -Fire-growth event tree for fire zone 24.
A

Fire in

Ca ble/T C/p anel

FGSI

8

Undamaged

systems

mitigate

accident given

FGSI

C

F ire sLIppressed

unprotected

raceway

(Failure gives FGS 2)

D

Undamaged

mitigate

accidenit given

FGS2:

E

Fire suppressed

beoloe damafjint

pro tec ted

race.wayl

(Fjilture gives FGS31

F

Undamaged

systems

mitigate

accident given

FGS3

Core

1St a Annual sequence frequeniicy

i

4.0-2

cabl - -pai

OK

OK

CM 1.1-5 .7-7 2.4-6

(see (see

note) ote)

CM 1.7-6 .0-7 3.5-7

CM 1, 9-7 it. 1-8 17. 9-9

C0

2.9-4/1.7-5

1.5-3

0. 43/0. 43/0. 04

1.4-2

4

6.5-4/6.5-4/5.3-6

i

1 .3-5 7.8-7 13.3-6

1.3-5 FGS = Fire growth stage

TC = Transient combustible

Note: Because of the evaluition of event E, the probability of event C

is not included in the evaluation of the sequoice frequeIIncy,

Total annual core-melt frequency = 1.7-5

Figure A.4. Fire-growth event tree for fire zone 44
A

Fire in

CablefTC/panel

FGS I

8

Undamaged

systems

mitigate

accidenh given

FGSI

C

F ire supprtessed

before damaging

unprotected

raceway

(Failure gives FGS,2)

D

Undamaged

systems

Mitigate

accident given

FGS2

E

,Fire suppressed

before damaging

protected

raceways

(Failuire gives FGS31

F

Undamaged

sys$tems

mitigate

accident given

F GS3

Core

ISt~i- fltl

Annual sequence frequency

;.875-2/4.875-2/6.50-2

OK -

OK

CM 9.2-6 8.3-7 1.7-6

(see (see

note) note)

CM 2.4-7 2.2-8 7.9-8

CM 1 .2-7 1.1-8 5.9-9

.9-4/1 .7-5/

ý.6-4

0. 43/0. 43/0. 04 11.0

3.0-3

6.5-4/6.5-4/9.0-6

9.6-6 8.6-711.8-6

FGS = Fire growth stage

TC =Transient combustible

Note: Because of the evaluation of event E, the probability of event C

is not included in the evaluation of the sequence frequency.

Total annual core-melt frequency I= 1 .2-5

Figure A.5. Fire-growth event tree for fire zone 45
A

Fire In

cablefTC/panel

FGS I

B

UnAdam aged

systems

mitigate

accident given

FGSI

C

Fitre suppressed

ibQfore damaging

unprotected

raceway

(Failure gives FGS ,2)

D

Undamaged

sy sterns

mitigate

accident given

FGS2

IE

Fire suppressed

before damaging

protected

raceways

(FailureC gives FGS3)

F

Undamaged

systems

mitigate

accident given

FGS3

FGS = Fire growth stage Total annual core-melt frequency 4 .4-6

JC = Transient combustible

Note: Because of the evaluaition of event E. the probability of event C

is not included in mte evaluation of the sequence frequency.

Figure A. 6 Fire-growth event tree' for fire zone 47
B-1

APPENDIX B: A Critique of "Seismic Ground Motion at Limerick

Generating Station." by ERTEC Rocky Mountain, Inc.

December 8, 1983

Professor Alan L. Kafka

Weston Observatory

Dept. of Geology and Geophysics

Boston College

Weston, MA 02193
B-2

INTRODUCTION

Although no theory has yet been develped that explains the cause of earthquakes in the Eastern United States, seismologists and engineers are still1

called upon to assess earthquake hazards in this region., As the trends of

urbanization and industrialization 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

their 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

parti cul ar.

(2) It is nevertheless incumbent upon seismologists to provide a practi -

cal guide for siting critical facilities that incorporates the present state of knowledge in the field.

"Seismic Ground Motion at Limerick. GeneratingStation," a report prepared

by ERTEC Rocky Mountain, Inc., is evaluated here 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

Limerick site in particular. On the other hand, despite this significant 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

at the Limerick Generating Station.

In particular, the results shown in Figure' 9 of the ERTEC report for the

"Decollement" hypothesis probably yield a 'reasonable estimate of seismic

ground motion at the site. This conclusion is ironic, since "Decollement"
B-3

.is possibly the most speculative of the four hypotheses considered.

Nonetheless, the practical application of "Decollement" is ultimately useful ,

since its essential feature (as far as the calculated seismic hazard is

concerned) is that it treats the entire Eastern seaboard as one seismogenic.

zone. This allows for the possiblity 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 Zones,

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 hazard~evaluation is to changes in the geometry of

seismogenic zones.

While these practical results can be gl~eaned from the ERTEC report,

Section 3 (Seisomogenic Zones) and Se'cion 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.

SEISOMOGENIC 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 earthquakes 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.
B-4

Decol 1 ement:

This hypothesis is based on an analysis of intensities reported for' the

1886 earthquake in Charleston, SC (Seeber and Armbruster, 1981) coupled wi th

results of'seismic reflection studies-of the deep crustal structure of the

southern Appalachians (Cook et al.,-1979). The seismic reflection profiles

have revealed a continuous shallow-dipping reflector beneath the southern

Appalachians that has been interpreted to be a maj~or decollement. The

inferred decollement has been proposed as the boundary of a sei-smically

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 et al., 1980) show a rather. low level of -seismicity in the Charleston

area, and the recent monitoring of the area with a dense -seismograph network

has also revealed a relatively low level of activity. Thus, studies of

microearthquake distribution, fault-plane solutions, and earthquake depth have

not been very abundant in this region (Hamilton, 1981). The hypothesis that

the current seismicity in the vicinity 6'f.Charleston, SC is occurring along a

major decollement surface is, therefore, not well supported by quantitati ve

seismological studies. The existence of an "Appalachian Detachment" should

thus be considered as interesting speculation, but speculation nonetheless.

Furthermore, although preliminary results from deep seismic reflection

profiles in the northern Appalachians (e.g., Ando et al., 198.1; Brown et al.,

1982) have also revealed shallow-dipping reflectors, the lateral extent 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 sourthern Appalachians, I have seen no- c onvincing

evidence to -suggest that this hypothesis should be applicable in the northern

Appalachians in general or in the vicinity of the Limerick site in particular.
B- 5

Figure 6 of the ERTEC report shows the northern boundary of the

Decollement zone at about 4140 . No reason for choosing this boundary was

given in the report.

Crustal Blocks:

According to this hypothesis, the occurrence of earthquakes in the Eastern

United States is controlled by large crustal blocks. Supposedly, the

boundaries of these blocks are seismically active and the interiors are

relatively inactive. While this hypothesis seems reasonable in principle, and

ma y eventually predict the locations of future large earthquakes, none of the

crustal block models that have been proposed (e.g., Diment et al., 1979)

correlate very well with historical or instrumentally located seismicity.

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

SeismicActivity 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 interpretation 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 rate 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
B-6

earthquakes in this region. Perhaps this approach should be referred to as

being "reasonable" rather than "realistic" I(see Table 1 of ERTEC report).

The ultimate test of such an approach to hazard assessment is, simply , an

accurate deterministic model of the causative mechanism of earthquakes in the

Eastern United States. At the present time, such a model is not available .

Thus, the assumptions used in the ERTEC report regarding rates of activity are

just that, assumptions.

Perhaps the historical earthquake activity in China studied by McGuire

(1979), for comparison with the Eastern United States, was anomalously

stationary due to a process that is at present unknown. Future investigators

may discover that 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-of plate boundaries; e.g., McCann et al ., 1979)

i's found to be applicable to intraplate earthquakes, then there might be long

periods of seismic quiesence premonitory to impending large earthquakes in

this region.

Does the rate of activity observed for the past 200 years in the East

represent an intraplate variation of a seismic gap, or is this rate a result

of many years of aftershocks of a large earthquake such as the New Madrid

event of 1811? Such questions can not be answered without an accurate

deterministic model of the cause of earthquakes in the East.

Maximum Magnitude:

It is not clear which hypotheses are being referred to in the ERTEC report

that restrict the *recurrence of Cape Ann, Massachusetts type earthquakes to

areas in New England; the author should have cited some references. I

suspect, however, that th e author is referring to an apparent association

between the northwest-southeast trend of seismicity in this region, and a

landward extension of the New England seamounts that was discussed by Diment
B3-7

et al. (1972), Sbar and Sykes (1973), and Fletcher et al. (1973). This trend

crosses the Ottawa-Bonnechere graben and Mesozoic intrusions that postdate the

initial separation of North America from Africa (Sykes, 1978). The

association between the trend of seismicity (the so-called "Boston-Ottawa

seismic belt") and these tectonic features (possible candidates for ancient

zones of weakness reactivated by the present-day stress field) has been

analyzed in detail by Sykes (1978). Further analysis of the correlation 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 Aggarwal, 1981). This gap (although not as

distinct) 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 A ggarwal , 1981). This observation suggests that earthquake

processes may be 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 origin. 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 shoul d

be.

LARGE EARTHQUAKES NEAR THE LIMERICK SITE

Appendix B of the ERTEC report discusses the credibility of hypotheses

that allow an earthquake of the size of the 1886 Charleston event to occu r in
B -8

the vicinity of the Limerick Generating Station. As stated in Appendix B,

calculations of the hazard at the site are sensitive to the subjective

probability assigned to such hypotheses. In the mai~n 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 earthquake the size of the Charleston event to occur at the Limerick site.

Since no explanation has been found for the cause of the 1886 Charleston

earthquake, there is no particular reason to exclude such an event from

anywhere along the 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 eastern Pennsylvania., Perhaps the

twenty-five to thirty percent probability for the scientific credibility of

such an hypothesis (as suggested by 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: i.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?

EARTHQUAKE 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

reporting and/or recording of events. While the lower bound Of mb=4.5 (MM

intensity V-VI) that was used for the part of the study estimating seismic

ground motion seems appropriate, it is not clear to what extent the

i ncompleteness of catalogues for smaller events could affect 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
B-9

serious evaluation of the completeness of the catalogues used is justifiably

beyond the scope of the study. Nonetheless, the report should state that completeness of catalogues could be a problem. This omission, again, creates an

impression that the phenomenon of Eastern United States earthquakes is better

understood than it really is.

SUMMARY AND CONCLUSIONS

The general writing style of "Seismi~c Ground Motion at the Limerick

Generating Station," a report prepared by ERTEC Rock 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 seismogenic zones "within which earthquakes are considered to

be a 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:

* The conclusion derived from studies by McGuire (1977), McGuire (1979),

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.

* 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 and all hypothesis that allows an earthquake the size of the

1886 Charleston event to occur in eastern Pennsylvania. This probability, suggested by at least one of the experts consulted in Appendix B,

is not unreasonable,.
B-10

*There is so 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, t he 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 earthquakes in the East. In

assessing the seismic hazard it is useful to know how sensitive the resulti ng

hazard evaluation is to changes in the geometry of seismogen~ic zones. This i s

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

yeild a reasonable estimate of the maximum seismic ground motion to be expected at the Lime-rick site. This conclusion is ironic, since "Decollement""

is possibly the most speculative of the four hypotheses considered. Nonetheless, the practical application of "Decollement" is ultimately useful,

since its essential feature (as far as calculated ,Peismic 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

reasonable estimate of the seismic hazard at the Limerick Generating Station .

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, EQS. Trans., Am. Geophysical Union, 62, No.45,

p.1046, 1981.
B-11

Barstow, N. L. , Brill1, K. G. , Nuttl i, 0. W. , and Pomeroy, P. W. , An approach

to seismic zonation for siting nuclear electric power generating

facilities in the eastern U.S., USNRC, NUREG/CR-1577, 1980.

Brown, L., Ando, C., Klemperer, S., Oliver, J. E., Kaufman, S., Czuchra, B.,

Walsh, T., and Isachsen, Y., Adirondack-Appalachian crustal structure:

The COCORP Northeast Traverse, EQS. 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., Oliver, J., and Hatcher,

R., Thin-skinned tectonics in the crystalline southern Appalachians:

COCORP seismic reflection profiling of the Blue Ridge and Piedmont, Geology, 7, p.563-567, 1979.

Diment, W. H., Muller, 0. G., and Lavin, P. M., Basement tectonics of New York

and Pennsylvania as revealed by gravity and magnetic studies, in

Caledonides on the USA, published by Virginia Poly. Inst. and State

Univ., Blacksburg, VA, 1979.

fliment, W. G., Urban, T. C., and Revetta, F. A., Some geophysical anomalies

in the e'astern 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.

Hamilton, R., Geological origin of eastern U. S. seismicity, in Earthquakes

and Earthquake Engineering, Eastern United States, Vol.1, Ed. J. E. Beavers, p.3-24, 1981.
B-12

McCann, W. R., Nishenko, S. P., Sykes, L. R., and Krause, J., Seismic gaps and

plate tectonics: Seismic potential for major plate boundaries, Pure and

Appl. Geophys., Vol.117, p.1083-1147, 1979.

McGuire, R. K., Effects of uncertainty in seismicity on estimates of seismic

hazard for the east coast of the United States, Bull. Seis. Soc, Am.,

Vol .67, No.3, p.827-848, 1977.

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, R. 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 seismicity

in eastern North America: An example of intraplate tectonics, Geol. Soc.

Amn. Bull., Vol.84, p.1861-1882, 1973.

Seeber, L., and Armbruster, J. G., The 1886 Charleston, South Carolina

earthquake and the Appalachian detachment, Journ. Geophys. Res., Vol .86,

No.B9, p.7874-7894, 1981.

Sykes, L. R., Intraplate seismicity; reactivation for preexisting zones of

weakness, alkaline magmatism, and other tectonism postdating continental

fragmentation, Rev. Geophys. and Space Phys., Vol.16, p.621-688, 1978.

Yang, J. P., and Aggarwal, Y. P., Seimotectonics of the northeastern United

States and adjacent Canada, Journ.* Geophys. Res., Vol.86, No.B6,

p.4981-4988, 1981.
MAC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER fAssigned by TIOC, add Vol. No, ji*,s~)

(2-84)

30,22 BIBLIOGRAPHIC DATA SHEET NUREG/CR- 3493

SEE INSTRUCTIONS ON THE REVERSE. BNL-NUREG-51711

2, TITLE AND SUBTITLE 3, LEAVE BLANK

A Review of the Limerick Generating Station

Severe Accident Risk Assessment_________________ 4, DATE REPORT COMPLETED

Review of Core-Melt Frequency MONTH Y EAR

S.ATOI) M.A. Azarm, R.A. Bari, J.L. Boccio, N. Hanan, July 1984

I.A. Papazoglou, C. Ruger, K. Shitj, J. Reed,6.DTREOTISD

M. ManA ak.MONTH I YEAR

7. PERFORMING ORGANIZATION NAME AND MAILING ADDRESS (hicl~de Zp Code) B, PRJET/AS /WR 1N 984E

Brookhaven National Laboratory 9. FIN OR GRANT NUMBER

Upton, NY 11973

A.-3 393

10. SPONSORING ORGANIZATION NAME AND MAILING ADDRESS (include Zip Code) 1Ila. TYPE OF REPORT

Division of Safety Technology

Office of Nuclear Reactor Regulationb.PROCVED nls dr)

U.S. Nuclear Regulatory Commission,b.PROCVED sc,,edl)

Washington, D.C. 20555

12. SUPPLEMENTARY NOTES

Pertains to Docket N6. 50-352 and 50-353.

13. ABSTRACT (200 eoOds or !e,)

A limited 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 frquencies

and their impacts on the plant components and structures. Particular

attention is given to uncertainties and critical assumptions. Limited

requantification is performed for selected core-melt accident sequences

in order to illustrate sensitivities of the results to the underlying

assumptions.

14. DOCUMENT ANALYSIS - . KEYWORDS/DESCRIPTDRS .18. AVAILABI1LITY

Limerick Generating Station STATEMENT

Core-melt frequency Unl imited

Severe Accident Risk Assessment________

Seismic- and Fire-Initiating Events 16. SECURITY CLASSIFICATION

(This Page)

b. IDENTIFIERSIOPEN-ENDED TERMS Unclassified

fThismrpore)

Unclassified

17. NUMBER OF PAGES

18 PRICE
I'

1~