NUREG/CR-5176

UCID-21425

Seismic Failure and Cask Drop

Analyses of the Spent Fuel

Pools at Two Representative

Nuclear Power Plants

Prepared by P. G. Prassinos, C. Y. Kimura, D. B. McCallen, R. C. Murray/LLNL

M. K. Ravindra, R. D. Campbell, P. S.. Hashimoto, A. M. Nafday, W. H. Tong/EQE

Lawrence Livermore National Laboratory

EQE Engineering, Inc.

Prepared for

U.S. Nuclear Regulatory

Commission
NOTICE

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NUREG/CR-5176

UCID-21425

R4, 9H

Seismic Failure and Cask Drop Analyses of the Spent Fuel

Pools at Two Representative Nuclear Power Plants

Manuscript Completed: June 1988

Date Published: January 1989

Prepared by

P. G. Prassinos, C. Y. Kimura, D. B. McCallen, R. C. Murray, Lawrence Livermore National Laboratory

M. K. Ravindra, R. D. Campbell, P. S. Hashimoto, A. M. Nafday, W. H. Tong, EQE Engineering, Inc.

Lawrence Livermore National Laboratory

7000 East Avenue

Livermore, CA 94550

Subcontractor:

EQE Engineering, Inc.

3150 Bristol, Suite 350

Costa Mesa, CA 92626

Prepared for

Division of Safety Issue Resolution

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

NRC FIN A0814

ABSTRACT

This report discusses work done in support of the resolution of Generic Issue82, "Beyond Design Basis Accidents in Spent Fuel Pools". Specifically the

probability of spent fuel pool failure due to earthquakes was determined for

the pools at the Vermont Yankee Nuclear Power Station (BWR) and the H. B.

Robinson S.E. Plant, Unit 2 (PWR). The dominant failure mode for each pool

was gross structural failure caused by seismic motion resulting in the loss of

pool liner integrity. The resulting sudden loss of water was then assumed to

lead to a self-propagating cladding failure and fission product inventory

release from the spent fuel elements in the pool. The mean annual frequency

of failure due to this failure mode was found to be 6.7E-06 at Vermont Yankee

and 1.8E-06 at H. B. Robinson. Other earthquake induced failure modes studied

but found to be less important were loss of pool cooling and make-up

capability, fuel rack damage and loss of liner integrity due to a cask drop

accident.

iii

TABLE OF CONTENTS

Page

ABSTRACT

LIST OF TABLES vii

LIST OF FIGURES viii

ACKNOWLEDGEMENT xi

EXECUTIVE SUMMARY xiii

Chapter 1. INTRODUCTION

1.1 Background 1-1

1.2 Objective of this Study 1-4

1.3 Report Organization 1-4

Chapter 2. SEISMIC RISK ANALYSIS METHODOLOGY

2.1 Introduction 2-1

2.2 Seismic Hazard Analysis 2-1

2.3 Seismic Fragility Evaluation 2-2

2.3.1 Failure Criteria for Spent Fuel Pools 2-3

2.4 Analysis of Systems and Accident Sequences 2-9

2.5 Evaluation of Accident Frequency 2-11

Chapter 3. SEISMIC FRAGILITY OF POOL STRUCTURES

3.1 Introduction 3-1

3.2 Vermont VYNPS Spent Fuel Pool (SFP) 3-1

3.2.1 VYNPS SFP Structure Description 3-1

3.2.2 VYNPS SFP Structure Evaluation 3-1

3.2.3 Conclusions of the Structural Analysis 3-5

3.3 H. B. Robinson, Unit 2 (HBR2) Spent Fuel Pool 3-6

3.3.1 HBR2 SFP Structure Description 3-6

3.3.2 HBR2 SFP Structure Evaluation 3-6

3.3.3 Conclusions of the Structural Analysis 3-10

Chapter 4. SPENT FUEL POOL SYSTEMS ANALYSIS

4.1 Analysis of the VYNPS SFP Systems 4-2

4.1.1 VYNPS SFP Systems Description 4-2

4.1.2 VYNPS SFP Systems Analysis 4-3

4.2 Analysis of the HBR2 SFP Systems 4-9

4.2.1 HBR2 SFP Systems Description 4-9

4.2.2 HBR2 SFP Systems Analysis 4-11

V
Chapter 5. EQUIPMENT FRAGILITIES

5.1 VYNPS Equipment

5.1.1 VYNPS Fuel Racks

5.1.2 VYNPS Cooling and Makeup Systems

5.2 HBR2 Equipment

5.2.1 HBR2 Cooling and Makeup Systems

Chapter 6. RISK EVALUATION

6.1 Objective

6.2 VYNPS Analysis

6.2.1 VYNPS Assumptions

6.2.2 VYNPS Accident Sequences and Failure States

6.2.3 VYNPS SFP Case Studies

6.2.4 Probability of VYNPS Spent Fuel Pool Failure

6.3 HBR2 Analysis

6.3.1 Probability of HBR2 Spent Fuel Pool Failure

Chapter 7. CASK DROP ANALYSIS

7.1 Background

7.2 Analysis Methodology

7.3 Cask Drop Analysis of Vermont Yankee.Fuel Pool

7.3.1 Finite Element Model

7.3.2 110 Ton Cask Drop

7.3.3 40 Ton Cask Drop

7.4 Cask Drop Analysis of H.B. Robinson Fuel Pool

7.4.1 Finite Element Model

7.4.2 Results of Analysis

7.5 Summary of Observations

5-1

5-1

5-2

5-4

5-4

6-1

6-1

6-1

6-2

6-2

6-3

6-3

6-3

7-1

7-1

7-3

7-3

7-4

7-4

7-5

7-5

7-5

7-5

8-1

8-2

9-1

A-1

B-1

C-1

Chapter 8. SUMMARY AND CONCLUSIONS

8.1 Summary

8.2 Conclusions

Chapter 9.

Appendix A

Appendix B

Appendix C

REFERENCES

Vermont Yankee Nuclear Power Station Spent Fuel

Pool Thermal Analysis

Vermont Yankee

Fault Trees

Nuclear Power Station Spent Fuel Pool

H. B. Robinson Unit 2 Spent Fuel Pool Fault Trees

vi
LIST OF TABLES

Page

2.1 Probability of Exceedance at Different Acceleration Levels 2-13

3.1 Median Factors of Safety and Variabilities of Spent Fuel

Structure at Vermont Yankee 3-11

3.2 Median Factors of Safety and-Variabilities of Spent Fuel Pool

Structure at H.B. Robinson 2 3-12

4.1 Components of VYNPS Spent Fuel Pool Systems 4-15

'4.2 VYNPS SFP Component Failures 4-17

4.3 HBR2 Nuclear Power Plant Component List 4-18

4.4 HBR2 SFP Component Failures 4-24

5.1 Summary of Equipment Fragilities for VYNPS SFP 5- 7

5.2 Median Factors of Safety and Variabilities for HBR2 CST 5- 8

5.3 Summary of HBR2 Equipment Fragilities 5- 9

6.1 Seismic Fragilities and Random Failure Rates' for VYNPS 6- 5

6.2 Results of the Analyses of VYNPS SFP 6- 6

6.3 Seismic Fragilities for HBR2 6- 7

6.4 Results of the Analyses of HBR2 SFP 6- 8

vii
LIST OF FIGURES

2-I Sample Seismic Hazard Curves

2-2 Example Fragility Curves for a Structure

2-3 Failure Mechanism for the Pool Floor

2-4 Failure Mechanism for the Pool Wall

2-5 Spent Fuel Storage Rack Schematic (10 x 10)

2-6 Typical Rack Elevation View

2-7 Pool Layout

2-8 Plan Fuel Storage Rack Arrangement

2-9 Fuel Pool Arrangement for Laterally Braced High Density

2-10 Typical Elevation - All Groups

3-1 Reactor Building Plan at Elevation 303'-0"

3-2 Cross-Sections of the Spent Fuel Pool

3-3 Reactor Building OBE Floor Spectrum, EL 303'-0" N-S Dir4

3-4 Reactor Building OBE Floor Spectrum, EL 303'-0" E-W Dir'

3-5 Reactor Building OBE Floor Spectrum, EL 303'-0" Vert. D

3-6 Spent Fuel Pool Structure Fragility

3-7 HBR2 Spent Fuel Pool Slab at EL 236'-9"

3-8 Cross-Section of HBR2 Spent Fuel Pool

3-9 HBR2 Spent Fuel Pool Structure Seismic Fragility Curves

4-I Fuel Pool Cooling System

4-2 Fuel Pool Filter Demineralizer System

4-3 VY Simplified Reactor Building Closed Cooling Water Sys

4-4 VY Simplified Reactor Building Closed Cooling Water Sys

Loop 1 & 2

4-5 VY Condensate Transfer System

4-6 VY Demineralized Water Transfer System

4-7 VYNPS Event Tree

4-8 VYNPS Fault Tree for Pool State 1

4-9 VYNPS Fault Tree for Pool State 2

4-10 H.B. Robinson Spent Fuel Pool Event Tree

4-11 Spent Fuel Pit Cooling

6-1 Seismic Hazard Curves for the VYNPS Site

6-2 Seismic Hazard Curves for the HBR2 Site

7-1 Finite Element Model of Wall and Canister

7-2 Dynamic Response Resulting from Impact

7-3 Bi-Axial Failure Envelope for Concrete

7-4 Estimated Region of Tensile Cracking in Concrete

7-5 Estimated Region of Vertical Steel Yielding

7-6 Estimated Region of Horizontal Steel Yielding

7-7 Yielding of Horizontal Steel

7-8 Yielding of Vertical Steel

7-9 Dynamic Response Resulting from Impact

7-10 Yielding of Vertical Steel

7-11 Yielding of Horizontal Steel

7-12 Finite Element Model of the Canister and Wall System

Fuel Racks

ection

ection

irection

tem

tem,

Page

2-14

2-15

2-16

2-17

2-18

2-19

2-20

2-21

2-22

2-23

3-13

3-14

3-15

3-16

3-17

3-18

3-19

3-20

3-21

4-25

4-26

4-27

4-28

4-29

4-30

4-31

4-32

4-33

4-34

4-35

6- 9

6-10

7- 7

7- 8

7- 9

7- 9

7-10

7-10

7-11

7-12

7-13

7-14

7-15

7-16

viii
Page

7-13 Time History Response of the Wall. 7-17

7-14 Estimating Damage to Vertical Reinforcing Steel 7-18

7-15 Estimating Damage to Horizontal Reinforcing Steel 7-19

ix

ACKNOWLEDGEMENT

This project was performed in cooperation with the U.S. Nuclear Regulatory

Commission. We greatly appreciate the support and guidance by Ed Throm,

Office of Nuclear Regulatory Research, Division of Reactor and Plant

Systems. This assistance was essential to the completion of this project.

We wish to thank the utility participants for their cooperation in providing

information on their plants.

xi

EXECUTIVE SUMMARY

Previous work [Sailor, et al., January 1987] in support of the resolution of

Generic Issue-82, "Beyond Design Basis Accidents in Spent Fuel Pools",

indicated that the risk due to seismic induced failure of nuclear power plant

spent fuel pools (SFP) may be significant and dominates the total risk

resulting from SFP failure. This result was based on use of generic seismic

hazard and fragility data and resulted in calculated risk numbers with wide

uncertainty ranges. To test the result, Lawrence Livermore National

Laboratory was asked to determine seismically induced spent fuel pool failure

probabilities for two specific plants, the Vermont Yankee Nuclear Power

Station, a boiling water reactor,'and the H. B. Robinson S.E. Plant, Unit 2, a

pressurized water reactor. Plant specific seismic hazard and fragility

information was used for each site. The results of this work is presented in

the report which follows.

The major concern about accidents in SFP's is the loss of water inventory and

its capability to cool the radioactive fuel. Without sufficient water, some

theoretical models suggest that the fuel's zircaloy cladding may initiate and

sustain rapid oxidation (fire) that can spread to adjacent fuel assemblies

(self-propagating) with the potential of releasing significant amounts of

long-lived radioactive isotopes. Although these isotopes may not present a

significant acute health hazard, their release could cause contamination of

surrounding property comparable or larger than would occur from a reactor core

melt accident.

The failure modes considered in this study were: (1) loss of liner integrity

precipitated by gross structural failure of the SFP, (2) loss of function of

the fuel pool support system (pool cooling and make-up) resulting in loss of

water inventory, (3) damage to fuel racks, and (4) loss of liner integrity due

to a cask drop accident. The first failure mode turned out to be the dominant

contributor to risk. The second failure mode, loss of pool cooling and makeup capability, was discounted since 3 to 7 days would be available to find

alternate water sources before fuel damage could occur. The third failure

mode was discounted after analysis showed that crushing of the fuel in the

fuel racks could not result in a criticality accident. The fourth failure

mode was discounted because the probability of a cask drop is believed low

(<1.OE-08/RY) as reported in the Sailor study. If this is not the case, the

cask drop accidents would be important since a dropped cask probably would

result in loss of liner integrity.

The first failure mode, gross structural failure of the pool, was analyzed by

first identifying the potential locations and issues for gross failure. In

the Vermont Yankee pool the out-of-plane shear failure of the pool slab

controlled. In the H. B. Robinson pool the out-of-plane bending of the pool

was wall controlled. It was found that the median acceleration capacity was

1. 4g and 2.Og at each pool respectively. The High Confidence of Low

Probability of Failure capacity (5% chance of failure with 95% confidence) was

0.50g and 0.65g respectively. Since the Safe Shutdown Earthquake at each

plant is 0.14g and 0.20g respectively, it can be seen that there is a large

safety margin inherent in each pool design. The mean annual frequency of

failure was calculated as 6.7E-06 for Vermont Yankee and 1.8E-06 for H. B.

Robinson. This compares with values of 2.2E-05 for a BWR and 1.6-05 for a PWR

found in the Sailor study.

xiii

1. INTRODUCTION

Increasing amounts of spent fuel are being stored at nuclear power plants

rather than being shipped to reprocessing facilities or disposal

repositories. This increasing amount of on-site spent fuel storage has

resulted in much larger inventories in spent fuel storage pools than

previously anticipated. While the radioactive material inventory of the spent

fuel will be less than an active core, there is a potential for accidents in

spent fuel pools involving the equivalent of several cores of spent fuel

bundles.

The major concern is for the spent fuel pool to lose its water inventory and

thus its capability to cool the spent fuel inventory. Without sufficient

water inventory, a theoretical model suggested that the zircaloy cladding

material of recently discharged spent fuel rods can initiate and sustain a

rapid oxidation (fire) that can spread to adjacent fuel assemblies with the

potential of releasing significant amounts of long-lived radioactive isotopes.

The overall objective of this report is to present the results of

probabilistic analyses of the seismic failure of two specific spent fuel

pools. The spent fuel pools considered are at the Vermont Yankee Nuclear

Power Station, a General Electric Boiling Water Reactor (BWR) located near

Vernon, Vermont, and the H. B. Robinson S. E. Plant, Unit 2, a Westinghouse

Pressurized Water Reactor (PWR), located in Hartsville, South Carolina.

1.1 Background

An analysis of the risk associated with spent fuel storage was performed in

the Reactor Safety Study (RSS) [USNRC, 19751. This analysis indicated that

the risk associated with spent fuel pool accidents was small compared to the

risk from accidegts involving the reactor core with failure probabilities in

the range of 10- (denoted E-6) per reactor-year. The RSS analysis included

external initiating events and heavy load drops. Release estimates were based

on accidents involving one-third of a core of spent fuel. The overall

conclusion in the RSS was that loss of the spent fuel pool water inventory,

and radioactivity release was unlikely based on design and operational

features.

Subsequently, because of the increased storage of spent fuel at nuclear power

plants and the possibility of zircaloy fire ignition and propagation, the NRC

sponsored a study of Generic Safety Issue 82, "Beyond Design Basis Accidents

in Spent Fuel Pools" [Sailor, et al., January 19871. This study revisited the

RSS analysis to see if there was a possibility for new events or more severe

consequences. The study involved a probabilistic analysis of the risk

associated with spent fuel pool failures initiated by random system failures,

seismic events and the dropping of heavy loads, primarily a spent fuel

shipping cask, on the pool wall. The analysis considered initiating event

frequencies, spent fuel pool systems response, accident phenomenology such as

cladding fires, and consequences such as radioactivity release, transport,

deposition and health effects.

1-1
The accident initiating events that were considered in the BNL study included:

- pool heat-up due to loss of cooling water inventory and circulation

capability,

- structural failure of the pool due to seismic events or missiles,

- partial drain down of the pool due to pneumatic seal failure, and

- structural failure of the pool due to heavy load drop.

Data and information from several plant sites including Ginna, Millstone 1,

Zion and Oyster Creek were used in the spent fuel pool analyses. These plants

were chosen specifically for their perceived Vulnerability, being designed to

less restrictive requirements than presently employed, in particular to

seismic events.

Estimates of spent fuel pool cooling system failure by BNL were derived from

previous probabilistic analyses for "trained" (parallel) systems, and included

component failures, and unavailabilities. Consideration was also given for

the capability of several alternative systems to supply makeup water to the

pool. The frequency for failure of fuel pool cooling leading to uncovering

the spent fuel and to radioactivity release was estimated to be about E-6 per

reactor-year for both PWRs and BWRs.

The BNL analysis of the spent fuel pool structural failure utilized the

seismic hazard representation for the Millstone site taken from the EUS

Seismic Hazard Characterization Study [Bernrueter, et al., 1985, 1987]. Since

no published seismic hazard study existed for the Ginna site, the Millstone

seismic hazard curves were scaled by the Ginna SSE peak ground acceleration at

200, 1,000, and 4,000 year return periods. The analysis for Millstone also

used seismic hazard curves at the same return periods. For the Millstone

analysis, four hazard curves (best estimate, median, 15% and 85% confidence

curves) were used. For the Ginna analysis, only one hazard curve was used.

In the BNL study [Sailor, January 1987], the seismic fragility of pool

structures was assigned by comparing with published PRAs. Since Millstone 1

is a BWR, the seismic fragility developed for the Oyster Creek reactor

building was judged appropriate for the Millstone 1 spent fuel pool. For the

Ginna spent fuel pool, the fragility of the Zion plant auxiliary building

shear walls was used.

Use of these representative fragilities and the assigned hazard curves,

resulted in a large range of values for seismically induced spent fuel pool

structural failure. For a PWR plant, the annual frequency of seismically

induced gross structural failure was given as 1.6E-11 to 2.6E-4 per reactor

year. For a BWR plant, this range was given as 4.5E-11 to 6.5E-4 per reactor

year. The mean frequency was estimated by BNL to be 2.2 E-5 per reactor year

for a BWR and 1.6 E-5 per reactor year for a PWR.

1-2
The structural failure of the spent fuel pool due to dropping of a spent fuel

transfer cask has been considered by the NRC as an Unresolved Safety Issue -

TAP A-36, "Control of Heavy Loads at Nuclear Power Plants". The NRC has

concluded in Generic Letter 85-11, that all licensees meet A-36 criteria "such

that the risk associated with potential heavy load drops is acceptably small",

with consideration for:

o single failure cranes,

o appropriate handling procedures and training, and

o analysis of heavy load structural and mechanical

capacities.

The BNL study also considered spent fuel pool structural failure due to

dropping of a spent fuel shipping cask. This analysis considered the

anticipated frequency of spent fuel shipment when a repository becomes

available, the time the cask would be over the pool wall, crane failure, and

human error in properly rigging, lifting and moving the spent fuel shipping

cask. In addition, it was assumed that the conditional probability of

structure failure of the spent fuel pool wall given a cask drop has a value of

0.1 with an uncertainty range of 0.01 to 1. The results of this analysis

indicates that failure of the spent fuel pool due to cask drop accidents has a

range of 3.E-12 to 3.E-5 per reactor year.

It was concluded from the BNL study that the dominant contributors to risk are

structural failures resulting from beyond design basis seismic events and from

heavy load drops. Large uncertainties were associated with both these

events. The uncertainty in the seismic failure of the pool resulted from

uncertainties in the seismic hazard curves and spent fuel pool fragilities,

and the use of generic information and data. The uncertainty in pool failure

from heavy load drops resulted from Uncertainties in human error rates and the

conditional probability of pool structural failure from shipping cask impact.

The estimate of the overall risk due to beyond design basis accidents in spent

fuel pools for seismic events at the PWR surrogate plant ranges from 600

person-rem/Ry to a negligible value with an estimated mean value of about 50

person-rem/Ry in the BNL study. The estimates for the BWR surrogate plant

were about a factor of three less. The overall risk due to a heavy load drop

event was estimated to range from 70 person-rem/Ry to a negligible value for

the PWR surrogate plant and also a factor of about three less for the BWR

surrogate plant. The unique character of spent fuel pool accidents

(potentially large release of long-lived isotopes) makes it difficult to

compare directly to reactor core melt accidents. There are no early health

effects associated with spent fuel pool accidents. Land interdiction area is

postulated to be larger for spent fuel pool accidents, as compared to reactor

core melt accidents.

The large uncertainties and the use of generic information and data (seismic

hazards, fragilities and systems descriptions) in the previous analyses has

1-3
provided the impetus for performing this present study of specific spent fuel

pools. This study, while providing results for specific plants, is being

performed to gain insight into representative values for the probability of

spent fuel pool failure.

1.2 Objective of this Study

The purpose of this effort is to provide technical assistance to the NRC

concerning Generic Issue-82, "Beyond Design Basis Accidents in Spent Fuel

Pools". The overall objective is to estimate the realistic seismic capacity

of spent fuel pool structures and systems, and to assess the seismic

contribution to risk from accidents in spent fuel pools. This objective is

accomplished by using plant specific analyses of actual spent fuel pools and

site specific hazard analyses. This effort includes:

1. Establish an appropriate definition of spent fuel pool failure.

2. Evaluate the seismic capacity of a typical BWR and PWR spent fuel

pool, and determine their seismic fragility and failure probability.

3. Evaluate the effect of dropping a spent fuel shipping cask on the pool

wall.

1.3 Report Organization

The remainder of this report is organized into six chapters. Chapter 2

present a discussion of the methodology utilized in the spent fuel pool

analysis followed by the fragility analysis of the spent fuel pool structure

and equipment presented in Chapter 3. The systems analysis of the spent fuel

pool cooling and make-up water systems is provided in Chapter 4, and Chapter 5

presents the equipment fragilities. The result of probabilistic evaluation

are presented in Chapter 6. Chapter 7 presents the analysis of the cask drop,

and the summary and conclusions are presented in Chapter 8.

1-4
2. 'SEISMIC RISK'ANALYSIS METHODOLOGY

2.1 Introduction

Seismic risk analysis methodology for reactor core damage accidents has now

been well established, with its application to over 25 nuclear power plants

and standardized by NUREG/CR-2300 [USNRC, 1983]. The objectives of such an

overall probabilistic risk assessment are to estimate the frequencies of

occurrence of severe core damage,.serious radiological releases, and

consequences in terms of early fatalities, long term adverse health effects

and property damage, and to identify significant contributors to plant

risks. The key elements of such an analysis are:

1. Seismic hazard analysis -estimation of the frequency of.variou.s

levels of seismic ground motion (acceleration) occurring:at the site.

2. Seismic fragility evaluation - estimation of the conditional

probabilities of structural or equipment failure for given levels of

ground acceleration.

3. Systems/accident sequence analysis - modeling of the various

combinations of structural and equipment failures that could initiate

and propagate a seismic core damage accident sequence.

4. Evaluation of core damage frequency and public risk -. assembly of the

results of the seismic hazard, fragility, and systems analyses to

* estimate the frequencies of core damage and plant damage states.•

Assessment of the impact of seismic events on the containment;.and

consequence analyses, and integration of these results with the core

damage analysis to obtain estimates of seismic risk in terms of

effects on public health..

This methodological sequence is modified by the use of a screening approach

based on component capacity to address the specific issue at hand,> which is

estimating the failure frequency of spent fuel pools initiated by earthquake

events. It is described in the subsequent sections.

2.2 Seismic Hazard Analysis

Seismic hazard is usually expressed in terms of the frequency distribution of

the peak value of a ground-motion parameter (e.g., peak ground acceleration)

during a specified time interval. The different steps of this analysis are as

follows:

1. Identification of the sources of earthquakes, such as faults and

seismotectonic provinces.

2. Evaluation of the earthquake history of the region to assess-the

frequencies of occurrence of earthquakes of different magnitude's or

epicentral intensities.

2-1
3. Development of attenuation relationships to estimate the intensity of

earthquake-induced ground motion (e.g., peak ground acceleration) at

the site.

4. Integration of the above information to estimate the frequency of

exceedance for selected ground motion parameters.

The hazard estimate depends on uncertain estimates of attenuation, upperbound

magnitudes, and the geometry •of the postulated sources. Such uncertainties

are included in the hazard analysis by. assigning probabilities to alternative

hypotheses about these parameters.. A probability distribution for the

frequency of occurrence is thereby developed. The. annual frequencies for

exceeding specified values of the ground motion parameter are displayed as a

family of curves with different probabilities. Sample hazard curves are shown

in Figure 2-1.

2.3 Seismic Fragility Evaluation

Seismic fragility evaluation has been described in the PRA Procedures Guide

[USNRC, 1983] and by [Kennedy and Ravindra, 1984]. The seismic fragility of a

spent fuel pool is the conditional probability of its failure for a given peak

ground acceleration. In the seismic probabilistic risk assessment studies, the

seismic.fragility is expressed in terms of the peak ground acceleration

capacity. Typically, seismic fragility is characterized by means of a family

of curves, each curve showing the conditional probability of failure as a

function of the peak ground acceleration. The family of curves is necessary

to represent the uncertainties in the failure modes and model parameters.

Usually a subjective probability or weight is assigned to each curve to

reflect the degree of belief associated with each hypothesis (i.e., a set of

model parameters).

In the seismic PRAs, the entire family of fragility curves for a structure or

equipment is developed by means of a. simple fragility model wherein the

capacity is represented by means of three fragility parameters: median

acceleration capacity, AmY and the logarithmic standard deviations PR and PU

representing, respectively, the random variability and the uncertainty in

median capacity. These parameters are estimated using the margins and

variabilities existing in the nuclear power plant seismic design and

qualification procedures. Figure 2-2 shows the seismic fragility curves for a

structure where the middle one is the median fragility curve and the outer two

show the fragility at 95% and 5% confidence.

A useful indicator of the seismic capacity presently being used in seismic

margin studies is the High Confidence Low Probability of Failure (HCLPF)

capacity of the structure which is defined as the acceleration value at which

we can state with 95% confidence that the conditional probability of failure

will not exceed 5%. The example in Figure 2-2 shows the median capacity of

the structure as 0.90g, the HCLPF capacity as 0.30g and the safe shutdown

earthquake as 0.17g. Therefore, the HCLPF capacity may be taken as the maximum

2-2
ground level motion that the structure could survive with a high degree of

confidence.

The 'starting point in estimating the fragility of a spent fuel pool structure

is to define what is considered to be failures and to identify the critical

failure modes. Guidance on the relevant failure modes is obtained by USNRC

and ANS standards [USNRC 1981; ANS 1976], although the failure of the

structures and equipment beyond the design basis earthquake is of concern

here. For each failure mode, the fragility'or equivalently the ground

acceleration capacity is estimated. The structure is modeled to fail if any

one of the 'failure modes occur and the fragility of the structure is obtained

as the conditional probability of occurrence of the union of these failure

modes. The failure modes of the spent fuel pool are discussed later in this

report.

For each failure mode, the ground acceleration capacity is expressed as the

product of the SSE acceleration times an overall factor of safety. The

overall factor of safety is further modeled as a product of the factors

reflecting the conservatisms and uncertainties in different variables such as

spectral shape, damping, mode combination, earthquake component combination,

material strength and ductility. Each factor is represented by its median

value, A, and 3.U* Using the lognormal model, the parameters of the

overall ;actor and of the ground acceleration capacity are calculated knowing

the parameters of the individual factors. The HCLPF capacity is expressed as

HCLPF Capacity = Am exp [-1.64 R + U

This fragility model and development of fragility parameters have been

utilized in over 25 seismic PRAs. For further details, the reader is referred

to [Kennedy and Ravindra 1984] and [Ravindra et. al 1987].

2.3.1 Failure Criteria for Spent Fuel Pools

Before the seismic fragility of a structure or equipment is estimated, the

potential failure criteria need to be established.

There are several different spent fuel pool designs and fuel rack designs.

The fuel rack designs depend on the vintage, fabricator and pool

configuration.

In analyzing the failure of the spent fuel pool structures and systems, we

considered the following:

o Loss of liner integrity precipitaued by structural failure of the

spent fuel pool

o Loss of function of the fuel pool support systems (e.g., pool cooling

and make-up water capacity) resulting in loss of water inventory

through boil-off or drainage

2-3
o Damage to fuel racks caused by fuel rack motion.

The loss of spent fuel pool water inventory either through a structural

failure or a failure in the support systems can result in degradation to

stored spent fuel and possibly zircaloy cladding fires. For the purpose of

this analysis, spent fuel pool failure and its probability of occurrence

(risk) is considered when;

" Spent fuel is uncovered long enough for a self-sustaining zircaloy

cladding oxidation initiation (-1650 deg. F), or

o Spent fuel is uncovered long enough for the zircaloy cladding to be

damaged (breached) and releases some or all of its gaseous and

volatile fission inventory. Sufficient fission products are released

to exceed 10 CFR 20 guidelines at the site exclusion zone boundary.

In the following sections we describe the different structures and equipment

comprising these systems and discuss the more credible failure modes under

earthquakes.

2.3.1.1 Spent Fuel Pool Structures

2.3.1.1.1 BWR Mark I & Mark II Plant Fuel Storage Pool

In BWR Mark I and Mark II containment 'system, the spent fuel storage pool is

located at the operating floor level (100'-150' above grade). The fuel

storage pool is typically -80' long x 40' wide x 44' deep. The thickness of

the pool floor and walls is on the order of 4'-6'. The horizontal and

vertical loads from the pool floor are transmitted to the two longitudinal

walls which are designed as deep girders supported at the peripheral wall of

the reactor building. The pool floor and walls are designed for dead and live

loads, hydrostatic pressure loads, seismic loads, hydrodynamic loads, thermal

loads and loads resulting from accidental drop of casks or heavy objects.

The critical locations for the design of the pool floor are at the mid span of

the floor for bending moments and at the intersection of pool floor and wall

for negative bending moment and shear loads. The critical locations for the

design of the pool walls are the intersection of pool walls, intersection of

the pool wall and floor and mid span of the walls.

Typical loads for the design of the fuel storage pool are:

Pool Floor = 9 kips/sq. ft.

Pool Walls = 100 kips/ft.

Failure mechanisms that need to be considered are given below:

1. The failure mode for the pool floor is that of a slab fixed at four

edges. Figure 2-3 shows the yield lines for the floor. Figure 2-4

shows the yield lines for the out of plane failure of the pool

wall. The girders supporting the pool are in reality long walls

2-4
acting as deep girders and are supported by the peripheral walls of

the reactor building.

2. Compressive and shear stresses at the reaction points of the girders

(onto the reactor building walls) for transmitting vertical and

horizontal seismic loads from the fuel storage pool to the foundation

should be investigated for structural adequacy.

3. Due to large concentrated loads (50 kips to 70 kips) at each foot of

the storage rack, bearing and punching shear stress in the pool floor

should also be investigated.

4. For laterally braced high density fuel storage racks, large

concentrated loads are transmitted to. the pool wall at either the

base level or at the base and the upper seismic bracing level. The

effect of concentrated load should be investigated.

Although the thermal loads are important in the design of the fuel pools,

their influence on the fragility of the pool is judged not significant because

the thermal loads are self-relieving.

2.3.1.1.2 PWR and BWR Mark III Fuel Storage Pools

The fuel storage pools for the PWR and BWR Mark III plants are located

typically at the ground level. Due to lower elevation, the seismic response

is relatively low. The vertical and horizontal loads from the pool floor are

transmitted to the ground for the case of free standing or floor mounted fuel

storage racks. For the case of laterally braced racks, the horizontal seismic

loads from the fuel storage racks are transmitted to the pool wall at either

the base level or at the base and the upper seismic bracing level.

Failure modes for the pool floor are:

1. Punching shear stress due to concentrated loads at the foot of the

storage racks.

2. Foundation settlements for soil; soil settlement may only be an issue

for piping relative displacements.

Failure modes for the walls are similar to that described for Mark I and Mark

II BWR's.

2.3.1.1.3 Failure Modes of Liner Plate

The fuel storage pool floor and walls are lined with 1/8" to 1/44" thick

stainless steel leak tight liner plate. The liner plates are welded to each

other by seam welds. Under the seam welds, leak detection (control) channels

are provided. Possible failure modes for the liner plates are:

1. Tearing of the liner plate or seam welds at the leak channels due to

2-5
vertical/horizontal, loads from fuel storage racks; this is of concern

only if the rack slides and the foot bears on leak channels.

2. Tearing of the liner plate due to sliding of the rack over any local

floor depression or wrinkles in the liner plate.

3. For a laterally braced rack, puncturing of the liner plate at the

knuckle in the vicinity of pool floor/walls intersection.

2.3.1.2 Fuel Racks

2.3.1.2.1 Types of Racks:

1. Low Density Racks (Cell Pitches 20" to 30") Subcritical

configuration is achieved by the physical separation of fuel

assemblies in an open-frame aluminum or steel structure. Structurally

the racks can be laterally braced at the upper and lower levels or

they could be bolted to the floor at four corners with upper and

lower grids connected by cross bracing.

2. Medium Density Racks (Cell Pitches 9" (BWR) to 13" (PWR))

Subcritical configuration is achieved by the flux trap principle.

The fuel assemblies are surrounded by stainless steel cans or cells

which prevent the neutrons in the water region between the cells from

returning to the fuel assemblies. The typical cell wall thickness is

1/8". Structurally the racks can be laterally braced at the upper

and lower grid levels, bolted to the pool floor at four corners with

upper and lower grids connected by cross-bracing, or they can be

cantilever cells (2 x 2 or 4 x 4 modules to reduce flexibility)

welded to a base structure.

3. High Density Fuel Racks (Cell Pitches 6" (BWR) 9" (PWR))

Subcriticality is achieved by addition of neutron absorbing poison

material between fuel assemblies. Poison is in the form of boron

containing material such as boron-carbide, borated stainless steel,

or borated aluminum. Storage cell walls have poison containing

pockets. Structurally the racks are mostly free standing or laterally

braced at the lower level. The honeycomb construction provides

structural integrity. The cell walls are typically 0.09" in

thickness. The cells are attached to each other by fusion or spot

welds.

4. Consolidated Fuel Racks - The fuel assembly is disassembled and

stored in a fuel canister. The canister is then stored in high

density storage racks. The consolidated ratio can be 2 to 1.

2.3.1.2.2 Failure Modes of Free Standing High Density Fuel Storage Racks

There are various types of fuel rack designs depending on the time (vintage),

fabricator and pool configuration. The design analysis of the fuel storage

2-6
racks is a complex task. There are non-linearities associated with gaps

between the fuel and cell wall, and friction at the base of rack., With a low

coefficient of friction, the racks could slide on the pool floor. With a high

coefficient of friction, the rack could lift up and fall back during a seismic

event. Due to the presence of water there is hydrodynamic coupling between

the fuel and racks, between adjacent fuel racks and between fuel racks and

pool walls [Habedank, et al., 1979; Dong, 1978].

The design analysis of the fuel storage racks is based on the requirements of

Appendix D to Standard Review Plan 3.8.4. The load combination and acceptance

criteria of SRP 3.8.4 Appendix D are based on subsection NF of ASME Code

Section III. Most of the high density fuel storage racks are of cold-form

construction. The design analysis procedure and acceptance criteria of coldform construction should be based on the AISI Cold Form Stainless Steel code

which is much more conservative (lower allowable stress limits and severe weld

spacing requirements) than the subsection NF of ASME Code Section III. Such

differences in the code criteria are taken into account in the fragility

analysis.

Due to multiplicity of storage rack designs, and fabricator/vendor(s), the

uncertainties in the design/analysis procedures and fabrication methods are

fairly large.

2.3.1.2.2.1 Free Standing/sliding Rack

A typical high density fuel storage rack consists of a welded assembly of

individual storage cells in a staggered checkerboard array. The storage cells

are comprised of double wall Type 304L stainless steel boxes welded to each

other with corner angles to maintain the desired pitch (6" for BWR & 8" for

PWR). Poison sheets in the form of boron containing materials such as borated

stainless steel, or boron carbide are placed between the pockets formed by

cell wall (0.08" to 0.1" thick) and poison retainer sheet (0.05") (See Figures

2-5 and 2-6). The unitized (honeycombed) construction of the racks provides

the structural strength to withstand vertical dead and live loads as well as

seismic loads.

The bottom of the storage cells sits on and is welded to the rack base plate

or base support structure. The base plate provides level support for each

fuel assembly and contains a cooling flow orifice for each cell. Each rack

module typically consists of a 15 x 18 array of fuel storage locations for a

BWR rack or a 10 x 10 array of storage locations for a PWR rack (Figures 2-7

and 2-8). These rack modules are supported by adjustable swivel feet. Each

support foot contains a remotely adjustable jack screw to permit the rack to

be leveled following wet installation. The support pad (feet) raise the rack

above the pool floor to the height required to provide an adequately sized

cooling water supply plenum.

The storage racks are positioned on the pool floor so that adequate clearances

are provided between racks and between pool structures to avoid impacting of

the racks during a design basis seismic event. The vertical dead weight and

2-7
seismic loads are transmitted directly to the pool floor by the' support

feet. The only horizontal seismic loads tra nsmitted from the rack structure

to the pool floor are those associated with friction between the rack

structure and the pool liner. The high density, free standing/sliding fuel

storage racks are popular.-with the utilities because they offer substantial

increase in storage capacity and they can be easily installed in an existing

contaminated wet pool.

2.3.1.2.2.2 Failure Modes of Free Standing/Sliding Racks

There are various types of design for free standing/sliding racks depending

upon the vendor (fabricator) and the time they were fabricated; The

critically stressed regions are the foot, storage cell wall 'in the vicinity 'of

the foot and the spot welds (fusion weld') with which the honeycomb structure

was built. Failure mod'es will vary depending upon the design feature of the

racks. The most likely failure mode however, will be the fallback impact load

in a seismic event. The storage cell wall above the foot weldment will buckle

and the fusion spot welds will open up resulting in the failure of the rack.

2.3.1.2.2.3 Laterally Braced High Density Fuel Racks

In the laterally braced storage racks, the storage racks are placed next to

each other without any gap between them. S'eismic bracing is provided around,

the periphery of the storage racks arrangement to transmit horizontal seismic

load to the storage pool walls. The seismic lateral bracing are provided at

the lower base level and sometimes at the lower base level and upper grid

level of the st'orage rack (See Figures 2-9 and 2-10). The vertical dead

weight and vertical seismic load are transmitted by the individual racks to'

the pool floor through the support foot. The horizontal seismic load is

transmitted from the storage racks to the peripheral seismic bracing and then

to the fuel pool walls. In order to account for variable distance between the

peripheral racks and the fuel pool walls, seismic bracing with 'rack screws at

their end are provided for easy field adjustment. The seismic bracings around

the periphery are very heavily loaded. Therefore, the most likely failure

mode for the laterally braced racks is the compression buckling failure of the

periphery seismic bracing. Once the bracing fails, the storage rack will

become the free standing/sliding type. As the load paths and design of a free

standing rack configuration is quite different from that of a laterally braced

desI~gn, the laterally braced racks without seismic bracing could fail. 'The

failure mode for the storage rack will be same as that of freestanding fuel

racks. Due to fallback impact load in a seismic event, the foot weldffent may

break or the storage cell wall above the foot weldment will buckle and the

spot weld (fusion weld) between storage cells will'open up, resulting in

failure of the storage racks.

2.3.1 .2.3 Consequence of Rack Failure.

Various failure modes for different rack systems were described above. The

exact nat'ure of the damage to the pool system by seismically induced failure

of racks has to be assessed by the nuclear systems engineers. The rack

2-8
failure under examination is not expected to lead to a sudden drainage of the

pool. The loadings introduced by different rack systems during a seismic

event will be considered in the structural failure mode evaluation of the

pool.

2.3.1.3 Pool Support Systems

Equipment failures in the systems that provide the spent fuel p901 cooling and

make-up water functions could also result in the loss of pool water inventory

through leakage and water boil-off. The components in these systems such as

the piping, valves, pumps, heat exchangers, tanks, and electrical distribution

systems are reviewed and the impact of their seismic failures on the spent

fuel pool behavior is examined. Seismic fragilities of these components are

developed using the failure modes and procedures described in the. PRA

Procedures Guide [USNRC, 1983] and [Ravindra et al., 1987]. The HCLPF

capacities of components and systems are derived using the seismic fragilities

of components and the Boolean expressions for the system failures. This

analysis is very specific to the particular spent fuel pool configuration.

For the two representative pools analyzed, the support systems are reviewed

and the seismic fragilities and HCLPF capacities estimated.

To determine the component included in the systems analysis, a screening

approach was utilized based on seismic capacity. This approach was developed

under the NRC Seismic Design Margins Program [Budnitz, et al., 1985, and

Prassinos et al., 1985] and uses an extensive table of generic seismic HCLPF

capacities for various components. Based on results of the Seismic Design

Margins Program HCLPF capacity of 0.3g was used as the screening value for

this analysis. It is judged that the components with HCLPF capacities larger

than 0.3g pga have no significant contribution to the failure of spent fuel

pools.

The systems that provide and support the cooling and make-up water functions

for the spent fuel pools examined were reviewed and inspected by performing a

plant walkdown. Those components within these, system that were not found to

have any characteristics that would reduce their seismic capacity, and were

adequately supported and anchored such that their HCLPF capacity was larger

than 0.3g were screened out from further consideration. Components with HCLPF

capacity less than the screening value remained in the analysis.

2.4 Analysis of Systems and Accident Sequences

The analysis of systems behavior is performed by combining plant logic models

with component fragilities and seismic hazard estimates. Event and fault trees

are constructed to identify the accident sequences and systems failures that

may lead to loss of water inventory in the pool and possible spent fuel

degradation.

Plant systems and sequence analyses used in seismic PRAs are based on the PRA

Procedures Guide [USNRC, 1983]. The analysis used in this study follows the

screening approach given in [Budnitz, et al., 1985] and is summarized as

follows:

2-9
1. Event trees were constructed reflecting the sequence of systems

successes and failures that will lead to loss of spent fuel pool

water inventory.

2. Fault trees were constructed reflecting the combination of component

failures that could lead to the systems failure identified. in the

event tree sequences. Only those component that were not screened out

of the analysis based on capacity were used in the fault trees

construction. Care was taken to include, however, all the failure

paths that could lead to system failure.

3. The Boolean expressions for the systems failures derived from the

fault trees were then substituted into the accident sequences

developed from the event trees to obtain a Boolean expression for

each accident sequence. The logical summation of all the accident

sequences resulted in an overall Boolean expression that described

the loss of water inventory from the spent fuel pool. The first,

second and third order cut sets of this Boolean expression were

obtained for further analysis.

4. The fragility of each component identified in the cut sets was then

estimated and those cut sets that had a component whose HCLPF

capacity was greater than the screening value (0.3g) were then

eliminated from further consideration.

5. The overall fragility for loss of spent fuel pool water inventory due

to systems failure was then determined by aggregating the fragilities

of the individual components as given in the overall Boolean

expression.

As an example, the Boolean expression for seismic core damage at a

representative plant could be:

MS = 4 + 8 + 10 + 14 + 17 + 21 + (12 + 22 + 26) * 9

The numbers represent components for which seismic fragilities have been

developed, as described earlier. The symbols "+" and "*" indicate "OR" and

"AND" operations, respectively. This Boolean expression represent the

combination of the Boolean expression for all accident sequences that lead to

seismic core damage. A plant level fragility curve can then be obtained by

aggregating the fragilities of the individual components using either Monte

Carlo simulation or numerical integration. The plant level fragility

represents the conditional probability of severe core damage for a given peak

ground acceleration at the site. The uncertainty in plant level fragility is

displayed by developing a family of fragility curves and assigning weights

(probability) to each curve derived from the fragility curves of components

appearing in the specific plant accident sequence.

2-10
2.5 Evaluation of Accident Frequency

The frequency of seismic induced accidents is evaluated by convolving the

seismic hazard distribution with the overall seismic fragility. If the seismic

hazard and the fragility are known, there will be a single hazard curve and a

single fragility curve,. and the result of the convolution will be a single

value for the annual frequency of the accident.

However, there is uncertainty in -the seismic hazard, this uncertainty is

represented by either a probability distribution (annual frequency'of

exceedance) or a discrete set of values of the frequency of exceedance along

with attached subjective weights'at each specific peak ground or spectral

acceleration.. Therefore, the uncertainty in the seismic hazard may be

represented by a family of curves with subjective weights attached to them or

by a family of curves at. different confidence values (subjective probabilities

of non-exceedance). Similarly,' the uncertainty in the seismic fragility can be

expressed by means of a family of fragility curves with subjective weights

assigned to them.

For the seismic hazard, the subjective weights are assigned by a hazard expert

for each hypothesis or seismic model that yielded the hazard curve.' In the

seismic fragility evaluation, the subjective weight for each fragility curve

is obtained by discretizing the uncertainty distribution which is also

assigned by a fragility analyst.

The convolution is performed by taking one hazard curve and one fragility

curve at a time, and the resulting frequency of the accident carries a

subjective weight equal to the product of the subjective weights assigned to

the specific hazard and fragility curves involved. The result of the

convolution of both families of curves (hazard and fragility) is a subjective

probability distribution for the. annual frequency of the seismic induced

accident. The convolution process can be accomplished by numerical integration

using the Discrete Probability Distribution method (DPD) [Kaplan, 1981] or by

simulation. The numerical integration procedure is described below.

Each discrete hazard curve may actually represent the output of one seismic

hazard model (by one hazard expert with subjective weight attached to the

curve); hence a large number of hazard curves may have to be convolved with

the fragility curve. Experience has shown that the use of five fragility

curves would be sufficiently accurate in representing the uncertainty in the'

fragilities of components [Ravindra, et al., 1984].

The uncertainty in seismic hazard is represented by a family of hazard curves

at different confidence values and an uncertainty distribution was used to

develop a new set of discrete curves. Five hazard curves (probability of'

exceedance vs peak ground acceleration) with subjective cumulative

probabilities of 5%, 15%, 50%, 85% and 95% were provided by [Bernreuter, et

al., 1987, Table 2-1]. A lognormal distribution was assumed for the

2-11
distrifbution of the uncertainty in probability of exceedance at each

acceleration value; the parameters of the lognormal distribution (i.e., median

and logarithmic standard deviation) were calculated by using the 50th and 95th

percentile values. If X is the hazard value at any acceleration, the median

of this distribution is X and the-logarithmic standard deviation of the

uncertai~nty is: 0= (s/1 .N) ln (X 5 /X5 0 ). However, since the tail of the

lognormal distribution extends to infinity, one might get values of variates

(probabilities of exceedance) greater than 1.0. A possible way to avoid this

problem is to modify the I ognormal distributi'on by truncating at the variate•

value ,of 1.0. Another option is to truncate the distribution at a selected

value of the variate (other than 1.0). In the present study, after examining

*i •the truncations at 99th, 98th and 97th'percentile values, it was judged that

. -the dfstribution could be truncated at X9 9, the 99th percentile value (i.e.,

99% of'the area of the uncertainty distri.ution is covered by the lognormal

distribution) where-X is given by X5 0 exp (2.33 0). The lognormal

distribution was norma ized to get a new distribution with cutoff at X The

sensi'tivity of this assumption was tested by running the case studies

1 (a),. (b), C() and (d) for the pool structure failure and the results are

shown,-in Table 6.2. The range of' the hazard represented'by this truncated

(lognormal) distribution at any acceleration was discretized into eleven

discrete values of the hazard (i.e., annual probability of exceedance) with

subjective probabilities of 0.03, 0.05, 0.07, 0.12, 0.15, 0.16, 0.15, 0.12,

0.07, 0.05 and 0.03. By repeating this procedure at other accelerations, a

family of hazard curves is obtained.

2"l 2
Table 2-1

Annual Probability of Exceedance at

Different Acceleration Levels

Peak Ground Cumulative Subjective Probability

Plant Acceleration

(9) 5% 15% 50% 85, 95%

Vermont

Yankee

Robinson

Unit 2

0.050

0.075

0.125

0.200

0.250

0.400

0.560

0.610

0.760

1.000

0.050

0.075

0.125

0.200

0.250

0.400

0.560

0.610

0.760

1.000

.1127E-03

.3970E-0O

.8001E-05

.151OE-05

.6306E-06

.8267E-07 .1291E-07

.7213E-08

.1144E-08

.1790E-10

.8451E-04

.3747E-04

.1063E-04

.2418E-05

.9409E-06

.9607E-07

.1067E-07

.4826E-08

.6465E-09

.2234E-10

.3192E-03

.1478E-03

.4473E-04

.1298E-04

.5792E-05

.8172E-06

.1455E-06

.8548E-07

.1992E-07

.2601E-08

.2548E-03

.1072E-03

.3355E-04

.8655E-05

.4223E-05

.6552E-06

.1250E-06

.7736E-07

.1967E-07

.3157E-08

.2594E-02

.1111E-02

.3544E-03

.1132E-03

.6117E-04

.1439E-04

.4383E-05

.3204E-05

.1319E-05

.3914E-06

.2196E-02

.8943E-03

.2871E-03

.8762E-04

.4682E-04

.1120E-04

.3330E-05

.2395E-05

.1013E-05

.2982E -06

.2902E-01

.1416E-01

*521SE-02

.1742E-02

.9989E-03

.2703E-03

. IOOSE-03

.7759E-04

.367GE-04

* 1448E-04

.1647E-01

* 7708E-02

.2746E-02

.9079E -03

.5118E-03

. 1465E-03

.5414E-04

.410BE-04

.1937E-04

.7717E-05

* 8137E-01

. 4217E-01

. 1572E-01

* 5554E-02

.3308E -02 .1032E-02

.41 53E-03 .3190E-03

. 1680E-03

* 7848E-04

.6474E -01

.3211E-01

.1247E-01

.4640E-02

* 2726E-02

.7376E-03

.2739E -03

.2078E-03

.1045E-03

.4173E-04

2-13
-1

10

-2 PGA DESIGN

10

-A

10

w

S10

S-7

10

4- ACCELERATION

Figure 2-1: Sample Seismic Hazard Curves

/0 • -

:IA/SECO-2

2-14
1.00

a 0.80 95% Confidence ' Median Fragility Curve Curve

LL

?D 0.60

.0

.0 0.50 ------------------ ---------------- 0.90g 0

PQ 0.40 -J C

0C 

0 

-Q- 5% Confidence 0.20 -Curve

0.05

0.00

0.00 0.17 0.30 0.40 0.80 1.20 1.60 2.00

Peak Ground Acceleration (g)

Figure 2-2: Example Fragility Curves for a Structure
--r Yield Lines

401

80'

Figure 2-3: .Failure Mechanism for the Pool Floor

2-16
.. -- Yield Lines

40f

Figure .2-4:

~*1

80'

Failure Mechanism for the Pool Wall

2-17
- .. II 

Lml . .. -,i I - I ,-i I

m 

ii II iIIalI

II lI II II l

,I tI II Ii I

II itItI , II ,

II i II i iI I .

li II l lII !

III Iii ii II '

II li IIIIII

I1 II liII " I

IIili i iiI 

00

II

II

II

II

II

II

'I

II

ii

II

II

II

tl

It

II

It

li

II

It

It

'I - II II

II II

II II II

Ii II II

II II II

II II II

II II II

It II II

Ii II II

II II II

II II~-

__.1 1I--9 3/4" Pitch Typ

Typical Section Thru

2 X 2 Module

Top View

--- 2 X 2 Module

I - ii U LI--- II 14 ~ I' .1 Leveling Pad (9)

I

Base Structure

Rack Base Plane

Side View

Pwr 10 X 10

Poisoned Fuel

Storage Rack

Figure 2-5: Spent Fuel Storage Rack Schematic (10 x 10)
2 3/4"

Typical Between Racks

At Top Of Cells

168 1/2"

4 1/2"

211

Typical Between Racks

Figure 2-6: Typical Rack Elevation View

2-19
168 112"

1/2"1

2 y1

TyPical 

Between 

Racks

Figure 2-6: T.Pical Rack Elevation 

View

2-19
2,

2,

Figure 2-7: Pool Layout

2-20
4.0" Minimum

rack to rack

and rack to poi

wall spacing Cask Area 8 X 10 10 X 10 10 X 10 10 X 10 for the East/

West direction / (typ) r•i

P•

8 X 10 8 X 10 10 X 10 10 X 10 10 X 10

LX1

3.0" Minimum rack to rack spacing

for the North/South direction.

(typ)

f1

1.25" Minimum rack to pool wall

spacing for the North/South

direction.(typ)

4 £ I

42' 0"

T

r

Figure 2-8: Plan Fuel Storage Rack Arrangement
LK-d I] LVL]1 14Uli LI

w

9 X 13 NJ

!)

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j I III II. I i L

Key: Crosshatched = Consolidated Region (1436 Avail. Cells) Bordered = Intact Cells (982 Avail. Cells)

Figure 2-9: Fuel Pool Arrangement for Laterally Braced High Density Fuel Racks
I r I I

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Figure 2-10: Typical Elevation - All Groups

2-23

3. SEISMIC FRAGILITY OF POOL STRUCTURES

3.1 Introduction

The evaluation described below was performed to determine the seismic capacity

of spent fuel pool structures. Failure of the spent fuel pool structure for

this study is defined to be a gross, rapid loss of the fluid which provides

cooling for the spent fuel rods. Minor leakage is not considered to

constitute failure. The investigation was limited to the structural

components forming the spent fuel pool.

3.2 Vermont (VYNPS) Spent Fuel Pool (SFP)

3.2.1 VYNPS SFP Structure Description

The Vermont Yankee spent fuel pool structure is an integral part of the

reactor building, and was designed as a Seismic Category I structure,

SSE=O.14g. It is situated at the south side of the reactor building, with the

top of the spent fuel pool located at the operating floor which is about 140

feet above the building foundation. The spent fuel pool is twenty six feet

wide in the north-south direction and forty feet wide in the east-west

direction. It is normally filled with water to a depth of 35.75 feet above

the bottom of the pool which is located at Elevation 306'-5"1.

The spent fuel pool structure is shown in Figures 3-1 and 3-2. It is built of

reinforced concrete. The floor slab is 4'-1" thick with an eleven inch thick

grout topping. The walls at the south, east, and west sides are from 4'-6" to

6,-0" thick. The drywell shield wall bounds the spent fuel pool on the north

side. The spent fuel pool structure is constructed integral with the reactor

building floors at Elevations 303'-0", 318'-8", and 345'2". The interior of

the spent fuel pool is lined with stainless steel plate anchored to the walls

and floor slab.

3.2.2 VYNPS SFP Structure Evaluation

A detailed evaluation of the Vermont Yankee spent fuel pool structure was

based on the maximum loading condition when the high density storage racks are

filled to their total storage capacity. The following plant specific

information formed the basis for this evaluation:

o Structural drawings

o Final Safety Analysis Report (FSAR)

o "Vermont Yankee Spent Fuel Storage Rack Replacement Report" (1986).

A number of potential failure modes of the spent fuel pool structure were

analyzed, including the following:

o Out of plane shear failure of the pool floor slab

o Out of plane bending failure of the pool floor slab

3-1
o Punching shear failure of the pool floor slab under the fuel rack

support pad

o Out of plane bending failure of the south pool wall

o Bending and shear failure of the girder under the south pool wall

o Bending and shear failure of the girder under the east pool wall

o Overall transfer of N-S and E-W inertial loads to the reactor building

Of these, the controlling failure mode with the lowest seismic capacity was

determined to be out of plane shear failure of the pool floor slab.

The spent fuel pool floor slab is 4'-1" thick with clear spans of 24'6" and

40'-0" in the north-south and east-west directions, respectively. It is

reinforced by steel bars in the two horizontal directions, but shear

reinforcement was not provided. The pool floor slab behaves as a two way,

member spanning between the girders and drywell shield wall at the

boundaries.

The spent fuel pool concrete was specified to have a minimum compressive

strength of 4000 psi. The results of non-destructive concrete strength

testing, believed to be Windsor probe,' from four locations at the bottom face

of the pool slab are reported in [Vermont Yankee 19861. These tests indicate

an average concrete compressive strength of 7000 psi. Because this test

method is reportedly very accurate, the median concrete compressive strength

was taken to be this value. A coefficient of variation of 0.15 was

estimated. This median strength and variability are within the range of

values used in previous fragility evaluations of nuclear plant structures with

4000 psi concrete.

Reinforcement for the spent fuel pool was specified to be Grade 40.

Reinforcement strength test data was not available. A median yield strength

of 48,000 psi and coefficient of variation of 0.10 were estimated based upon

past nuclear plant fragility evaluations.

The slab was evaluated for out-of-plane loading resulting from dead load plus

seismic load. Sources of dead load are the weights of the slab, grout, water,

fuel racks, and attached equipment. Sources of seismic load are:

o Vertical seismic response of the slab and attached masses

o Fluid impulsive and convective mode responses induced by horizontal,

seismic excitation

o Horizontal seismic response of the spent fuel racks

Thermal gradients across the walls and slab can introduce significant stresses

at elastic levels. However, these stresses are stiffness dependent and

diminish as the structure stiffness decreases. Under increased transverse

slab loads, the concrete will crack and crush and the reinforcement will

yield, with resulting reduction in stiffness. The slab stiffness approaches

zero asymptotically as the slab is loaded to failure. This evaluation is

performed to assess structure failure capacity. Because the slab stiffness is

small when its failure capacity is reached, thermal stresses at this load

level are small and can be neglected.

3-2
Maximum dead load shear at the slab supports due to uniform transverse

loadings from the weight of-slab, grout,. water, and attached equipment was

determined by modeling the slab as a one way member. Although the slab

actually develops two way action, this approximation is sufficiently accurate

since the difference is minor for the slab aspect ratio.

Seismic response of the slab was derived from the reactor building floor

spectra, shown in Figures 3-3 to.3-5, from [Vermont Yankee 19861. These floor

spectra were reportedly generated from a seismic analysis using the N 69' W

component of the 1952 Taft earthquake normalized to O.07g for the operating

basis earthquake (OBE). Five percent damping was assigned to the structure.

Safe shutdown earthquake (SSE).input, was taken to be .twice the OBE. An actual

plot of-the ground spectra corresponding to- the Taft ti-me history is not•

available. They were therefore assumed to be consistent with the spectra

shown in Figure B-2a of [Kennedy et al., 1985].

For the fragility evaluation, design seismic input to the spent fuel pool were

factored to reflect input that would result from the use of median centered

seismic response parameters. The factors used varydepending onthe response

considered as discussed below. Median site specific ground spectra are not

currently available for Vermont Yankee. Therefore, the shape of the median

Vermont Yankee ground spectrum was assumed to correspond to the median

spectrum recommended for rock sites by NUREG/CR-0098 [Newmark and Hall,

1978]. Ten percent structure damping was considered to be a median value

based on the recommendations of NUREG/CR-0098 for reinforced concrete

structures at or near the yield point. Conservatism in the design basis floor

spectra due to broadening and smoothing was also considered in the fragility

analysis by inclusion of a median factor of 1.1 for subsystem spectral shape.

The slab fundamental vertical frequency was determined using a closed form

solution. Accounting for potential cracking and continuity at the boundaries,

a vertical frequency of about 25 Hz was estimated. Ten percent damping is

considered to be a median value for the slab. Median slab shears due to

vertical response were determined by factoring the dead load shears. by the

estimated median spectral acceleration of 0.14g. This value is increased by a

factor of 1.15 above the SSE spectral acceleration to account for the

difference between the median site specific ground. spectrum at median

structure damping and the SSE ground spectrum at five percent structure design

damping for the structure fundamental vertical frequency of about 13 Hz. The

increase factor reflects potential unconservatism,.of the design ground

spectrum at the structure frequency relative to the median ground spectrum.

Seismic response of the pool water in the impulsive and convectivemodes,

causes transverse pressures to be exerted on the slab and walls under

horizontal excitation. Even accounting for wall flexibility, response of the

impulsive mode was determined to be essentially rigid. Fluid response was

thus determined following the provisions of TID-7024 [USAEC, 1963]. Resulting

seismic loads on the slab due to- fluid response were found to be small and

were therefore neglected. ..

3-3
The new, high. density spent fuel storage racks are freestanding and not

anch'r'ed to the structure. In this design, rack loads are transmitted to the

slab only by the rack support pads. Seismic response of the racks was

determined by dynamic, nonlinear analysis described in [Vermont Yankee,

1986]. This report contains maximum reactions imposed on the slab by

individual support pads as well as maximum total reactions for the rack

modules. Inspection of the individual support pad reaction time histories

indicates that they are nearly harmonic with long impulse durations relative

to the slab natural period. Based upon approximate methods for analysis of

structure response to impulse loads described in [Biggs 19.64], the rack loads

acting 'on the slab have a dynamic amplification factor of unity. Maximum slab

shears were therefore calculated using maximum total rack module reactions as

static loads, with shears from separate modules combined by the square root of

the sum of the squares (SRSS) since their phasing is random. Rack module

reactions are estimated to have a median factor of safety Of about 1.7,

accounting for potential conservatism in the input ground motion and the use,

of bounding base coefficients of friction.

The highest applied slab shear load occurs at the interface with the drywell

shield wall. Median slab shear strength at this location was determined from

available test data for reinforced concrete beams. From Figure 7.13 of [Park

and Pauly 1975], the' ultimate concrete shear stress was estimated to be

2.6 (f) 2, which represents an increase of about 25% above the ACI 349-80 Code

value, cnot including the strength reduction factor. This shear strength

implicitly accounts for moment-shear interaction. Reduction in shear strength

due to axial tension imposed by wall reactions is relatively small at seismic

levels corresponding to the lower tails of the fragility curves, which

dominate the seismic risk and HCLPF capacity. A median ultimate shear

strength of 120 kips per foot was calculated with the median concrete

compressive strength of. 7000 psi. This capacity was based upon- one way action

for the slab spanning in the short direction between the north and south

supports. [ASCE 1985] recommends a very small increase in shear capacity for

this slab accounting for two way action. However,- due tO the lack of shear

reinforcement,", shear failure will be nonductile. Because-the applied loading

is cyclic and the resulting failure mode is nonductile, any load '

redistribution will be unlikely and was therefore not included in this

evaluation.

Following the methodology described above, the slab was' fdund to have the

following overall fragility parameters:

Am =1.4g

aR= 0.26

B 0.39

The median slab fragility curve with 5% and 95% confidence bounds are shown in

Figure 3-6. Table 3.1 lists the median values and variabilities of the

individual factors. Because seismic loads used to' define the median strength

factor were already modified to account for median response parameters, the

median response factors are typically unity.

3-4
The HCLPF capacity for slab shear failure is calculated to be 0.5g. This

acceleration is about 3.5 times the SSE pga and represents a significant

margin of'safety. This failure mode is considered to be brittle and it is

probable that shear failure will lead to a.loss of structural integrity with

liner rupture and subsequent rapid loss of spent fuel pool water inventory.

Additional redundancy provided by alternative remaining load paths,. such as

catenary action of the slab, may be possible. However, these load paths were

not evaluated since the shear failure capacity is already well in excess of

the SSE.

The HCLPF capacity determined for the Vermont Yankee spent fuel pool structure

is consistent with NRC trial guidelines for seismic margin reviews. Following

Table 2.1 of [Prassinos et al., 1985], the spent fuel pool would not require

evaluation for a margin review earthquake with peak ground acceleration less

than 0.3g, since the HCLPF capacity of shear wall structures is considered to

be much greater -than'this input level. This is the case tor the Vermont

Yankee spent fuel pool structure which has a HCLPF capacity of 0.5g, even

including higher loads from the new high density storage racks. The HCLPF

capacity based upon the existing storage racks is estimated to be about 0.65g,

which is well above the 0.3g pga at which shear wall structures.are screened

out.

In a generic study on seismic risk from spent fuel pool accidents, [Sailor et

al., July 1987] assigned the greatest "weight" to the following structural

fragility for the Oyster Creek reactor building developed by [Kennedy et al.,.

1980].

Am = 0.75g

1R = 0.37

JeU = 0.38

The HCLPF capacity from these fragility parameters is.0.22g, which is much

less than the value for the Vermont Yankee spent fuel pool structure. There

are two major reasons for this difference. First, there has been considerable

evolution in techniques to quantify nuclear -plant seismic fragilities since

the Oyster Creek parameters were developed. These techniques generally result

in greater seismic capacities than.were previously calculated. Second, the

Vermont Yankee spent fuel pool structure fragility has been derived' from a

plant specific analysis that considers the actual seismic capacities and

loadings. This approach is preferable to the use of generic fragilities

obtained for, a structure whose configuration, resistance, and load paths may'

be markedly different.

3.2.3 Conclusions of the Structural Analysis

A detailed evaluation of the VYNPS spent fuel pool structure shows that there

is high confidence of a low probability, of failure due to an earthquake With

an average peak ground acceleration of 0.5g. The HCLPF capacity is.about.3.5

times the plant safe shutdown earthquake peak ground acceleration. Therefore,

the Vermont Yankee spent fuel pool-structure has significant safety margin

above its design basis when more realistic evaluation criteria are employed.

3-5
3.3 H. B. Robinson, Unit 2 (HBR2) Spent Fuel Pool

3.3.1 HBR2 SFP Structure Description

The Robinson spent fuel pool structure is a seismic Category I, SSE=O.20g,

reinforced concrete structure containing 35,167 cubic feet of water and is

housed within the Fuel Handling Building (FHB). The spent fuel pool is 43

feet wide in the north-south direction and 53 feet 6 inches wide in the eastwest direction measured from outside to outside face. It is normally filled

with water to a depth of 37 feet above the bottom of the pool which is located

at Elevation 236 ft 9 in.

The spent fuel pool structure is shown in Figures 3-7 and 3-8. It is built of

reinforced concrete. The floor slab is 4'-6" thick with reinforcement of #11

bars at 6 in. spacing in the negative moment regions, and #10 bars at 12 in.

spacing in the positive moment regions. A single steel column was installed

below the pool slab during the plant modification for the spent fuel storage

capacity expansion in 1982. The peripheral walls are all 6 ft thick with

reinforcement of #9 bars at 12 in. spacing each way and each face. The

interior of the spent fuel pool is lined with stainless steel plate anchored

to the walls and slab.

3.3.2 HBR2 SFP Structure Evaluation

A detailed evaluation of the Robinson Unit 2 spent fuel pool structure was

based on the maximum loading condition when the high density storage racks are

filled to their total storage capacity. The following plant specific

information formed the basis for this evaluation:

Structural drawings

Final Safety Analysis Report (Updated FSAR)

"Spent Fuel Pool Storage Extension New Column Under Fuel Pool Floor"

Calculation No. HB-102 prepared by Ebasco Services Incorporated for

Carolina Power & Lighting, March 1982.

A number of potential failure modes of the spent fuel pool structure were

analyzed, including the following:

Out of plane bending failure of the East or South wall.

Out of plane shear failure of the East or South wall

Out of plane shear failure of the pool floor slab

Out of plane bending failure of the pool floor slab

3-6
Punching shear failure of the pool floor slab at the added

steel column

Failure of the added steel column underneath the spent

fuel pool

Overall seismic stability of the spent fuel pool

Of these, the controlling failure mode with the lowest seismic capacity was

determined to be out of plane bending failure of the East wall.

The East or South wall resists the lateral forces of the old fuel racks. It

is modeled as a slab fixed on three sides and free at the top. The loads

considered in the evaluation of the seismic capacity of this wall are:

Hydrostatic loads (normal water level El. 273.67')

Hydrodynamic loads induced by horizontal and vertical

accelerations in earthquakes.

Wall inertia force

Reaction forces from the old fuel racks.

A realistic evaluation of the out-of-plane failure capacity of the wall was

done by using the yield line analysis under a uniformly distributed load. The

triangular hydrostatic pressure was converted into an equivalent uniform load.

Similarly, the pool slab was analyzed for static and dynamic loads. The

static loads included dead weight of the pool water, dead weight of the slab,

and dead weight of the fuel racks and cask. The dynamic loads included

hydrodynamic loads due to vertical acceleration, slab inertia, vertical

inertia of the fuel racks and cask, and reaction forces of the high density

fuel storage racks due to horizontal ground motion. Since the factors of

safety of the pool slab based on the elastic analysis were found to be much

higher than those of the east wall, no further analysis such as the yield line

analysis was performed for the pool slab.

The various factors needed in the estimation of the ground acceleration

capacity of the pool were evaluated as follows:

Strength Factor

The spent fuel pool concrete was specified to have a minimum compressive

strength of 3,000 psi at 28 days. Since no cylinder test results were

available, the median compressive strength (taking into account aging effects)

was estimated by using the previous PRA results as 4,900 psi and the

logarithmic standard deviation of 0.17 was also estimated.

3-7
Reinforcement for the spent fuel pool was specified to be Grade 40. Rebar

strength test data was not available. A median strength of 48,000 psi and a

logarithmic standard deviation of 0.10 were estimated based upon past nuclear

plant fragility evaluations.

The median strength factor was estimated to be 7.7 for the critical failure

mode of flexural failure of the pool wall based on the yield line analysis.

The uncertainty in this factor comes from the material strength uncertainty

(i.e., steel yield strength (0 U = 0.10)), uncertainty in the moment capacity

equation (estimated to be (8 = 0.15)) and other uncertainties in the yield

line analysis and torsional ýffect (aU = 0.15). Therefore, the total

uncertainty in the strength factor is estimated to be 0.23.

Inelastic Energy Absorption Factor

Based on the recommendations in Appendix C of ACI 349, a ductility factor of

10 was estimated for the east wall. The elastic natural frequency of this

wall, which is a subsystem of the spent fuel pool, was estimated to be about

50 Hz which is in the rigid range. Hence, the inelastic energy absorption

capacity factor was estimated to be equal to (10 x 0.13 = 1.3). The

logarithmic standard deviations in this factor were estimated to be

BR = 0.06 and 8U = 0.04.

Spectral Shape Factor

For the fragility evaluation, design seismic input to the spent fuel pool were

factored to reflect input that would result from the use of median centered

seismic response parameters. Median site specific ground spectra are not

currently available for Robinson. Therefore, the shape of the median Robinson

ground spectrum was assumed to correspond to the median spectrum recommended

by NUREG/CR-0098 [Newmark and Hall, 1978]. The median fundamental natural

frequency of the spent fuel pool including the soil and the piles that the

structure is founded on is estimated to be 3.8 Hz based on the information

provided in the Updated FSAR. Ten percent damping is estimated to be the

median value for the system, in consideration of the interaction between the

reinforced concrete structure and the foundation soil and piles. Since the

dynamic response of the spent fuel pool was estimated by using the NUREG/CR0098 10% median ground response spectrum, the median spectral shape factor is

1.0. The logarithmic standard deviations of this factor taking into account

the peak-to-peak variability and earthquake directional variability are

estimated as 8 R = 0.25 and U = 0.10.

Damping Factor

The response of the pool was calculated using 10 % damping which is considered

to be a median damping value. Therefore, the median damping factor is 1.0.

The system frequency is estimated to be 3.8 Hz, considering the foundation

media of the spent fuel pool. This frequency falls within the amplified

acceleration region of the NUREG/CR-0098 spectrum. The variabilities in

damping factor were estimated by taking 7% as a 16 percentile value. The

3-8
estimated valu'es Of ýR and u f6r damping factor are 0.03 and '0.14

respectively.

Modeling Factor

The median modeling factor-was assumed to be 1.0 with the value of 6U

estimated to be 0.15.

Soil Structure Interaction Factor

Robinson Fuel Handling Building is founded on piles. The median soilstructure interaction factor is estimated to be 1.0 with logarithmic standard

deviations 8R = 0.05 and ýU = 0.20. Note that .a higher value of ýU is

assigned for this factor compared to the Vermont Yankee evaluation because the

Robinson spent fuel pool is supported on piles&.

Subsystem Analysis Factors

The subsystem analysis of the pool is considered to be median centered; since

the load estimation on the wall is relatively straight forward compared to the

estimation of loads acting from the high density fuel racks on the pool slab

(as in Vermont Yankee), the uncertainty in the modeling of the subsystem. was

estimated to be lower and equal to 6U = 0.10. The median factor for

conservatism or unconservatism arising out of the modal combination method

used is estimated to be 1.0 with a nominal value of BR- = 0.05. With the

natural frequency~of the East wall falling in the rigid region, the'

uncertainty in the spectral shape and damping factors was estimated to be

zero.

Earthquake Component Combination Factor

The median response analysis was done by combining the seismic responses from

different directions by the 100%, 40% and 40% rule proposed by Newmark '

[Newmark'and Hall, 1978].' If we assume that the maximum vertical-earthquake

component-occurring simultaneously'with the horizontal component as the 3

logarithmic standard deviation case,ý the value of BR is estimated as 0.06.

Following the methodology described'above, the'overall fragility parameters

for the spent fuel pool wall were estimates as:

Am = 2.Og

8R = 0.28

BU = 0.40

The median slab fragility curve along with the 5% and 95% confidence bounds

are shown in Figure 3-9. Table 3.2 lists the median values and variabilities'

of the individual factors.' Because seismic loads used to define the median

strength factor were already modified to account for median response

parameters, the median'response factors are typically unity.

3-9
The HCLPF capacity for pool wall flexure failure is calculated to be 0.65g.

This acceleration is about 3.3 times the SSE pga and represents a significant

margin of safety.

3.3.3 Conclusions of the Structural Analysis

A detailed evaluation of the Robinson Unit 2 spent fuel pool structure shows

that there is high confidence of a low probability of failure due to an

earthquake with an average peak ground acceleration of 0.65g. The HCLPF

capacity is about 3.3 times the plant safe shutdown earthquake-peak ground

acceleration. Therefore, the Robinson Unit 2 spent fuel pool structure has

significant safety margin above its design basis when more realistic

evaluation criteria are employed.

3-10
Table 3.1

Median Factors of Safety and Variabilities

of Spent Fuel Pool Structure at Vermont Yankee

Factor Median OR OU

Strength 8.8 0. 0.30

Inelastic Energy Absorption 1.0 0. 0.

Spectral Shape 1.0 0.24 0.10

Damping (Structure) 1.0 0.02 0.08

Modeling (Structure) 1.0 0. 0.13

Soil-Structure Interaction 1.0 0.01 0.05

Spectral Shape (Subsystem) 1.1 0. 0.05

Damping (Subsystem) 1.0 0. 0.05

Modeling (Subsystem) 1.0 0. 0.15

Modal Combination 1.0 0.02 0.

Comb. of EQ Components 1.0 0.10 0.

TOTAL 9.7 0.26 0.39

Median pga capacity = 9.7 (0.14g) = 1.4g

HCLPF capacity - 1.4 exp [-1.64(0.26+0.39)] = 0.50g

3-11
Table 3.2

Median Factors of Safety and Variabilities of

Spent Fuel Pool Structure at H.B. Robinson 2

FACTOR Median 8 R U

Strength 7.7 0. 0.23

Inelastic Energy Absorption 1.3 0.06 0.04

Spectral Shape 1.0 0.25 0.10

Damping (Structure) 1.0 0.03 0.14

Modeling (Structure) 1.0 0. 0.15

Soil-Structure Interaction 1.0 0.05 0.20

Spectral Shape (Subsystem) 1.0 0. 0.

Damping (Subsystem) 1.0 0. 0.

Modeling (Subsystem) 1.0 0. 0.10

Modal Combination 1.0 0.05 0.

Comb. of EQ Components 1.0 0.06 0.

TOTAL 10.0 0.28 0.140

Median pga capacity = 10 (0.20g) = 2.Og

HCLPF Capacity = 2.0 exp (- 1.64 (0.28 + 0.40)) = 0.65 g

3-12
Nt

'St

'2LOk~~

Figure 3-1: Reactor Building Plan at Elevation 303'-0"
S/TA/F

POOL

(A

-~Z ~

*EZ ~

.•ETO 10V/-I/ Iý SeCT/O/N 2-2

Figure 3-2: Cross-Sections of the Spent Fuel Pool
z I

<

.. 00 __0.00

I I J ' 

FRE0UENCY ( C 5

Figure 3-3: Reactor Building OBE Floor Spectrum, Elevation 3031-0",

N-S Direction, 5% Damping
Z.

0

-J

uJ

L-)

1.00__ _ _ _ _______f

.10 1.0 0.0 0

F.R E 0 U E N C Y ( CPS)

Figure 3-4: Reactor-Building OBE Floor Spectrum, Elevation 303'-0", E-W Direction, 5% Damping
I *I I I I I II

I I I

--- 'I E A .I -- I - I- I I

I I' I m

IJ

m-1

z

0

b-i

U

-LJ

•C4.~-.&-.--.b-~---I-I-f--&--~l.-- I

I

r

I.-.. - - ~ -

___ -11144

10 - - - - - - - - - - - - - - - - - -

______ - ~- - - . - -

I

.0000

__________ 4I44'-4-4-4~44 III-I-t1~4~1

I + i 00 .I,0

R ' E 0 U E N CY ( CP s)

10.00

Figure 3-5: Reactor Building OBE Floor Spectrum, Elevation 303'-0"0 Vertical Direction, 5% Damping,
VERMONT YANKEE SPENT FUEL POOL

co

LUI

Dr_1

I.-H

ULL

0

CD

I--"

I-J

I-H

cn

0

ccClz

0

z

0

U

X 10

0.

X 10

PEAK GROUND ACCELERATION (g)

Figure 3-6: Spent Fuel Pool Structure Fragility
N

orh

!A

,t.o

A A (Figure)

Note: Crosshatched Area Indicates High Density Fuel Storage Racks

Figure 3-7: -HBR2 Spent Fuel Pool Slab-at EL 236'-9" (Top of Slab)
53' -6"

Section A-A

Figure 3-8: Cross Section of HBR2 Spent Fuel Pool

3-20
HBR2 SPENT FUEL POOL

liJ II

a:

I

LL

0

Ian

(n

0

0

_J

z

0

U-- 0

z

0

U

X to0

1.0"

-J

X 0 0

PEAK GROUND ACCELERATION (g)

Figure 3-9: HBR2 Spent Fuel Pool Structure Seismic Fragility Curves

4I. SPENT FUEL POOL SYSTEMS ANALYSIS

The spent fuel pool systems analyses was performed to gain an understanding of

the seismic response and capacity of pool cooling and makeup water systems at

selected nuclear power plants: Vermont Yankee Nuclear Power Station and H.B.

Robinson, Unit 2, Nuclear Power Plant. The analyses follows a seismic margins

approach utilizing the event tree, fault tree, and screening techniques

discussed in Chapter 2.

The ini~tiating event considered for this analysis is the occurrence of an

earthquake at the plant site. As a result, loss of offsite power is assumed

to occur due to the low seismic capacity of ceramic insulators in power plant

switchyards. Emergency electrical power sources were assumed available for

the analysis.

For the systems analysis, spent fuel pool failure is defined as loss of water

inventory leading to spent fuel rupture or degradation and possible

radioactive material release. Loss of pool inventory occurs due to water

boil-off following the failures of the pool cooling system and the systems

that provide water makeup. Loss of pool inventory can also occur due-to

leakage from the pool concurrent with failure of the makeup systems.

The scope of the systems analysis included only those front-line systems that

perform the primary functions of pool cooling and inventory makeup, and the

immediate systems or components that support these systems. The scope of this

effort did not allow a complete analysis of all the support systems within the

plant.

For the electrical support systems, only the cabinets, transformers, and motor

control centers that contain the primary power sources for the front-line and.

support system components were Included. The analysis of emergency electrical

power sources was not relevant to this effort and was not performed.

For secondary cooling of spent fuel heat exchangers, only those components

needed to perform secondary cooling were included. The complete analysis of

the closed loop component cooling systems that perform fuel pool secondary

cooling was not performed. In addition, the systems that provide cooling for

these closed-loop component cooling systems, like service water systems, were

not analyzed as part of this effort. An evaluation of these systems was

performed for VYNPS and found to be not significant.

The remainder of this chapter describes the spent fuel pool systems analyses

performed for this study. The analysis of the Vermont Yankee Nuclear Power

Station is presented .in Section 41.1 followed by the analysis of the H. B.

Robinson Unit 2 Nuclear Power Plant (Section 41.2).

4I-.1
4.1 Analysis of the VYNPS SFP Systems

The Vermont Yankee Nuclear Power Station is a General Electric (GE) boiling

water reactor (BWR) located outside of Vernon, Vermont. The Vermont Yankee

plant commenced operation in 1972 with a rated output of 514 MWe.

The spent fuel pool (SFP) at Vermont Yankee is elevated above grade and forms

part of the operating deck. The SFP concrete structure, metal liner and fuel

storage racks are seismic category Class I. A description of the structure is

given in Chapter 3. The plant is currently licensed to store a maximum of

2000 spent fuel assemblies using both low and high density fuel storage racks.

4.1.1 VYNPS SFP Systems Description

The fuel pool cooling and demineralizer system is shown in Figure 4-I [VYNPS

October 1985]. The cooling system transfers the spent fuel decay heat through

heat exchangers to the reactor building closed cooling water system (RBCCWS)

[VYNPS September 1985]. Water purity and clarity in the pool is maintained by

circulation through a filter-demineralizer system shown in Figure 4-2 (VYNPS,

FSAR).

The SFP cooling and demineralizer system consists of two circulating pumps

connected in parallel, two heat exchangers, two filter-demineralizers, and the

required piping, valves and instrumentation. Each pump has a rated capacity

equal to the system design flow rate and is capable of simultaneous-operation

with the other pump. Two filter-demineralizers are provided, each with a

design capacity equal to the design flow rate. The pumps circulate the pool

water in a closed loop, taking suction. from the pool, circulating the water

through the heat ,exchangers and filters, and discharging through diffusers at

the bottom of the pool. Water suction lines from the pool are equipped with

anti-siphon valves or a series o'f holes located below the water surface to

limit inadvertent pool drainage. Water discharge lines to the pool are

equipped with check valves.

The SFP cooling system is designed to maintain the pool water temperature

below 125 deg. F. during normal refueling operations. If the complete

reactor core is discharged to the pool, the water temperature may increase to

greater than the 125 deg. F. If this occurs, the RHR system using the train A

heat exchanger can be used to supplement the SFP cooling system to maintain

the pool temperature below 150 deg. F.

The RBCCW system provides cooling to the secondary side of the SFP heat

exchangers [VYNPS September 1985, VYNPS, FSAR]. Figures 4-3 and 4-4 show the

pertinent portions of the RBCCW system that provide this secondary cooling.

The RBCCWS is a closed loop cooling system consisting of two 100% capacity

pumps and two 100% capacity heat exchangers. A RBCCW surge tank is provided

to accommodate differential expansion of the water and is located on the

4-2
suction side of the pumps. The system consists of four parallel cooling loops

with the ability to isolate portions of the system to maintain essential

cooling water services. Service water is Used to cool the secondary side of,

the RBCCW heat exchangers.

Normal makeup to the pool is supplied from the condensate transfer system

shown in Figure 4-5. Water removal from the pool via the pumps is through

filter-demineralizer units to the condensate storage tank. Water makeup to'

the pool, however, can be provided by any of the following means (VYNPS

ON3157):

o normal pool fill line through valve DW-111 on the refueling floor, if

radiation levels permit (from the demineralized water/storage tank

(DWST), Figure 4-6).

o 8-inch condensate transfer fill line through valve FPC-29.

o charge any accessible fire hose and direct stream toward the pool.

o if the reactor building is unaccessible, through valves CST-14A, IFPCSP25A and FPC-FCV-16A or CST-14B, FCP-SP25B and FPC-FCV-16B (from the

condensate storage tank).

o rotate the spectacle flange at FPC-50 and initiate service water

makeup through RHR-183 and RHR-184, bypassing the pool demineralizers

(from service water system).

The SFP instrumentation is provided for both automatic and remote manual

operation to detect, record and control pump operation, pool temperature and

level, and system flow. A pool leak detection system is provided to monitor

pool leakage that consists of interconnected drainage paths behind the liner

welds that lead to an instrumented collection system.

4.1.2 VYNPS SFP Systems Analysis

The analysis methodology is described in Chapter 2. An event tree was used to

identify possible accident sequences that could lead to loss of water

inventory. Fault trees'were Used to analyze the failure of the systems givenzin the accident sequences. The results of the systems analysis is a Boolean

expression indicating the component failures that will lead to pool failure

and possible radioactive material release.

4.1.2.1 VYNPS SFP Event Tree Analysis

The initiating event is'the occurrence of an earthquake at the plant site.

The event tree is then drawn for this initiating event. The VYNPS spent fuel

pool event tree is shown in Figure 4-7. This event tree depicts the success

or failure of gross pool structural integrity, water tight integrity

(leakage), pool cooling and water makeup. These events are displayed across

the top of the figure and given an alphabetic designation, A through E. Pool

4-3
water makeup was divided into two events on the event tree, primary water

makeup and backup water makeup. The alphabetic designations are given below:

A = Gross pool structural integrity

B ='Pool water tight integrity (leakage)

C Primary pool water makeup

D = Pool cooling

E = Pool backup 'water makeup.

At each event in the tree, failure is represented by an upward branch and

success by a downward branch. Each path through the event tree represents an

accident sequence. Those sequences that end with a status of success indicate

successful operation of the pool systems and adequate water inventory in pool

following the initiating event. Those sequenfc'es with a status of failure

indicate possible loss of pool water inventory and spent fuel uncover as a

result of the earthquake. Total pool failure is the logical summation of the

failure accident sequence.

Each event in the failure accident sequences is analyzed to determine its

failure. Events that consist of systems failures are analyzed using fault

trees. The result of the systems analysis is minimum cut sets and a Boolean

expression for pool failure.

Inspection of the VYNPS event tree indicates that there are four accident

sequences that lead to possible pool failure. These failure accident

sequences are given below. Failure is indicated by the alpha designator

(i.e. A), 'success is indicated by 'not' preceding the alpha'designator

(notA).

SI = A

S2 = notA B C E

S3 = notA B notC D E

S4 = notA notB C D E

The first failure sequence is Si, gross structural failure of the pool leading

to complete loss of water inventory. This event is analyzed in Chapter 3.

The next accident sequence is S2. This sequence occurs following success of

the gross pool structure integrity but loss of pool water tight integrity and

complete failure to provide makeup water to the pool. Loss of water tight

integrity of the pool is defined as pool leakage resulting from partial pool

structural failure or the failure of the piping-connected to the pool

requiring both primary and backup water makeup to keep the spent fuel

4-4
covered. Notice in this accident sequence, pool cooling is not considered

since this function will not replace the water-leaking from the pool.

The next accident sequence, S3, occurs following success of gross pool

structural integrity and primary water makeup but failure in pool water tight

integrity. Pool failure occur due to failure of both pool cooling and backup

water makeup. If loss of fuel pool water tight integrity is defined as above,

pool cooling should not make a difference, since cooling will not supply water

to the pool.

This situation was considered in more detail by performing a conservative

thermal analysis of the pool to estimate the timing involved for water boiloff due to loss of pool cooling. Two failure states were considered that

correspond to the cooling requirements of the pool:

Pool State 1 (PSI): Maximum number of fuel subassemblies in the fuel pool as

presently licensed (2000), approximately 1/3 of a reactor core recently

discharged to the pool, reactor operational and at power, recently discharged.

fuel has undergone maximum fuel burn-up.

Pool State 2 (PS2): Maximum number of fuel subassemblies in the fuel pool as

presently licensed (2000), full core discharge to the pool 10 days after

reactor shutdown, fuel has undergone at least 30 days burn-up at full power,

reactor is down for inspection or maintenance.

Both analyses assumed all decay heat went into heating and boiling the pool

water. No heat transfer was considered to the pool walls or out of the pool

surface. The thermal analysis is given in Appendix A.

The results of the thermal analyses indicate that for PS1, it would take about

seven days for the pool water to heat up and boil-off such that the pool level

would decrease to the top of the spent fuel bundles following the loss of pool

cooling. For PS2, the time for the level to decrease to the top of the spent

fuel bundles was estimated to be about three days following the loss of pool

cooling.

In light of the long lead times, it was considered that for accident sequences

with coolant makeup from the primary makeup water source and successful

operation of the pool cooling system, there would be sufficient time to obtain

additional pool makeup water. Therefore, these accident sequence were

considered as successes and not analyzed further.

The last accident sequence, S4, occurs following the success of both gross

pool structural and water tight integrity. The failure sequence results from

complete loss of water makeup and pool cooling.

4.1.2.2 VYNPS Fault Tree Analysis

For the pool failure accident sequences, fault tress were developed for each

of the front-line system failures. Fault trees were also developed for the

4-5
support systems to the front-line systems. The individual fault trees were

then combined with respect to the accident sequence to analyze pool failure.

The individual fault trees were also analyzed separately.

The fault tree analysis was carried out using the FTAP computer code [R. R.

Wylie Computer Aided Fault Tree Analysis, Operation Research Center,

Universidy of California, August 19781 which gives the combination of single

(first order), double (second order), etc., component failures that lead to

the top event. The analysis of the fault trees was limited to the

consideration of cut sets up to fourth order (consisting of, at most, four

component failures). Cut sets greater than fourth order were not considered.

The overall pool failure cut sets were then inspected for components with

HCLPF capacity greater than 0.3g. If a cut set was found to contain a high

capacity component, it was eliminated from further consideration. The

remaining cuts sets were then combined into the Boolean expression for overall

pool failure.

Front-line system fault trees were developed for;

o loss of water tight integrity

o loss of primary water makeup

o loss of pool cooling from SFP cooling system

o loss of pool cooling from RHR system

o loss of backup makeup

Support system fault trees were developed for;

0 loss of secondary cooling to the SFP heat exchanger (failure of the

RBCCW system to support the secondary cooling)

o loss of the secondary cooling to RHR heat exchanger A (failure of the

RHR service water system to support secondary cooling)

The fault trees for these systems are given in Appendix B. The scope of the

fault trees include only those immediate components that support the operation

of the front-line and support systems. The components used in the

construction of the fault trees are listed on Table 4.1.

The fault tree for loss in pool water tight integrity considered failures

consisting of pipe breaks, failures of anchorages on heat exchangers, pumps,

and the demineralizer tanks leading to piping breaks and failures of check

valves and anti-siphon valves from stopping or limiting pool drainage. The

4-6
analysis of this fault tree indicated that there were nine second order cut

sets consistilng of either pipe breaks or equipment anchorage failures coupled

with failures of either check Valves or anti-siphon valves. There were

thirteen third order cut sets.

The fault tree for loss of primary water makeup consists of failures in makeup

from the demineralized water systems (DWS) and the condensate storage system

(CSS) along with a failure of the pool level instrument to indicate makeup is

needed. The DWS component failures consist of demineralized water storage

tank failure, the transfer pumps and human errors in not manually aligning the

DWS for SFP makeup. The CSS component failures consist of similar events.

Transfer pump failures consist of associated anchorage failure or motor

control center failure.

The analysis of the primary water makeup fault tree indicated that there is

one single order cut set and six second order cut sets along with several high

order cut sets. The single order cut set consisted of failure of the pool

level alarm system to indicate loss of water inventory. The second order cut

sets consist of combinations of opposite system (DWS and CSS) failures in

tanks, electrical power and human errors in aligning manual makeup water

valves.

The spent fuel pool cooling system fault tree considers failures in pool

suction from alternate paths, failures in the parallel pump/heat exchanger

arrangement and failures in the two parallel demineralizer-filters. The

demineralizer-filters also have a bypass line that was included in the

model. The component failures consist of a clogged strainer, inadequate

anchorage, human error, loss of electrical power and failure of secondary

cooling to the SFP heat exchangers.

The loss of secondary cooling to the SFP cooling system heat exchangers was

determined by developing a fault tree for the RBCCWS which provides this

function. This fault tree considers failures in the two parallel RBCCW pumps,

the two parallel heat exchangers, and failure of the RBCCW surge tank.. The

component failure consist of inadequate anchorage, loss of electrical power

and failure of secondary cooling to the RBCCW heat exchangers which is

provided by the service water system (SWS). The loss of secondary cooling to

the RBCCW heat exchangers (failure of the SWS) was not developed further.

The analysis of the SFP cooling fault tree after combining it with the RBCCWS

fault tree indicated that there is one single order cut set, sixteen double

order cut sets, twelve third order cut sets, and one fourth order cut. The

singleton failure was the RBCCW surge tank. The doubleton failures consisted

of combinations of opposite SFP pump train failures in pump and heat exchanger

anchorage and electrical power.

Spent fuel pool cooling provided by the RHR system was analyzed by developing

a RHR fault tree. This fault tree considers human error in properly aligning

the RHR system for pool cooling, failures of the two RHR pump trains with

associated motor operated valves (MOV) and failure of the RHR A heat exchanger

4-7
to provide pool cooling. The component failures consist of anchorage failure

and loss of electrical power to the pumps and MOVs.

Loss of secondary cooling to the RHR A heat exchanger was analyzed by

developing a fault tree of a portion of the dedicated RHR service water system

which provides this function. This fault tree considers failure to properly

align the RHR A secondary cooling MOV, which automatically closes on loss of

electrical power and requires manual opening, failure of the two pumps within

this system to provide cooling flow and failure of the normal service water

system which can provide this function. The normal SWS was not analyzed. The

component failures consist of anchorage failure and loss of electrical power.

The analysis of the RHR pool cooling fault tree after combination with the RHR

SWS tree indicated there are five single order cut sets, twenty second order

cut sets, eight third order cut sets and one fourth order cut set. The single

order cut sets consisted of human error in aligning the RHR system for pool

cooling, MCC 9B failure and RHR A heat exchanger anchorage failure.

The fault tree for the loss of backup water makeup to the SFP considers makeup

from the fire water system, the SWS and an alternate path from the CST

(through valves SP-25A(B) and CST-14A(B)). Due to the long recovery time,

failure of the fire water system and SWS to provide makeup to SFP was not

analyzed. Therefore, this fault tree only consisted of human errors in

properly aligning an alternate makeup water path from the CSS. The analysis

of this fault tree indicated that there is one first order cut set and ten

second order cut sets. The singleton is failure of the CST. The doubletons

consist of combinations of electrical power and anchorage failures from the

two trains of the CSS system.

4.1.2.3 VYNPS Accident Analysis

The development of a Boolean expression for the SFP involves combining the

individual fault trees as indicated by the failure accident sequences and

logically summing over all the sequences. This logical summation is

accomplished by developing a high level fault tree for pool failure consisting

of the four failure accident sequences given in Section 4.1.2.1.

Two overall high level pool fault trees were developed. One for each of the

Pool States discussed in Section 4.1.2.1 because of the different pool cooling

requirements. For PSI, pool cooling is accomplished by the SFP cooling

system. The high level fault tree for PSI is shown in Figure 4-7.

For PS2, the SFP cooling system is not adequate and the RHR system is required

to cool the pool. The high level fault tree for PS2 is shown in Figure 4-8.

The high level fault trees were analyzed to obtain the cut sets for pool

failure. The analysis considered cut sets only up to fourth order. The

results of the analysis indicated that for PSI case, there was 1 first order

cut set, 135 fourth order cut sets and over 20,000 higher order cut sets. For

the spent pool with fuel loading described by the assumption of PSI, pool

4-8
failure result from gross structural failure or many combinations of component

failures, one from each the four system that support the pool: primary water

makeup, pool cooling from the SFP cooling system, pool cooling from the RHR

system and backup water makeup.

For the PS2 case, there was 1 first order cut set, 35 third order cut sets,

365 fourth order cut sets and over 20,000 higher order cut sets. Spent fuel

pool failure with a fuel loading as assumed by PS2 results from gross

structural failure, or many combinations of component failures, one from each

of the three system that support the pool: primary water makeup, pool cooling

from the RHR system and backup water makeup. Pool cooling from the SFP

cooling system was not sufficient under the condition of PS2.

Each third and fourth order cut set was examined and those cut sets containing

high capacity component were eliminated. The remaining cut sets were combine

into the Boolean expressions for each pool state given below:

PS1= 12 + ( 7 + 11 + 1 * 2 )*(5 * {1 4 + 9 + 8 + 10)

+ 4 * { 3 + 6 })

PS2 =12 + (7 + 11 + 1 * 2 ) 4 + 9 + 10)

where the number represent the component seismic fragilities and random

failure rates given in Table 4.2. It may be noted that components 1, 2, 7,

and 11 are all in the non-safety electrical systems of the plant.

4.2 Analysis of the H. B. Robinson Unit 2 (HBR2) Spent Fuel Pool Systems

The H. B. Robinson Steam Electric Plant, Unit 2 is a Westinghouse pressurized

water reactor (PWR) located near Hartsville, South Carolina. The H. B.

Robinson plant commenced commercial operation in 1971 with a rated output of

665 MWe.

The spent fuel pool at HBR2 is elevated and forms a separate structure from

the containment and auxiliary building. The SFP structure is seismic category

Class I with a stainless steel liner of all welded construction. A

description of the SFP structure is given in Chapter 3.

Fuel subassemblies are transferred to and from the pool via a fuel transfer

tube that connects the SFP structure with the containment building. The pool

is licensed to store 544 spent fuel subassemblies using both low and high

density storage racks.

4.2.1 HBR2 SFP Systems Description

The Spent Fuel Pool Cooling and Purification System consists of the pool

structure and piping loops through which borated water is circulated to remove

4-9
decay heat dissipated from the spent fuel subassemblies and through

filter/demineralizers as a means to purify and maintain water clarity [HBR2

November 1986]. The SFP Cooling and Purification System is shown in

Figure 4-9.

The SFP cooling system consists of two 100 percent capacity pumps, one heat

exchanger, filter, demineralizer, associated valves and piping, and

instrumentation. The two SFP cooling pumps, one operating and the other as

backup, draw water from the pool to circulate through the heat exchanger and

return to the pool. The demineralizer is sized to pass 5 percent of the loop

flow and the purification loop can be used to circulate water from the

refueling water storage tank (RWST) in a loop back to the RWST using the

refueling water purification pump.

The SFP pumps are manually controlled locally or from the SFP operating

deck. Two emergency connection are provided across the SFP pumps. One at the

suction of pump "B" and the other downstream of the pumps to allow the use of

another."portable" pump for cooling under abnormal condition.

The pool has two suction lines to the pumps. The upper suction line

penetrates the pool wall four feet below the normal water level and.about nine

feet above the top of the spent fuel assemblies. The other suction line is..

located at the bottom of the pool and provides for complete pool drainage.

This line is equipped with a closed and locked valve inside the pool to

prevent drainage if an external severance of this line should occur. The

bottom line is also equipped with an external closed and locked valve.

The return piping to the pool, from the pumps, discharges at six feet above

the fuel assemblies and is equipped with a 1/2 inch hole, six inches below the

normal water level to serve as a vacuum breaker to prevent siphoning.

The SFP cooling system is designed to maintain the pool water temperature

below 120 deg. F. during normal refueling operations when 1/3 of a core is

discharged to the pool. During the discharge of a complete reactor core. into

the pool, the water temperature is maintained at or below 150 deg. F. (HBR2

FSAR). Should the pool water temperature exceed 150 deg. F., fuel assemblies

are transferred back to the reactor cavity until the temperature falls to 150

deg. F. or less.

During normal operations the SFP pool water level is maintained at 22 ft. 5

in. above the fuel assemblies. The pool water level is monitored continuously

with alarm set points at 2 inches above and 7 inches below the normal water

level.

The component cooling system provides cooling to the secondary side of the SFP

heat exchanger. The component cooling water (CCW) system is a closed loop

cooling system consisting of three pumps, two heat exchangers, a surge tank,.

and the associated piping, valves and instrumentation [HBR2 June 1987].

4-10
During normal full power operation, one CCW pump and one heat exchanger

accommodates the heat removal loads. All three pumps and two heat exchangers

are needed to remove sensible and residual heat during plant shutdown.

Component cooling water is supplied to the SFP heat exchanger to maintain pool

temperature within specified limits. Temperature is control by throttling a

manual valve on the outlet of the heat exchanger. Emergency cooling

connections are provided across the secondary side of the SFP heat exchanger

so an external cooling system (i.e. fire system) can be used if the CCW has to

be isolated.

The service water system (SWS) supplies secondary cooling to the CCW heat

exchangers. The SWS was not modelled as part of the systems analysis of the

HBR2 spent fuel pool.

Normal water makeup to the pool [HBR2 April 19851 is supplied from the

refueling water storage tank (RWST) via the refueling water purification (RWP)

pump. The reactor coolant drains pump can also deliver water to the RWP pump

from the reactor coolant drains tank. Makeup water can also be provided to

the pool from the condensate storage tank (CST) via the demineralized water

pump (DW) delivered directly to the SFP purification loop or to a serviceconnection (faucet) at the operating deck the fuel handling building.

The SFP instrumentation is provided for remote manual operation to detect,

record and control pump operation, pool temperature and level, and system

flow. A pool leak detection system is provided to monitor pool leakage that

consists of 10 one-inch diameter pipes embedded in the concrete along the

bottom of the pool where the walls join the floor behind the liner. These

pipes are interconnected and lead into a valved end. The leak detection

system is checked daily during shift change to monitor pool leakage.

4.2.2 HBR2 Systems Analysis

The HBR2 systems analysis follows the same general process that was used in

the VYNPS analysis. The systems analysis of the HBR2 spent fuel pool utilized

a seismic margins screening approach as described in Chapter 2. An event tree

was developed to identify possible accident sequences that could lead to loss

of pool water inventory. Fault trees were used to analyze the failure of the

systems given in the accident sequences.

4.2.2.1 HBR2 Event Tree Analysis

The HBR2 spent fuel pool event tree is shown in Figure 4-10. This event tree

was developed with consideration of an earthquake at the plant site as the

initiating event and depicts the success or failure of gross pool structural

integrity, water tight integrity (leakage), pool cooling and makeup water.

These events are displayed across the top of the event tree and given

alphanumeric designators as shown.

Inspection of the HBR2 event tree indicates that there are three accident

4-11
sequences that lead to possible complete loss of water inventory. These three

accident sequences are given below. Failure is indicated by the alphanumeric

designator (i.e. W), success is indicated by 'not' preceding the alphanumeric

designator (notW).

ASI = W

AS2 = notW X Z

AS3 = notW notX Y Z

The first accident sequence is ASI, gross failure of the spent fuel pool

structure which results in the complete loss of pool water inventory. This

event is analyzed in Chapter 3.

The next accident sequence is AS2. This accident sequence occurs following an

earthquake where the spent fuel pool has survived a gross failure but the

water tight integrity of the pool has not been maintained. Pool leakage occurs

through possible pipe breaks or severe leakage from the pool liner. Pool

leakage occurs over a relatively long time, however, pool make-up is required

to prevent fuel uncover, damage and possible radioactivity release. SPF

failure occurs from seismic failure of the water make-up systems.

The third accident sequences is AS3, which occurs following an earthquake

where pool water tight integrity has been maintained. Fuel uncover, damage

and possible radioactivity release occurs from failure to cool the water in

the pool resulting in water boil-off and failure to provide make-up water in

the event cooling fails. If pool cooling is successful or the pool makeup

systems are available, then SFP success is maintained.

A thermal analysis of the HBR2 SFP was reviewed in [USNRC 1982]. This review

indicated that with the maximum heat load in the pool (full core discharge),

the time it would take for water temperature to rise from its allowable

maximum temperature of 150 deg. F. to boiling would be 6.83 hours and that the

boil-off rate would be 41.23 gpm (5.51 cu.ft./min.). Using this boil-off rate

and the water heat time, it takes about 3.2 days for the pool inventory

(23352.7 cu. ft.) to decrease to the top of the fuel assemblies (22.41 ft.)

after a full core discharge to the pool following loss of pool cooling and

water make-up.

4.2.2.2 HBR2 Fault Tree Analysis

Fault trees were developed for those front-line and support systems failures

that were indicated in the accident sequences derived from the event trees.

Front-line system fault trees were developed for:

o loss of SFP cooling, and

o loss of water make-up to the pool.

4-12
A support system fault tree was developed for failure to supply secondary

cooling to the SFP heat exchanger. The HBR2 fault trees are given in Appendix

C. The component that were Used to develop these fault trees are given in

Table 4.3.

The fault tree analysis considered only those component and systems that,

support the operation of the front line systems. The service water system that

provides secondary cooling to the CCW system was not analyzed.

The loss of water tight integrity event was not modeled. Pipe breaks and tank

failures were considered in the other two fault trees for the SFP systems that

are connected to the pool water.

The SPF cooling fault tree considers failure of the.SFP heat exchanger to

provide cooling and failure of the SFP pumps to circulate the cooling flow.

The component failures consist of inadequate anchorages on the heat exchanger

and the pumps, loss of electrical power, failure of the pool temperature

instruments to indicate temperature increase, human error in properly aligning

the alternate pump and cooling loop pipe failure.

The loss of secondary cooling to the SFP heat exchanger was determined by

developing a fault tree of the component cooling system that provides this

function. The CCW fault tree considers failure of the CCW pump to provide

cooling flow, failure of the CCW surge tank, and failure of the CCW heat

exchangers. The component failures consist of inadequate anchorage of the CCW

pumps and heat exchangers, loss of electrical power, human error in properly

aligning the alternate pumps and heat exchangers and failure to utilize the

fire water system for alternate cooling of the SFP heat exchanger secondary.

The fire water system was assumed available and not analyzed.

The analysis of the SFP cooling fault tree indicated that there were four

first order cut sets, 2 second order cut sets and 26 third order cut sets. The

first order cut sets were pipe breaks, failure of the CCW surge tank and

failure of the SFP heat exchanger. The second order cut sets consisted of a

combination of the temperature instrument failure and pump failures. One pump

fails because of inadequate anchorage and the other due to failure in

electrical power. Failure in the temperature instrument results in no

indication of pool heat up and thus no alignment of an alternate cooling path

for the pool.

The loss in water make up fault tree considered failure of pool make up

through the purification loop utilizing the refueling water purification pump

and failure to provide make up flow from the faucet located on the SFP

operating floor. Component failures consist of inadequate anchorage on the RWP

pump and the deminetalizer. Water pump, failure of the condensate storage,

refueling water storage and demineralizer tanks, loss of electrical power,

failure in pool level and area radiation instruments and human error in

properly aligning the primary and alternate make up water paths.

4-13
The analysis of the pool water inventory make up fault tree indicates that

there are 21 second order cut sets. These cut sets consist of combinations of

failures in the systems that monitor and supply make up water to the pool.

Double component failures of the CST, RWST, demineralizer tank, electrical

-power cabinets, pump anchorage, level sensor and area radiation monitor, and

human error in properly aligning the make up supply.

4.2.2.3 HBR2 Accident Analysis

The development of the Boolean expression for the SFP involves combining the

individual fault trees as indicated by the accident sequences and logically

summing over all the sequences. For the HBR2 analysis, the summation was

accomplished by a logical union of the three accident sequence equations.

The analysis resulted in one first order cut set, 29 second order. cut sets, 84

third order cut sets and 92 fourth order cut sets., The first order cut set was

gross failure of the pool structure leading to complete loss of pool

inventory. Each cut set was examined, and those containing high capacity

component were eliminated. The remaining cut sets were combined into the

Boolean expression for pool failure given below:

HBR2 pool failure = 1 + ( 4 + 5 + 7 + 9) * (6 + 8 + 12)

where the numbers represent component seismic fragilities and random failure

rates given in Table 4.4.

4-14
Table 4.1

Components of VYNPSSpent Fuel Pool Systems

Description

MCC 7b fails (SS Transformer 7)

MCC 7C fails (SS Transformer 7)

MCC 8b fails (SST 8)

MCC 9b fails (SST 9)

SWGR 8 fails (SST 8)

SWGR 9 fails (SST 9)

4160 SWGR bus fails (SST 3)

4160 SWGR bus 4 fails (SST 4)

no SFP low level alarm

condensate storage tank fails

demineralized water storage tank fails'

ops fails to open manual valve DW-111

ops fails to open manual valve CST-14a

ops fails to open manual valve CST-14b

ops fails to open manual valve SP-25a

ops fails to open manual valve SP-25b

ops fails to open manual valve FPC-28

ops fails to open manual valve FPC-29

ops fails to open manual valve DW-35c

ops fails to open manual valve V-53

condensate transfer pump P-4-1a fails

condensate transfer pump P-4-1b fails

demin transfer pump P-63-la fails

demin transfer pump P-63-1b fails

rbccw pump P-59-1a fails

rbccw pump P-59-1b fails

rbccw heat exchanger E-8-1a fails

rbccw heat exchanger E-8-1b fails

rbccw surge tank TK-7-1a fails

RHR heat exchanger E-14-1a fails

ops fails to open MOV V-13a

ops fails to open MOV V-13c

ops fails to open MOV V-15a

ops fails to open MOV V-15c

ops fails to open MOV V-89a

ops fails to open manual valve V-25

ops fails to open manual valve V-50

4-15
Table 4.1

(Continued)

Components of VYNPS Spent Fuel Pool Systems

Description

RHR pump P-10-la fails

RHR pump P-10-1c fails

SWS RHR pump P-8 la fails

SWS RHR pump P-8-le fails

SFP demineralizer/filter DM-2-1a fails

SFP demineralizer/filter DM-2-1b fails

SFP heat exchanger E-9-1a fails

SFP heat exchanger E-9-1b fails

ops fails to open manual valve V-39

ops fails to open manual valve V-46

ops fails to open manual valve SF-77

SFP cooling pump P-9-1a fails

SFP cooling pump P-9-1b fails

fuel pool outlet strainer fails (clogged)

dmin/filter DM-2-1a strainer S-12-la fails

dmin/filter DM-2-1b strainer S-12-Ib fails

dmin/filter DM-2-1a fails causing pipe break

dmin/filter DM-2-1b fails causing pipe break

HTX E-9-1a fails causing pipe break

HTX E-9-1b fails causing-pipe break

pump P-9-1a fails causing pipe break

pump P-9-1b fails causing pipe break

pipe.break in segment fpc-19 below RHR connection

anti-siphon valve SF-11 fails

check valve V-18 fails

check valve V-21a fails

check valve V-21b fails

undeveloped event-failure of fire water system to provide make-up

station service transportation

4-16
Table 4.2

VYNPS SFP Component Failures

Number Component Failure

1 Seismic failure of MCC 7B

2 Seismic failure of MCC 7C

3 Seismic failure of MCC 8B

4 Seismic failure of MCC 9B

5 Seismic. failure of RBCCW Surge Tank (TK-7-1A)

6 Seismic failure of the SFP Heat Exchanger (E-9-1A)

7 Seismic failure of the Station Service Transformer (SST 7).

8 Operator fails to open motor operated valve (MOV V-89A)

9 Operator fails to open manual valve (V-25)

10 Operator fails to open manual valve (V-50)

11 Seismic failure of Switchgear (4160V SWGR Bus 4)

12 Gross seismic failure of the SFP Structure

4-17
Table 4.3

H. B. Robinson 2 Nuclear Power Plant

Component List

SPENT FUEL POOL COOLING (SFPC) SYSTEM

Valves

type

MV

MV

MV

MV

MV

MV

MV

CV

designation

SFPC-796

SFPC-836A

SFPC-793

SFPC-797

SFPC-837

SFPC-836

SFPC-819

SFPC-820

position*

NO

NO

LC

NC

NO

NO

NO

Pumps

SFPC Pump

SFPC Pump

A

B

- 480v Bus No.3

- 480v Bus No.1

SFPC Heat Exchanger

Instruments

Pool Level - LA-651

Pool Temperature - TIC-651

SFPC HTX Temp. - TI-653

PURIFICATION LOOP OF SFPC SYSTEM

Valves

type

MV

MV

MV

MV

MV

MV

MV

MV

MV

MV

designation

SFPC-798A

SFPC-800A

SFPC-799B

SFPC-799C

SFPC-801

SFPC-811

SFPC-812

SFPC-813

SFPC-800B

SFPC-798B

position

NO

NC

NC

NC

NO

NO

NO

NO

NC

NO

*See legend at end of table.

4-18
Tanks

SFP Pit Demineralizer

Filter

SFP Filter

Instruments

Pressure - PI-670

PI-655A

PI-656A

PI-656B

Flow - FI-654

FE-654

SFPC MAKE-UP WATER

From the Condensate Storage Tank (CST)

(SFPMU-CST)

Valves

type designation position

MV SFPC-808 LC

MV DW-255 NO

MV DW-252 NO

CV DW-293 --

CV DW-290 --

Tanks

Condensate Storage Tank (CST)

Pumps

Demineralized Water Pump - 120v/208v MCC No.8

From the faucet in the SFP Pit

(SFPMU-FA)

Valves

type designation position

MV DW-269 NC

MV DW-267 NO

MV DW-257 NO

Hose stored in SFP Pit area

4-19
Instruments

Flow - FI-6420

From the Liquid Disposal System

(SFPMU-LDS)

Valves

type

AOV

AOV

AOV

AOV

AOV

AOV

CV

CV

Tanks

Reactor

Pumps

Reactor

Reactor

designation

LCV-1003A

LCV-1003B

FCV-1018A

WD-1721

FCV-1018B

WD-1722

WD-1724

WD-1725

position

NC

NC

NO

NO

NO

NO

failure mode

fails closed

fails closed

fails closed

fails closed

fails closed

fails closed

Coolant Drain Tank (RCDT)

Coolant Drain Pump A

Coolant Drain Pump B

- 480v MCC No.2

- 480v MCC No.1

Instruments

Pressure - PI-101 8A

PI-1018B

Temperature- TE-1058

From the incoming header IVSW

(SFCMU-IVSW)

between FCV 1018A & 1018B

Valves

type

MV

MV

CV

designation

SFPC-804

SFPC-802A

SFPC-803

position

NC

NO

Pump

Refueling Water Purification (RWP) Pump - 48ov MCC No.1

4-20
From the RWST

(SFPMU-RWST)

Valves

type

MV

MV

Tanks

Refueling

designation

SFPC-805A

SFPC-805B

position

NO

NO

Water Storage Tank (RWST)

SECONDARY SIDE TO THE SFP HEAT EXCHANGER

(SFP-HTX-CCS)

Component Cooling System (CCS) Components

Valves

type

MV

MV

MV

MV

MV

MV

MV

MV

MV

MV

MV

MV

MV

MV

designation

CC-701A

CC-701B

CC-701C

CC-703A

CC-703B

CC-703C

CC-710

CC-713A

CC-713B

CC-712A

CC-712B

CC-772

CC-776

CC-775

position

NO

NO

NO

NO

NO

NO

LO

CO

CO

CO

CO

NO

CO

NO

Pumps

Component

Pump

Pump

Pump

Cooling

A - 480v

B - 480v

C - 480v

Dedicated Shutdown Bus

Bus No.El

Bus No.E2

4-21
Component Cooling Heat Exchangers

CC HTX A

CC HTX B

Tank

Component Cooling Water Surge Tank

Instruments

flow element - FE-622

Temperature - TI-621

Service Water Cooling

Cooling Water System

to Secondary Side of Component

Valves

type

MOV

MOV

MOV

MOV

MV

MV

MV

MV

MV

MV

MV

MV

CV

CV

CV

CV

designation

V6-1 2B

V6-1 2C

V6-12A

V6-1 2D

SW-20

SW-21

SW-739

SW-740

SW-5

Sw-6

SW-7

SW-8

SW-374

SW-375

SW-376

SW-377

position

NO

NO

NO

NO

NO

NO

NO

NO

LO

LO

LO

LO

failure mode

fails as is

fails as is

fails as is

fails as is

(alternate

- 480v MCC No.5

- 480v MCC No.6

- 480v MCC No.5

- 480v MCC No.6

- 480v MCC No.5)

Pumps

Service Water

SW Pump A

SW Pump B

SW Pump C

SW Pump D

Strainers

SW $6-IA

SW S6-1B

Pumps

- 480v Bus No.El

- 480v Bus No.El

- 480v Bus No.E2

- 480v Bus No.E2

(alternate - 480v Dedicated Shutdown Bus)

4-22
Instruments

Pressure - Px-1615A

Px-1615B

Temperature - Tx-1 682A

Tx-1 682B

Tx-1 683A

Tx-1 683B

TI-1655

TI-1656

LEGEND

MV - manual valve

MOV - motor operated valve

AOV - air operated valve

CV - check valve

NO - normally open

LO - locked open

NC - normally closed

LC - locked closed

CO - chained open

4-23
Table 4.4

HBR2 SFP Component Failures

Number Component Failure

1 Gross seismic failure of the SFP Structure

4 Seismic failure of the Demineralizer tank

5 Seismic failure of the RWST

6 Operator fails to align CST for makeup

7 Seismic failure of 480v MCC No.1

8 Seismic failure of 480v MCC No.8

9 Operator fails to align RWST for make up

12 Seismic failure of the CST

4-24
To IER From

FPC 37 6MP 19

41.

In Prom

CondensateTransfer Pump

FPC 53

IfFilter

Demineralizer I

I

J Radvaste Building

LrCCI

-Sample

Figure 4-1: Fuel Pool Cooling System
. r 1.0. . * 0

Li 04-A 7 fs ___ ___- I___ A

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VERMUOT YANKEE

NUCLEAR POWER STATION

FINAL SAFETY ANALYSIS REPORT

Fuel Pool Filter Deaneralizer System

Figure IO.S-2

Figure 4-2: Fuel Pool Filter Demineralizer System
IIS

!

r•

"-,,I

P59-IA-I PRIMARY CONTANMANT

pan

CORWHEUSS

MON SEMMMAL

LOADS

ALTERNATE COLGRETUMN

117

Figure 4-3:' VY Simplified Reactor Building Closed Cooling Water System
mm

sm.y

PCV 70

DENMI

WATER

TRAN&

Me N.

PRUNAR V

CONTAINMENT

AIR

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ALL VALVES RCW

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Figure 4-4: VY Simplified Reactor Building Closed Cooling Water System Loops 1 & 2
- UT

~j I

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-- IMATIM lmm TO 

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MOO mUss

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"I-

&UCm.

TO HM

KIWICTN

NI I I

I U I i

cr'wVS~zU -r - es' 

M we- 00

sywvrm

Figure 4-5: VY: Condensate Transfer System
00T PIL POOL MU MOKS

EASU FUEL POOL UMMin POK

I-' S

S

I

L.

10 LOQUID COGTIlml

HIILLED WATE EMW*MM6TAI

CANK PILLUS AMA

1riV fOILSE

CHEMCAL TATUET INS

STATOS WHOMN GCOOLUI WATEE

US COOLS VATEN SUME TANK

MS SAMPL SMN

CSNINCAL LAWS

AIR couOý PAUIHUS CowN

POST ACCOM BSDUM lYSTM

TO KfE gum

CONEEE

Figure 4-6: VY Demineralized Water Transfer System
IE I SFP GROSS I SFP WATER I SFP I SFP I SFP I

I STRUCTURAL I TIGHT I MAKEUP I COOLING I BACKUP I OUTCOME

I INTEGRITY I INTEGRITY I SYSTEM I SYSTEM I MAKEUP I

I I I iSYSTEM I

(A) I (B) i (C) I (D) I (E) I

------------------- 7-------------------------------------------------FALR

---- --- --- --- --- --- ---- --- --- --- --- --- --- FA ILURE

I Failure (Si)

---------- FAILURE

-------------------- (S2)

------------------ SUCCESS

------------ ---------- FAILURE

Success I--------- (S3) I. I I

I -------- SUCCESS

----- SUCCESS

I-------- FAILURE

I I ($4)

------------ SUCES

--- SUCCESS

I--------------------- SUCCESS

------------------- I-------- SUCCESS

II------I

I I-------- SUCCESS

SI J SUCCESS

Figure 4-7: VYNPS Event Tree

4-31
Na

Figure 4-8: VYNPS Fault Tree for Pool State 1
Fuel Pool Failure

For Pool State

2

(.Al

Figure 4-9: VYNPS Fault Tree for Pool State 2
1E 

I

W

SFP Gross

Structural

Integrity

x Y

SFP Water SFP

Tight (leak) Cooling

Integrity

z

SFP

Make-up

Water

Outcome

failure

(ASI)

failure

(AS2).

success

failure

(AS3)

success

success

Figure 4-10: H. B. Robinson Spent Fuel Pool Event Tree

4-34
so,

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H. S. ROINSON

UNIT 2

Cwollna POWS, & LighI Comwswy

UPDATErD FINAL SAFErTy ANALYSIS REPODRT

FLOW DIAGRAM

SPENT FUEL PIT COOLING

FIGURE |.I.3i I

Figure 4-11: Spent Fuel Pit Cooling

5. EQUIPMENT FRAGILITIES

In this chapter, seismic fragilities of equipment that comprise cooling and

makeup systems for the spent fuel pool are discussed. The seismic capacity of

the fuel racks is also evaluated.

5.1 Vermont Yankee Nuclear Power Station (VYNPS) Equipment

5.1.1 VYNPS Fuel Racks

The seismic capacity of the fuel rack assembly was derived using the design

information and results of analysis performed for the proposed replacement

racks [Vermont Yankee, 1986]. The racks are freestanding with no connections

to the spent fuel pool walls. The design of the racks has been performed to

the requirements of the Standard Review Plan Section 3.8.4. Racks are of cold

formed stainless steel honeycomb construction. There is a clearance of 2 3/8

in between the racks and the pool walls; the clearance between the racks is 2

3/4 in. The above reference [Vermont Yankee, 1986] has shown that the maximum

displacement during an SSE is 0.62 in which includes sliding, tilting, flexure

and rotation; sliding of the racks alone is 0.31 in.

Rack behavior under seismic loading is nonlinear and extrapolation of linear

analysis results is not very accurate. But, for purposes of a rough comparison

to other fragilities reported herein, such an extrapolation is made. It is

assumed that impact would occur when the sliding of the racks exceeds about 2

in. Hence, there is a factor of 2/0.31 = 6.5 over the SSE acceleration for

this to occur. Such extrapolation for other failure modes (tilting, flexure,

and rotation) is judged not appropriate. The median acceleration capacity of

the racks for incipient impact is estimated to be 1.0g. Failure of the racks

would require more than impact; perhaps, 1.5 to 2 times the acceleration is

required to cause impact and damage. Therefore, the earthquake peak ground

acceleration should exceed a value of 1.5 to 2.Og before failure occurs. Even

then, the fuel rack design is such that the assembly cannot be compressed into

a critical mass. According to the analysis reported in Chapter 4, no serious

effects can result (i.e., criticality and melt) as long as the fuel is

covered. Thus margins above SSE for impact and stress allowables are not

particularly meaningful. Therefore, it is concluded that crushing of fuel and

assemblies is a not a credible failure mode of the spent fuel pool system.

The cited reference also indicates that the maximum rack stress under SSE as

14.2 ksi where the allowable stress, S , is 25 ksi. The report does not state

the type of stress or the location. FoJ a free standing rack, the stresses are

expected to be flexural. Since the rack slides, the flexural stresses are

limited and will not increase with the increasing peak ground acceleration.

Therefore, failure of the rack by exceeding the flexural capacity is not a

credible failure mode.

The seismic failure of fuel pool racks and fuel assemblies is therefore not

modeled further in this risk analysis.

5-1
5.1.2 VYNPS Cooling and Makeup Systems

A preliminary list of components comprising these systems is provided in

Table 4.2. The generic categories of components in this list include motor

control centers, switchgear, pumps, tanks, valves, heat exchangers, filters

and strainers.

During a visit to Vermont Yankee Station, the focus was only on the spent fuel

pool structure and a detailed examination of the cooling and makeup systems

was not performed. However, a brief walkdown of these systems was conducted;

some photographs were taken and notes on the existing conditions of the

equipment were made. Detailed equipment drawings showing anchorage and

support details were not reviewed for this evaluation.

The equipment fragilities were estimated using generic information and based

on the applicable design codes. A brief description of the procedures used in

this fragility evaluation is given. First, the deterministic capacity called

the High Confidence Low Probability of Failure (HCLPF) capacity is

estimated. Assuming the values of PR and PU to be 0.3 and 0.4 respectively,

the median acceleration capacity, Am, of the component is calculated.

Motor Control Centers

It was assumed that the motor control centers were all located at Elevation

303'. The peak spectral acceleration was compared to the Generic Equipment

Ruggedness Spectrum GERS, [EPRI 1987] at 5% damping using a reduction factor

of 1.3 as recommended in the EPRI Seismic Margins Methodology [NTS, 1987].

The HCLPF capacity corresponding to a non-recoverable failure mode was

estimated to be 0.27g. The inspection of the anchorage photos indicated that

there are 10 cabinets in a row. Assuming 2 bolts per cabinet of 1/2" diameter

non-shell type, the spectral acceleration capacity of the cabinet against

anchorage failure was estimated from [EPRI, 1987] to be 1.7g. The

corresponding HCLPF capacity of the cabinet was derived as 0.16g. The median

capacity of the cabinet was estimated as 0.50g. Anchorage failure is the

lower of the two failure modes i.e., cabinet/device failure and anchorage

failure. In the above calculation, only two bolts were assumed to be

installed per cabinet; however, if 4 bolts of 3/8" nonshell type were

installed instead, the median capacity would be about 0.72g with a HCLPF of

0.23g which is much closer to the generic capacity of the MCC units as defined

by GERS.

Station Service Transformer # 8

There are 4 1/2" diameter expansion anchors on this transformer. The bolts are

subjected to bending due to the presence of shims below the transformer. The

weight of the transformer was estimated to be 6.3 kips. The HCLPF capacity of

the anchorage was evaluated using the interaction formula for bending and

shear. The HCLPF capacity obtained was 0.16g and the median ground

acceleration capacity was 0.51g.

5-2
Station Service Transformer # 9

A review of the photographs of the item indicated that one anchor bolt was

missing but unlike SST #8, the flange of the transformer is bolted directly to

the floor without any spacer being provided. The HCLPF capacity of this

transformer is considered to be higher than #8 and is 0.28g with a median

capacity of 0.89g.

4 kV Switchgear/Bus

The photograph shows that angles have been added to the inside of the cabinet

which were fastened to the floor with 3/8" expansion anchors. The cabinet is

2' x 5' and is 7' tall. The anchorage in the back was not observed. Assuming

non-shell anchors, the spectral acceleration capacity was estimated to be

1.6g. The HCLPF capacity and the median capacity were estimated to be 0.14g

and 0.44g respectively.

RBCW Surge Tank

The photograph shows that the tank is mounted on 4 legs resting on concrete

pads. There is lateral bracing but no longitudinal bracing is present. The

tank is about 3.5 ft in diameter and 5.5 ft long. The total weight of the

tank with water and framing is estimated to be 2.5 kips. The tank support is

fairly flexible but the strength of the anchorage could not be ascertained

from the photograph. The diagonal bracing looked minimal. Two failure modes

were investigated: failure of the diagonal bracing and overturning. The

overturning failure mode was found to govern the capacity. The HCLPF capacity

was estimated to be 0.09g; the median capacity was calculated as 0.29g.

RB CCW Heat Exchanger

This is a large horizontal tank. The photograph showed that it is supported

on saddles which are adequately reinforced in the longitudinal direction. On

the fixed end, there are 4 bolts of probably 1/2" diameter. Assuming the

HCLPF capacity corresponds to the earthquake level when the shear stress

exceeds the allowable value of 0.6 Sy, it is estimated to be 0.46g and the

median capacity as 1.46g.

Spent Fuel Pool Heat Exchanger

This heat exchanger has an over and under horizontal cylinder arrangement.

The photographs did not show details of the anchorage. There does not appear

to be any top brace on the assembly. Lateral overturning appears to be the

critical failure mode. Field notes indicated that there is one 3/14" bolt in

each support leg. Assuming that the bolts are cast in place, the HCLPF

capacity is estimated to be 0.23g and the median capacity as 0.73g.

5-3
RHR Heat Exchanger

Based on a comparison of this component with the heat exchangers reviewed in a

number of seismic PRAs, the HCLPF capacity is estimated to be larger than 0.5g

with the median capacity in excess of 2.0g.

RHR Pumps

The HCLPF capacity of this component is estimated larger than 0.5g based on a

comparison to similar pumps in the seismic PRA database.

Fuel Pool Demineralizer Tank

The demineralizer tank is located low in the radioactive waste building. It

has four legs of 3" angle and is bolted to the floor using 7/8" bolts. The

governing failure mode is cantilevered bending of the angle legs. The HCLPF

capacity is estimated to be 0.14g with the median capacity about 0.44g.

Piping and Valves

Based on the extensive PRA and experience data base, the HCLPF capacities of

piping and valves were estimated to be in excess of 0.5g.

Table 5.1 gives a summary of equipment fragilities for the spent fuel pool

cooling and makeup systems.

5.2 H. B. Robinson, Unit 2 (HBR2) Equipment

5.2.1 HBR2 Cooling and Makeup Systems

A list of components comprising these systems was prepared and shown in

Table 4.3. The generic categories of components in this list include motor

control centers, switch gear pumps, tanks, valves, heat exchangers, and

instruments.

During a visit to the Robinson plant, some of these items were examined with

the major focus on the spent fuel pool structure. Based on a plant walkdown,

information available in the updated FSAR, and equipment anchorage drawings,

most of the equipment items were screened out as having high seismic

capacities. This was also guided by the seismic fragility parameters of

component categories reported in [Budnitz et al., 1985] and [Campbell et al.,

1985].

Fragility estimation were made for the following equipment items:

CCW Surge Tank

Demineralizer Tank

Refueling Water Storage Tank

5-4
480v MCC No. 1

480v MCC No. 8

Condensate Storage Tank

The equipment fragilities were estimated using generic information and based

on applicable design codes. In the following, we describe the procedure used

and the results of the fragility evaluation of these components; for

components other than the ground mounted yard tanks, the deterministic

capacity called the High Confidence Low Probability of Failure (HCLPF)

capacity is estimated. Assuming the values of ýR and 6 to be 0.3 and 0.4

respectively, the median acceleration capacity, Am, of The component is

calculated.

CCW Surge Tank

This is a steel tank 4 ft in diameter and 14 ft long; it is mounted on the

roof of the reactor auxiliary building using two saddle supports. The saddles

are bolted down to the floor with two 3/4 in. bolts on each saddle. The

weight of the tank including water is estimated to be 13 kips. The surge tank

is judged to be rigid hence the maximum acceleration acting on it at the SSE

is estimated to be 0.40g; at this load, the shear stress in the bolts is

calculated as 4.1 ksi; comparing this stress with the allowable shear stress

of 0.6 Sy = 21.6 ksi, the HCLPF capacity is estimated as 1.05 g. The median

ground acceleration capacity Am, is estimated to be 3.31g.

Demineralizer Tank

The demineralizer tank is located at Elevation 226'0" in the Auxiliary

Building. It is a 32" diameter steel tank mounted on four legs of 41" angle

and bolted to the floor using 1" bolts. The governing failure mode is the

cantilevered bending of the angle legs. The HCLPF capacity is estimated to be

0.44g with the median capacity to be 1.41 g.

Tanks

Yard tanks which were identified by the systems analysis for the Robinson Unit

2 spent fuel pool risk analysis are the refueling water storage tank (RWST)

and the condensate storage tank (CST). In the following, the evaluation of

the CST is described in detail.

The condensate storage tank is a 35 ft diameter steel tank fabricated using

1/4 in. stainless steel (A 240 Type 304) plate; it is anchored to the

foundation reinforced concrete ring wall with 12 anchor bolts of 1-3/4 in.

diameter. Maximum depth of water is 28 ft 3 in. The median ground response

spectrum used for this evaluation is NUREG/CR-0098. Since the tank is at

grade on a soil site, a median damping is increased from 7% used typically for

welded tanks founded on rock to 10% to account for soil structure interaction

effects. Based on previous PRA studies, the critical failure mode of the tank

is identified to be the overturning moment failure. The median overturning

5-5
moment capacity of the tank is estimated to be 20,930 ft-kips. This includes

the resistance provided by the tank weight, fluid weight, base plate, shell

buckling and anchor bolts. The median overturning moment at the 0.2 g SSE

level is calculated to be 6,681 ft-kips resulting in a median strength factor

equal to 20,930/6,681 = 3.1. The uncertainty in this strength factor is

estimated taking into account the uncertainties in the shell buckling

coefficient, location of the anchor bolt force, bottom plate holding force,

yield strength of anchor bolts and modeling of the failure mode. This value

is PR = 0.18. This overturning moment failure is somewhat ductile and a

median inelastic energy absorption capacity factor of 1.1 is estimated.

It is assumed that the average of the two horizontal components is used in

.specifying the median ground response spectrum. Because of the axisymmetry of

the tank, no matter how the two horizontal components are oriented the tank

maximum response will result from the larger of the two horizontal components

rather than the average of the two. Therefore, a factor of 0.9 is estimated

to reduce the tank capacity calculated on the basis of the spectrum averaged

over two components.

The other safety factors are estimated both in terms of the median values

and PR and pU values as explained for the spent fuel pool structure and are

tabulated in Table 5.2.

The refueling water storage tank (RWST) at Robinson Unit 2 is a 42 ft diameter

tank fabricated from 1/4 in steel plates; it is anchored to the ring wall

using 20 anchor bolts of 1-3/4 in. diameter. The maximum water height in the

tank is 34 ft 3 in. By comparing the overturning moment capacities of RWST and

CST, the median ground acceleration capacity of the RWST is estimated to be

0.60g with PR = 0.21 and PU = 0.27.

Motor Control Centers

It was assumed that the motor control centers are all located at Elevation

275' 0". From the Updated FSAR, the maximum floor spectral acceleration in

the auxiliary building is 2.19 g. This spectral acceleration was compared to

the Generic Equipment Ruggedness Spectrum GERS [EPRI 1987] at 5% damping using

a reduction factor of 1.3 as recommended in the EPRI Seismic Margins

Methodology ENTS, 1987]. The HCLPF capacity corresponding to a nonrecoverable failure mode was estimated to be 0.28 g. As a representative MCC,

we reviewed the MCC No. 6; it has 10 cabinets in a line with a total of six

1/2" diameter shell anchors. The lineup is supported at the top through three

struts made up of two equal 2 1/2" angles. The struts are bolted to the side

wall using two anchor bolts of 1/2" diameter. This anchorage and lateral

support are considered to be sufficiently rugged that the structural failure

of the cabinets is not a critical mode. The median ground acceleration

capacity of the MCCs is estimated to be 0.89 g based on the GERS discussed

above.

Table 5.3 gives a summary of the equipment fragilities.

5-6
Table 5.1

Summary of Equipment Fragilities

for VYNPS SFP

Component Am U HCLPF Notes

(g) Capacity

(g)

Motor Control Centers 7b,7c,8b,9b 0.50 0.30 0.40 0.16 1

Station Service Transformer 8 0.51 0.30 0.40 0.16 2

Station Service Transformer 9 0.89 0.30 0.40 0.28 5

4 kV Switchgear 0.44 0.30 0.40 0.14 3

RB CCW Surge Tank 0.29 0.30 0.40 0.09 4

SFP Demineralizer Tank 0.44 0.30 0.40 0.14 8

RB CCW Heat Exchanger 1.46 0.30 0.40 0.46 5

SFP Heat Exchanger 0.73 0.30 0.140 0.23 6

RHR Heat Exchanger >2.00 >0.50 7

RHR Pump >2.00 >0.50 7

Piping >2.00 >0.50 7

Valves >2.00 >0.50 .7

1. Assumes 2 non-shell 1/2" bolts per cabinet

2. Anchor bolts in bending and shear

3. Assumes 4 non-shell 3/8" bolts per cabinet

4. One 3/8" bolt per leg from photographs; assumed non-'shell expansion anchors

5. Bolt shear; all dimensions from photographs and notes

6. 3/14" diameter 4 bolts per field notes; assumed to be cast in place

7. Based on PRA results

8. Support leg bending

5-7
Table 5.2

Median Factors of Safety and Variabilities for HBR2 CST

Factor Median

Strength 3.1

Inelastic'Energy Absorption 1.1

Spectral Shape 1.0

Damping 1.0

Modeling 1.0

Soil-Structure Interaction 1.0

Modal Combination 1.0

Combination of EQ Components 1 .0

Horizontal Component Direction

TOTAL 3.1

Median PGA capacity = 3.1 (0.20g) = 0.62 g

HCLPF Capacity = 0.62 exp (- 1.64 ( 0.21 + 0.27

0

0.02

0.20

0.03

0

0.01

0.03

0.01

0.9

0.21

) 0.28 g

oU

0.18

0.02

0.00

0.1 4

0.10

0.10

0

0

00

0.27

5-8
Table 5.3

Summary of HBR2 Equipment Fragilities

Component gAm . HCLPF

(g) Capacity

(g)

CCW Surge Tank 3.31 0.30 0.40 1.05.

Fuel Pool Demineralizer Tank 1.41 0.30 0.40 044

Refueling: Water Storage Tank 0.60 0.21 0.27 0.27

Motor Control Centers No. 1 and 8 0.89 0.30 0.40 0.28

Condensate Storage Tank 0.62 0.21 0.27 0.28

5-9

6. RISK EVALUATION

6.1 Objective

The objective of this study is to estimate the probability of failure of the

Spent Fuel Pool due to seismic events. Failure of the pool is defined to be

gross rapid loss of pool contents that would empty all the cooling water and

expose the fuel rods. It has been determined that the failure of cooling and

make-up systems would not result in immediate uncovering of the fuel rods.

There is a response time of at least 3 days, maybe as much as 7 days before

fuel uncovery would occur. Thus, SFP failures attributed to these system

failures and are not directly comparable to SFP failures caused by structured

degradation leading to sudden loss of water. The damage to fuel racks in a

seismic event are judged to be minor and fuel rack damage is not significant

as long as there is water in the pool.

The estimation of probability of failure is accomplished by using the seismic

probabilistic risk assessment (PRA) methodology, described earlier in Chapter

2.

Preliminary seismic hazard curves for the two sites are shown on Figures 6-1

and 6-2. These curves correspond to the confidence values of 15%, 50%, and

85%. Note that these preliminary curves have not been finally reviewed by the

NRC and are used herein only to get a preliminary understanding of the seismic

induced spent fuel failure. The hazard curves may change after NRC review and

guidance for their proper use will be developed.

Using these hazard curves, a set of discrete hazard curves with associated

subjective probability weights were generated. This was done by fitting a

lognormal distribution to the uncertainty in the hazard at any peak ground

acceleration value. This lognormal probability distribution was assumed to

have the median value of the hazard (i.e. annual frequency of exceedance)

given by the 50th percentile curve and the 95th percentile value given by the

top most curve in the figure. This probability distribution was discretized

into 11 ranges with probabilities of 0.03, 0.05, 0.07, 0.12, 0.15, 0.16, 0.15,

0.12, 0.07, 0.05, 0.03. Note that the lognormal probability distribution was

cutoff at the 99th percentile value. By repeating this procedure at other

accelerations, a family of hazard curves was generated.

6.2 VYNPS Analysis

6.2.1 VYNPS Assumptions

The following assumptions have been made in the seismic risk calculations

reported herein:

Seismic fragility of the spent fuel pool structure was estimated earlier using

detailed information on the structure. The median capacity with all high

density racks is estimated to be 1.4 g with the HCLPF capacity of 0.50 g. For

the existing fuel rack configuration, the HCLPF capacity is estimated to be

0.65g. Case 1 shows the annual failure frequencies of the pool structure.

6-1
Seismic fragilities of mechanical and electrical components have been

estimated by using both site-specific information from walkdowns, as well as

the walkdown notes and photographs, and generic information on the component

categories from previous PRAs. The HCLPF capacities were estimated using

deterministic methods and the median capacities were calculated assuming PR

and 8U to be 0.3 and 0.4, respectively.

6.2.2 VYNPS Accident Sequences and Failure States

Seismic fragilities and random failure rates of different components relevant

to this analysis are shown in Table 6.1. Based on the systems analysis

presented in Chapter 4, two pool states were developed, PSI and PS2.

The Boolean expressions for these two pool states are given by:

PS1 = 12 + (7 + 11 + *2) *( 5*{ 4 + 9 + 8 + 10

+4 * { 3+6})

PS2 = 12 + (7 + 11 + *2) *( 4 + 9 + 10)

The quantification of these Boolean expressions was accomplished using the

component seismic fragilities and random failure rates given in Table 6.1.

The fragilities were assumed to be perfectly independent both in the

randomness and in the uncertainty aspects.

It may be noted that the components 1, 2, 7 and 11 are all in the non-safety

electrical systems of the plant. As shown in Table 6.1, their low capacities

seem to contribute significantly to the seismic induced failure of the spent

fuel pool support systems.

6.2.3 VYNPS SFP Case Studies

Case studies 1, 2 and 3 were conducted using the seismic fragilities estimated

as shown in Table 6.1, case definitions are given in Table 6.2. They also

include the operator failure rates or random failures. The influence of

random failures on the seismic system failure frequencies is shown in case

studies 4 and 5.

The anchorage of the MCC cabinets could not be determined by the walkdown

notes and photographs; therefore it was assumed in case studies 2 through 5

that the MCC cabinets were anchored using 2-1/2 in non-shell expansion anchors

per cabinet and the seismic fragility parameters Am and 3S were estimated.

Sensitivity case studies 6 and 7 were conducted that assumed MCC cabinets had

4-3/8 in non-shell anchors per cabinet. The median capacity was estimated to

be 0.82 g as contrasted with the earlier value of 0.50 g. There was not a

large change in the failure probability because of the dominance of the

cabinets as seen in Table 6.2.

6-2
6.2.4 Probability of VYNPS Spent Fuel Pool Failure

The total probability of pool failure is determined by combining the

probabilities of the two Pool States, PSI and PS2, based on the time the plant

is in that mode of operation (normal operation or complete core discharge to

the spent fuel pool). For this analysis, it was assumed that the plant could

have a complete core discharge to the pool once every five years and that the

full core could be in the pool for a duration of one month. The analysis is

given below:

Let:

P(PS1) = probability that fuel pool is in pool state 1, per year

P(PS2) = probability that fuel pool is in pool state 2, per year

P(fjPSI) = probability of fuel pool failure while in pool state 1,

per year

P(fJPS2) = probability of fuel pool failure while in pool state 2,

per year

then, P(f), the total probability of fuel pool failure is given by

P(f) = P(fJPSI) P(PS1) + P(fjPS2) P(PS2)

P(PS1) and P(PS2) have been estimated to be 59/60 and 1/60 respectively. The

mean conditional probabilities for different failure states are given in Table

6.2. The results of this analysis indicate that for:

Case 2 & 3 P(f) = 1.5 E-4

Case 4 & 5 P(f) = 1.2 E.-4

Case 6 & 7 P(f) = 9.4 E-5

6.3 H. B. Robinson, Unit 2 (HBR2) Analysis

6.3.1 Probability of HBR2 Spent Fuel Pool Failure

Seismic fragilities and random failure rates of different components relevant

to the HBR2 analysis are provided in Table 6.3.

Gross spent fuel pool structural failure results in loss of the pool liner

integrity and complete loss of the pool water inventory in a short time

leading to fuel damage, possible cladding fire and radioactivity release. The

seismic fragility for this failure mode is discussed in Section 3.3.2

Fuel uncovery, damage and possible radioactivity release occurs from failure

to cool the water in the pool resulting in water boil-off and failure to

provide make-up water in the event cooling fails. If pool cooling is

successful or the pool make-up systems are available, then SFP success is

maintained. The Boolean expression is :

6-3
Pool Failure = 1 + ( 4 + 5 + 7 + 9 ) * ( 6 + 8 + 12 )

The quantification of this Boolean expression was done using the component

fragilities and random failure rates given in Table 6.3. The fragilities were

assumed to be perfectly independent both in the randomness and in the

uncertainty aspects. The results of the qualification are given in Table 6.4.

6-4
Table 6.1

Seismic Fragilities and Random Failure Rates for VYNPS

COMPONENT FRAGILITIES Am aR aU

#1

#2

#3

#4

#5

#6

#7

#11

#8

#9

#10

MCC 7b FAILS (SS TRANSFORMER 7)

MCC 7c FAILS (SS TRANSFORMER 7)

MCC 8B FAILS (SS TRANSFORMER 8)

MCC 9B FAILS (SS TRANSFORMER 9)

RBCCW SURGE TANK TK 7-1A FAILS

SFP HEAT EXCHANGER E-9-1A FAILS

STATION SERVICE TRANSFORMER

SWITCHGEAR

OPERATOR FAILURE RATES

OPERATOR FAILS TO OPEN MOV V-89A

OPERATOR FAILS TO OPEN MANUAL VALVE V-25

OPERATOR FAILS TO OPEN MOV V-50a

0.50 0.30

0.50 0.30

0.50 0.30

0.50 0.30

0.29 0.30

0.73 0.30

0.50 0.30

0.44 0.30

RATE/DEMAND

0.04

0 . 4 0

0.40

0.40

0.40

0.40

0.40

0.40

0.40

ERROR FACTOR

3.0

3.0

3.0

0.08

0.08

6-5
Table 6.2

Results of the Analyses

of VYNPS SFP

Case Annual Seismic Failure

Frequencies (per reactor-yr)

Mean 5% 50% 95%'

1. Pool Structure only

Am= 1.4 g; HCLPF= 0.5 g

(a) Cutoff at X= 1.0

(b) Cutoff at 99%

(c) Cutoff at 98%

(d) Cutoff at 97%

2. SFP Support Systems Failure

State PSI

3. SFP Support Systems Failure

State PS2

4. SFP Support Systems Failure

State PS1 w/o random failures

5. SFP Support Systems Failure

State PS2 w/o random failure

6. SFP Support Systems Failure

State PS1 w/revised MCC

fragilities

7. Pool State PS2 with

revised MCC fragilities

3.8E-05

6.7E-06

4.9E-06

3.8E-06

1 .5E-04

1 .7E-04

1 .2E-04

1 . 1 E-04

9.4E-05

9.OE-05

1 .8E-1 1

3. 1 E-1 1

3.1E-1 1

3.1E-1 1

7.7E-08

8.3E-08

8.OE-08

7 .7E-08

3.%6E-05

1 .9E-05

1 .9E-05

1 .4E-05

2.OE-07 1.2E-05 6.7E-04

2.3E-07 1.3E-05 6.7E-04

1.2E-07 9.5E-06 5.3E-04

7.1E-08 8.1E-06 4.6E-04

1.OE-07 6.6E-06 3.9E-04

1.1E-07 6.2E-06 3.2E-04

Note: Cases 2 through 4 do not include alternate makeup.

Failure to align normal makeup is considered.

6-6
Table 6.3

Seismic Fragilities for HBR2

Component Fragilities Am R aU HCLPF

(g) (g)

#4 Fuel Pool Demineralizer 1.41 0.30 0.40 0.144

Tank

#5 Refueling Water Storage 0.60 0.21 0.27 0.27

Tank

#7¶" 480v Motor Control Center 0.89 0.30 0.40 0.28

No. 1

#8 480v Motor Control Center 0.89 0.30 0.40 0.28

No. 8

#12 Condensate Storage Tank 0.62 0.21 0.27 0.28

OPERATOR FAILURE RATES RATE/DEMAND ERROR FACTOR

#6 Operator fail to align 0.08 3.0

CST for pool makeup

#9 Operator fails to align 0.08 3.0

RWST for pool makeup

6-7
Table 6. 4

Results of the Analyses

of HBR2 SFP

Case Annual Seismic Failure

Frequencies (per reactor-yr)

Mean 5% 50% 95%

1. Pool Structure only

Am = 2.0 g; HCLPF = 0.65 g

(a) Cutoff at X= 1.0

(b) Cutoff at 99%

(c) Cutoff at 97%

8.6E-06

1 .8E-06

9.9E-07

6. 1E-1 2

9.9E-1 2

9-5E-1 2

1. 3E-08

1 .5E-08

1 . 4E-08

1 .9E-05

8.6E-06

5.OE-06

3.5E-06

2. SFP Support Systems Failure 1.5E-04 5.5E-07 6.OE-04

Note: Case 2 does not include alternate makeup.

Failure to align normal makeup is considered.

6-8
E..U.S SEISMIC HAZARD CHARACTERIZATION

LOWER M•AGNITUDE OF INTEGRATION IS 5.0

PERCENTILES = 15., 50. AND 85.

HAZARD CURVES USING ALL EXPERTS

L.J Cr

0-

L.J

LIJ

LIJ

0

I-.

l-J

¢0

a-

-1

10

-2

10

-3

10

-4

1 0

-5

1.0

-6

10

-7

to '- e N on "q tn t N C3

÷w ACCELERATION CM/SEC#02

VERMONT YANKEE

Figure 6-1: Seismic Hazard Curves for the VYNPS Site

6-9
E.U.S SEISMIC HAZARD CHARACTERIZATION

LOWER MAGNITUDE OF INTEGRATION IS 5.0

PERCENTILES = 15.. 50. AND 85.

HAZARD CURVES USING ALL EXPERTS

LX

w

¢L

0

Lo

41

CD)

0

ix

O0

-1

10

-2

10

-3

10

1 0

-5

10

-6

10

-7

10

0

ACCELERATION CM/SEC'-2

ROBINSON

Figure 6-2: Seismic Hazard Curves for the HBR2 Site

6-10
7. CASK DROP ANALYSIS

7.1 Background

Cask drop analyses were performed as part of this study because the BNL study

indicated sigificant uncertainty as to whether damage would occur. The spent

fuel pools at Vermont Yankee and H.B. Robinson were analyzed for their

response to cask drops. The purpose of this study is to assess the affects of

a shipping cask drop on a spent fuel pool wall, assuming that the VY and HBR

wall designs (thickness, strength and reinforcement) are typical. The walls

selected for analysis were determined after discussion with the NRC and were

based on the fuel pool configuration, and the path the casks would follow when

crossing the pool. The cask weights and drop heights were obtained from

actual plant operating criteria. The following cask drops were analyzed:

Vermont Yankee - "40 ton 6" Drop

110 ton 6" Drop

H. B. Robinson - 68 ton 4' Drop

7.2 Analysis Methodology

Characterization of the nonlinear behavior of reinforced concrete is currently

a topic of intense academic research. Attempts at analytically modeling the

complicated interaction between reinforcing steel and concrete and the ratedependence of concrete constitutive relationships have had limited success,

and there is no one universally accepted model for nonlinear concrete finite

element analysis. Efforts in this area have been hampered by the lack of an

acceptable failure criteria for tri-axially loaded concrete. According to

[Winter and Nilson 1979]:

"In spite of extensive and continuing research, no general theory of the

strength of concrete under combined stress has yet emerged. Modifications of

various strength theories, such as the maximum-tension stress,ý the MohrCoulomb, and the octahedral-stress theories, all of which are discussed in

strength-of-materials texts, have been adapted with varying degrees of success

to concrete".

These observations were made with regard to concrete under static, monotonic

slow strain rate loading. In the case of extreme environmental loads

(seismic, blast, impact), with resulting high strain rates, the difficulties

are magnified. It is noted therefore that in absence of experimental testing,

the estimation of damage in a reinforced concrete structure can be very

subjective. Concrete design has necessarily been a mixed bag of structural

mechanics fortified with empirical data from experimental tests. Predictive

capabilities of the ultimate strength of typical reinforced concrete

structures such as simple beams and columns is fairly well understood but more

exotic or lesser used structural systems are difficult to design and analyze.

7-1
For the impact of the spent fuel cask on the wall, the load on the wall is inplane and the configuration is essentially a deep beam subjected to transverse

load. The American Concrete Institute [ACI 19771 has design formulas for deep

reinforced concrete beams. The design formulas are based on standard Whitney

stress block formulas. The underlying assumption of the ultimate strength

formulas is that plane sections of the beam remain plane after loading. As a

beam becomes very deep, the plane section assumption becomes less valid. The

ACI code places limits on the aspect ratio (span/depth) outside the range of

applicability of the ACI formulas, the code states that the beams shall be

designed "taking into account the nonlinear distribution of strain", however,

the code does not provide guidance on how this should be performed. The

design profession has handled deep beams by utilizing elasticity solutions to

estimate the stress distribution in very deep beams [L. Chow, et al., 1977 and

Design of Deep Girders, 1977]. The stresses from elasticity solutions are

integrated to determine the location and magnitude of stress resultants and

steel is apportioned in order to supply tensile stress resultants. A check of

the aspect ratio of a typical pool wall shows that they are very deep beams

and are below the range of application of simple ACI formulas. In addition,

from finite element impact analysis (discussed in detail below).the stress

distribution is quite complex and changes continually during impact loading.

Consequently, a simple application of ACI type formulas or use of elasticity

solutions which have been derived for static loads is inadequate. Information

on the experimentally determined strength of moderately deep reinforced

concrete beams can be found in the literature [K. N. Smith et al., 1982, and*

H. A. Rawden et. al 1965] however, the data is restricted to static loadings

and the aspect ratio of the tested beams are much shallower than the fuel pool

walls. Work on impact loading of reinforced concrete beams has been conducted

by Loo and Santos [Y. C. Loo and A. P. Santos, 1986]. The authors developed

analytical expressions for beam response to an impact load at mid-span of a

pinned-pinned beam. The results are only applicable in the linear response

range and the derived expressions are for thin beams (i.e., transverse

deflections dominated by flexure rather than shear) and consequently are not

directly applicable to the current problem. Extensive effort has gone into

developing empirical formulas for missile impact on reinforced concrete (an

excellent summary is given by Sliter [G. E. Sliter, 1980]). These formulas

deal primarily with impact of small missiles and local impact response such as

cratering, crushing, etc., and do not address the global response of the

structure.

All of these cited works and design formulas are not applicable to the

particular geometry and loading of the spent fuel wall impact problem.

Consequently one must resort to a numerical analysis to obtain estimates of

the damage the wall sustains as a result of impact.

A numerical methodology which has been used by practicing engineers to

approximate the behavior of concrete structures is to perform a finite element

analysis in which the reinforced concrete structure is idealized as a linear

elastic structure [A. Davito et al., 1978, and Personnel Communication

1979]. When this approach is used damage estimates under extreme loads can'be

made based on the stresses obtained from the linear analysis. It is

7-2
understood that this method does not account for stiffness degradation due to

concrete/steel damage nor does it account for energy dissipation due to

inelastic action. However, because of limitations on state-of-the-art

nonlinear finite element techniques, more rigorous analysis may not be

justified. For the current study it was decided to exploit the contact

surface algorithms in existing nonlinear finite element codes in order to

rigorously model the impact of the cask on the wall. The wall is taken as a

homogenous, isotropic, linear elastic structure. Based on the stresses from

the linear elastic analysis, judgements on wall damage can be made. All

analyses were conducted with NIKE2D [J. Hallquist 1986], a nonlinear finite

element program with the capability to model contact/impact problems. Further

simplification of the problem is obtained by assuming the impacting cask is

rigid.

7.3 Cask Drop Analysis of Vermont Yankee Fuel Pool

7.3.1 Finite Element Model

A two-dimensional finite element model was constructed for the south wall of

the Vermont Yankee spent fuel pool (see Fig. 7-1). This wall was selected as

a worst case example. The wall was treated as a plane stress structure with

the appropriate thickness. Consideration of all possible load combinations

was beyond the scope of the current study, thus dead weight, hydrostatic loads

and thermal loads were not included in the analysis. Consideration of

hydrostatic and thermal loads requires a full 3-D analysis and is more

computationally intensive than the 2-D analysis which have been performed.

The stresses caused by the cask impact are much larger than the stresses due

to the self-weight of the wall, and the impact load stresses different regions

of the wall than does the gravity load. In light of this, it is felt that

inclusion of the gravity load component would not substantially change the

damage estimates. The boundary conditions were taken as fixed at the outer

corners of the wall where the north-south running beams support the pool

structure. There is also an east-west running beam under the south wall.

This beam was not included in the finite element model, which essentially

neglects its stiffness. Because of the nature of the dynamic response, and

the fact that the in-plane stiffness of the pool wall is so great, neglecting

the beam for impact loading is acceptable. The ends of the wall in the finite

element model are left free and thus any stiffening provided by the northsouth walls and floor diaphragms is neglected. For the finite element

analysis, the wall is taken as a linear elastic material with an elastic

modulus of 57000 (f)Y 2 . The original design strength of the concrete was f'-

cc 4000 psi (12). In-situ testing of the concrete in the pool bottom showed c

present day strength of about 7000 psi. Because of uncertainty in wall

concrete strength (it was unclear if the concrete in the floor was the same as

the wall concrete), the 4000 psi value was used for conservatism. Two cask

weights were considered, the first cask was taken to be 110 tons, the second

cask was 40.tons. Both casks were assumed dropped from a height of 6 inches

(the operating criteria dictates this as the maximum height the cask can be

lifted (7)). The cask is modeled as a solid steel body with a width of 8

feet.

7-3
The geometrically nonlinear portion of the problem (i.e., the impact in which

the contact surface is changing) was modeled with the NIKE2D. contact

algorithm. The equations of motion were integrated at a time step of .0005

seconds, based on some limited sensitivity analyses this time step appears

sufficiently short to capture the problem dynamics.

7.3.2 110 Ton Cask Drop

The displacement of the system during impact is summarized in Fig. 7-2 which

shows the vertical displacement of a point at mid-span and mid-depth of the

wall. Also shown is the displaced shape of the wall from initial impact until

the cask rebounds. In order to estimate the damage which the wall incurs, the

contours of principal stresses and vertical and horizontal strains were

plotted for each time step in the response. The principal stress values were

used to estimate regions of potential concrete cracking. While a trn-axial

failure criteria for concrete is not currently available, a bi-axial failure

envelope has been established. This failure envelope is shown in Fig. 7-73.

In the region where both principal stresses are compressive, the failure

criteria is simply the exceedance of the concrete compressive stress by either

stress component (i.e., quadrant iii in Fig. 7-3). From Fig. 7-3 it is

observed that if either principal stress component is positive and greater

than the modulus of rupture of the concrete (taken as 4 (f )'2) then a tension

crack will develop. Figure 7-4 shows regions of the wall in which the

principal stress is positive and larger than the modulus of rupture at some

time during the impact analysis. This figure indicates that virtually the

entire wall is subject to tensile cracking at some point during the

response. Minor cracking is generally not a significant problem and this does

not necessarily indicate any catastrophic failure or even loss of water from

the pool. The pool has a thin metal liner which provides containment of the

pool water. If the liner remains intact, the water will be contained despite

concrete cracking. Another indicator of wall damage is the possibility of

yielding of the reinforcing steel. In order to estimate the regions in which

the reinforcing steel might yield during impact loading, contour plots of the

vertical and horizontal strains at each time step were used. The contour

values were set equal to the yield strain values of the steel which allowed a

direct view of yield regions at each time step. The contour plots for each

time step were then superimposed, to show cumulative damage as indicated in

Fig. 7-5 and Fig. 7-6. The results of these overlays are shown in Fig. 7-7

and Fig. 7-8. These plots indicate the regions of the wall where the

horizontal and vertical strains exceed the strain which reinforcing steel

would have at initial yield at some time during the impact response (yielding

in different regions does not occur simultaneously). These plots allow an

estimation of areas in which the reinforcing steel yields and thus significant

structural damage may occur.

7.3.3 40 Ton Cask Drop

The impact analysis was repeated for a 40 ton cask, all other parameters

remained the same. The response time history of the wall is shown in

Fig. 7-9, where the displacements are significantly less than the 110 ton cask

7-4
case (Fig. 7-2). As with the 110 ton cask the maximum principal tensile

stress exceeds the modulus of rupture throughout the wall and thus concrete -

cracking potentially occurs throughout the wall. The estimated regions of-'

reinforcing steel yielding are indicated in Fig. 7-10 and Fig. 7-11. For the

40 ton cask the vertical steel region of yield is still substantial, but the

region of yield of horizontal steel is greatly reduced from the 110 ton

case. The reason for the significant reduction in the horizontal steel

yielding appears to be due to the fact that the wall does not experience such

large overall deflections'for the 40 ton case (compare displaced shape plots

in Fig. 7-9 and Fig. 7-2 at the maximum response times) and the yielding of

the horizontal steel occurs as a result of the global or overall deflections

of the wall. The vertical steel yielding, on the other 'hand, occurs in the

early stages of impact as the impact stresses propagate through the depth of

the wall, prior to the overall response of the wall.

7.4 Cask Drop-Analysis of H.B. 'Robinson Fuel Pool

7.'4.1 Finite Element Model

A 68 ton cask was considered dropped from a height of 4 feet onto the

reinforced concrete pool wall. A plane stress finite element model was

constructed for both the cask and the pool wall (see Fig. 7-12). The hole in

the wall is the entrance to the waste gas decay tank area. Dead weight,

hydrostatic, and thermal loads were not included in the analysis. The cask

was given an initial velocity equal to that which would be attained in a four

foot drop.

7.4.2 Results of Analysis

The time history response of the wall/cask system, expressed in terms of the

vertical displacement of a node on the wall directly under the cask, is shown

in Fig. 7-13. Also shown in Fig. 7-13 are the displaced shapes of the wall at

selected time steps. The wall is observed to "crush" directly under the

cask. In the last displaced shape plot, the cask has bottomed out and is in

the process of rebounding. As with the Vermont Yankee pool, cumulative damage

estimates of the concrete and reinforcing steel were obtained by superimposing

strain plots from each time step (see Fig. 7-14 and 7-15).. The strain

contours which were plotted were set equal to yield strain values for steel.

The stress values in the wall exceed that required for concrete cracking

throughout the wall. The wall is observed to suffer significant damage as a

result of the 4 foot drop. Based on the previous results for the Vermont

Yankee pool (drop height of 6 inches) it is clear that the relatively high

drop height has dramatically increased the damage.

7.5 Summary of Observations

The dynamic response of the walls resulting from cask impact loading is more

like an elastic half space than a deep beam. The stresses from impact

propagate through the wall radially from the point of impact prior to the

overall transverse displacement of the wall. As a result, the deformed shape

7-5
of the wall at any given time is a combination of global transverse

displacement (like a shear beam) and deflections caused by propagation. and

reflections of stress waves in the wall. The resulting displaced shapes are

quite complex and do not resemble the simple transversedeflection one would

expect from a shear beam. Because of the dimensions of the walls (i.e., the

walls would be classified as "very deep beams" by the ACI code) and the

observed complex dynamic response, simple classical methods of strength

estimation such as ACI code formulas are inadequate. Iln addition, estimation

of cracked sections for more rigorous analysis is essentially impossible.

The finite element analysis which was performed had the desirable feature of

rigorously modeling the impact between the cask and wall with the NIKE2D

impact algorithm, however a number of simplifying assumptions had to be made

to conduct the analysis ;in a timely fashion. The most significant

simplifications were the assumptions of homogeneous, isotropic linear elastic

material behavior for both the cask and the wall. Obviously as the walls

incur damage the material does not remain elastic. It is also noted that the

assumption of a rigid cask results in higher stresses than if the cask

flexibility is included. The analysis included only the planar load due to

impact and did not address added the load due to thermal, wall gravity or pool

water loads. Based on the stress distributions from the impact analyses, it

appears that hydrostatic loading would have the most interaction with impact

loads. A simulation of hydrostatic and impact load combination would require

a three-dimensional modeling effort.

The results of the idealized analyses indicate that pool walls similar to

those of both the Vermont Yankee and Robinson plants would suffer severe

damage as a result of the worst-case cask drops. The indicated regions of

potential reinforcing steel yield are quite extensive. While the integrity of

the pool liner is difficult to predict, it seems likely that the liners could

be severely damaged (tearing, cracking of welds, and pull-out of anchors).

Based on these results, loss of pool water certainly cannot be ruled out.

Even with the most advanced nonlinear finite element techniques, it would be

difficult to obtain a very high level of confidence in the complete accuracy

of the impact analysis of the wall. A definitive estimate of wall damage

would require full scale drop tests on a wall. However, the simple linear

analysis which has been performed gives insight into both the level of damage,7

and for the Vermont Yankee plant, the relative damage between the 110 ton and

40 ton casks.

7-6
.110 TON CASK

.......... .. ....

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Figure 7-1: Finite Element Model of Wall and Cask

7-7
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7-8
Figure 7-3: Bi-Axial Failure Envelope for Concrete

:::::INDICATES REGION

:":::OF CRACKED

CONCRETE

CUMULATIVE DAMAGE...

ESTIMATED REGION OF

TENSION CRACKS IN

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7-9
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7-10
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7-11
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Figure 7-8: Yielding Of Vertical Steel

7-12
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7-13
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7-14
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Figure 7-11: Yielding of Horizontal Steel

7-15
SI marBB

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FINITE ELEMENT MODEL OF WALL AND CANISTER

Figure 7-12: Finite Element Model of the Cask and Wall System

7-16
ROBINSON CASK DROP 1 1MAR88 I I

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7-17
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7-18
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Figure 7-15: Estimating Damage to Horizontal Reinforcing Steel

7-19

8. SUMMARY AND CONCLUSIONS

8.1 Summary

The objective of this study was to provide technical assistance to the NRC

concerning the Generic Issue - 82, "Beyond Design Basis Accidents in Spent

Fuel Pools". The realistic seismic capacity of spent fuel pools at two

representative plants was estimated and the spent fuel pool's seismic failure

probability was assessed and site specific information was used for this

analysis. For this purpose, the spent fuel pools at Vermont Yankee (BWR) and

H. B. Robinson Unit 2 (PWR) were studied. The seismic PRA and margins

approach was used in this study to assess the seismic risks from spent fuel

pool failures.

The failure modes for the spent fuel pools were: 1) loss of liner integrity

precipitated by gross structural failure of the spent fuel pool, 2) loss of

function of the fuel pool support system (e.g., pool cooling and make-up water

capacity) resulting in loss of water inventory, 3) damage to fuel racks and

4) cask drop accidents. Gross structural failure of the pool was analyzed by

identifying the potential locations for gross failure. In the Vermont Yankee

spent fuel pool, the out-of-plane shear failure of the pool slab was

determined to be the controlling failure mode. In the Robinson spent fuel

pool, the controlling failure mode was identified to be the out-of-plane

bending of the pool wall. The fragility of the pool structures under these

failure modes was estimated taking into account the seismic margins inherent

in design along with the uncertainties. The so-called High Confidence of Low

Probability of Failure capacity of the pool structure from seismic motion was

estimated to be 0.50g for Vermont Yankee and 0.65g for H. B. Robinson Unit

2. The higher capacity for the Robinson fuel pool may have resulted from

structure and fuel rack modifications included with installation of the high

density fuel racks.

The cooling and make-up water systems for the two pools were reviewed; event

and fault trees were constructed to identify the accident sequences that

result from failure of these systems. For the components appearing in these

accident sequences, seismic fragilities were estimated based on the design

information and plant walkdowns. Boolean equations developed from the fault

tree analysis were quantified using the seismic fragilities and hazard

curves. The dominant components to spent fuel pool systems failure are

similar to component that have been found to contribute to seismic risk in

several PRAs; poorly anchored electrical equipment and tanks. The

consequences of the cooling and make-.up water system failure on the spent fuel

pool system were assessed by performing a thermal analysis. It was concluded

that the fuel assembly uncovery would occur only after 3 to 7 days from the

time of failure of the cooling and make-up water systems; this response time

is considered to be sufficiently long for any recovery action.

The free-standing high density fuel racks at Vermont Yankee were reviewed for

potential failure under seismic loads. It was concluded that crushing of fuel

and assemblies and the flexural failure of the rack are not credible failure

8-1
modes. Therefore, the fuel assemblies cannot be compressed into a critical

mass; no serious effects can result as long as the fuel is covered.

In this study, preliminary site-specific seismic hazard. curves developed by

LLNL for the NRC under the Eastern United States Seismicity Program were used.

The mean annual frequency of seismic failure for the spent fuel pool structure

was estimated to be 6.7E-06 for Vermont Yankee and 1.8E-06 for H. B. Robinson

Unit 2. The 5% to 95% confidence range for these frequencies was about 3.0E11 to 2.4E-05. The mean annual frequency of failure of cooling and make-up

water systems in the two spent fuel pool systems was about 1.5E-04/reactoryr. As explained before, the consequence of this failure is not significant

because of the long response time available to recover cooling or provide

alternate makeup.

The effect of a spent fuel cask drop on spent fuel pools was studied by

performing a nonlinear two-dimensional finite-element analysis of the impact

problem using Vermont Yankee and HBR as representative designs. For Vermont

Yankee, two cask drops were considered: 40 ton and 110 ton both dropping from

6 inches. For Robinson, a 68 ton cask dropping from 4 feet was considered.

The results of the analyses indicate that the pool walls would suffer severe

damage as a result of cask drop. The indicated regions of potential

reinforcing steel yield are quite extensive and while the integrity of the

pool liner in difficult to predict, it seems likely that the liner would be

severely damaged. Based on these results, loss of pool water certainly cannot

be ruled out. The probability of this event is not estimated in this study;

it is a function of the rate of spent fuel shipping, amount of storage

available in the plant, handling equipment, reliability and human error in the

handling of the cask. In the BNL report, the estimated cask drop probability

was less than E-8 per reactor-yr which is consistent with the implementation

of A-36 requirements.

8.2 Conclusions

Based on the insights and results obtained in this study the following

conclusions can be drawn:

1. The spent fuel pool structure which is designed to retain large amounts of

water and to withstand gravity and lateral loads from the fuel racks has a relatively high seismic capacity (the HCLPF capacity of the pool structure

is estimated to be more than 3 times the SSE value); this is true for both

the BWR spent fuel pool which is mounted high in the reactor building and

the PWR pool which is generally at the ground level. Therefore, seismic

risk contribution from spent fuel pool structural failures is negligibly

small.

2. The assessment of potential failure modes of the fuel pool racks and fuel

assemblies has indicated that the fuel rack design is such that the

assembly cannot be compressed into a critical mass thereby leading to a

severe accident.

8-2
3. Parts of the cooling and makeup systems for spent fuel pools are not

designed as seismic class 1 systems and as such their failure is expected

at relatively low seismic levels. However, the failure of cooling and

makeup systems would not uncover the spent fuel assemblies for about 3 to

7 days; it is expected that some recovery action could be taken in this

time period.

4. This study has shown that the use of plant-specific and site-specific

information would reduce the uncertainty bands on the seismic risk

estimates; previous BNL study using generic information had estimated the

annual frequency of gross structural failure of the pool to vary from

1.6E-10 to 2.6E-O4 for PWR; the current study shows this range to be

3.1E-11 to 1.4E-05. Similar results are obtained for the BWR spent fuel

pool. Note that the reduction in the uncertainty on the higher end is

more significant. The range of uncertainty in the risk estimate is

governed by the range of uncertainty in the seismic hazard at the site.

Note that the hazard curves used in this study are preliminary and subject

to change following NRC and peer reviews. It is expected that the

uncertainty range in'hazard curves will be reduced, thereby reducing the

uncertainty in the'risk estimates.

5. This study used two-representative spent fuel pools- a.PWR and a BWR.

These pools have been designed to the seismic design criteria existing in

the late 1960s. Their large seismic capacities lead us to conclude that

,the pools designed to current seismic standards should have higher seismic

capacities and should not contribute significantly to seismic risk.

8-3

9. REFERENCES

American Concrete Institute, Building Code Requirements for Reinforced

Concrete, ACI 318-77, Detroit, Michigan, 1977.

American Nuclear Society, "Design Objectives for Light Water Reactor Spent

Fuel Storage Facilities at Nuclear Power Stations", ANSI N210-1976/ANS-57.2,

April 1976.

American Society of Civil Engineers, Design of Structures to Resist Nuclear

Weapons Effects", Manual and Report on Engineering Practice No. 42, 1985.

Bernreuter, D. L., Savy, J. B., and Mensing, R. W., UCID-20421, Seismic

Characterization of the Eastern United States, Volume 1: Methodology and

Results for Ten Sites, April 1985.

Bernreuter, D. L., Savy, J. B., and Mensing, R. W., NUREG/CR-4885, UCID-20696,

Seismic Hazard Characterization of the Eastern United Stated: Comparative

Evaluation of the LLNL and EPRI Studies, May 1987.

Biggs, J.M., Introduction to Structural Dynamics, McGraw-Hill Book Company,

1964.

Chow, L., Conway, H., and Winter, G., Stresses in Deep Beams, Proceedings of

the American Society of Civil Engineers No. 127, May 1952.

Davito, A., Murray, R., Nelson, T., and Bernreuter, D., Seismic Analysis of

High Level Neutralized Liquid Waste Tanks at the Western New York State

Nuclear Service Center West Valley New York, Lawrence Livermore National

Laboratory, UCRL-52485, 1978.

Dong, R.G., "Effective Mass and Damping of Submerged Structures", UCRL-52342,

Lawrence Livermore National Laboratory, April 1978.

EPRI, 1987a, "Seismic Verification of Nuclear Plant Equipment Anchorage", EPRI

NP-5228, Prepared by URS Corp., May.

EPRI, 1987b, "Generic Seismic Ruggedness, of Power Plant Equipment", EPRI NP5223, Prepared by ANCO Engineers for EPRI, May.

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9-1
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9-2
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9-3
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9-4
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9-5

Appendix A

Vermont Yankee Nuclear Power Station

Spent Fuel Pool

Thermal Analysis

Mass of water In fuel pool

Fuel pool volume

26'0" x 40'0" x 39'3/4"

260" 40'0"

39'3/4"

39.063'

26' x40' x39.063' = 40625 It33

FSAR pg. 10.3-3

Sec. 10.3.4

1 gal

.13368 ft 3

0~

303,897.4 gal

From FSAR pg 10.5-7, Table 10.5-1

Vol = 41,600 ft 3

Use 40625 ft 3

Specific volume of water @ 1 atm pressure, T = 150OF (Tech Spec limit) from

Keenan, Keyes, Hill & Moore Steam Tables

-of= 0.016343 ft 3

T= 150'F

ft 3

= 0.016716

T= 212'F

Ug = 96.99 It3'g = 26.80

V 40625 ft 3

•f ft 3

0.016343

lbm

= 2,485,774 Ibm

A-1
Time required-to heat fuel pool water to boiling temperature

Failure state 1

Q = 6.54 x 106 Btu heat Input from"FSAR, pg 10.5-8, Table 10.5.2 hr

Energy Increase = heat Input - heat loss

Energy Increase = ( c AT) -m = mass of fuel pool water, Ibm

t

Cp= specific heat of water

Btu

Ibm OF

AT = temperature rise

= (212-F - 150-F) = 62oF

t = time for temperature rise

Heat loss = heat conduction through walls + heat convection from fuel pool

and floor of fuel pool water surface

Assume = 0 =m h ATair

(mCp AT) =-

m Cp -Q m h ATair ,for 1st cut calculation assume heat convective loss =0

2,485,774 Ibm ( 1 Btu 620F

- mr Cp AT ibm OF) 6tBt

6.54 X 106 Btu

hr

= 23.57 hrs

A-2
Rate of Water elevation drop due to boll-off of fuel, pool

Failure state 1

0= m hfg where m = water mass loss rate

hfg = heat of vaporization

S970.3 Btu water, T -- 212 0F, P I atm, from Keenan, Keyes, Ibm ... Hill iSand Moore Steam Tables,,

m_-

hfg

= 6.56 x 106 Btu 6760.8 ibm

970.3 Btu hr hr

Ibm

=m g= 6760.8. 0.016716 f

= 3 ft31 

rhr

From Keenan, Keyes, Hill and Moore Steam Tables

ft3

0.016716 1 T- T 212'F

H V

W 113 it" (26.40) ft 2

h r

= 0.109 fthr

= 1.3 in.

hr

A-3
Time required to heat fuel pool to boiling temperature

Failure state 2

6 Btu Q = 15.56 x 10 hr heat input from FSAR, pg 10.5-9, Table 10.5.3

2,485,774 Ibm 1 Btu 62 0F

t= m Cp AT 2Ibm OF)

0 15.56 x 106 Btu

hr

= 9.91 hrs

A-4
Rate of water elevation drop due to boll-off of fuel pool

Failure state 2

0 m hfg

0 Q 15.56 x 106 Btu 1 .Ibm

Btu hr hr hfg 970.3 ýt

V =m 1 gIbm ft316036.3 -h- 0.016716 ftbm

V H A

= 268.1. ft3

hr

f3

= 268.1

hr

= 0.258 hr hr

= 3.09 I n.

hr

1 (26 x 40) ft 2

A-5

Appendix B

Vermont Yankee Nuclear Power Station

Spent Fuel Pool

Fault Trees

,c
w
w

(A)
7)
w•
VYNPS

wX0)
"4
ca
w
w-A0
Pump

sws anchor

fails

Appendix C

H.B. Robinson Unit 2

Spent Fuel Pool

Fault Trees

HBR2

9-L
HBR2

9

r:3
9
0
HBR2

9
9
,.) -J

18-O7) U.&U NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER IA$sgr dby PPM5: LPS, add Vol. No., ifany)

NRCM 1102,

3201,.3202 BIBLIOGRAPHIC DATA SHEET NUREG/CR-5176

SEE INSTRUCTIONS ON THE REVERSE UCID-21425

2. TITLE AND SUBTITLE 3. LEAVE BLANK

Seismic Failure and Cask Drop Analyses of the Spent Fuel

Pools at Two Representative Nuclear Power Plants 4. DATE REPORT COMPLETED

June"°' I 198'AR

B.oVyT 11 .06 P r a s s i n o s , C . Y . K im u r a , D . B . M cC a l l e n , J. DAT E RE OR I U

R. C Murray, LLNL MON. DATE REPORTSSUED

M K. " Ravindra, R. D. Campbell, P. S. Hashimoto, January 1989

A. M. Nafday. W. H. Tong. EQe Engineering 7. PERFORMING ORGANIZAl ION NAME AND MAILING ADDRESS (Include Z•p CodeI B. PROJECT/TASK/WORK UNIT NUMBER

Lawrence Livermore Subcontractor:

National Laboratory EQE Engineering, Inc. 9 FIN OR GRANT NUMBER

7000 East Avenue 3150 Bristol, Suite 350

Livermore, CA 94550 Costa Mesa, CA 92626 FIN A-0814

10. SPONSORING ORGANIZATION NAME AND MAILING ADDRE (In.lu~eZ~ p Cod.) 11a. TYPE OF REPORT

Division of Safety Issue ResolAui on

Office of Nuclear Regulatory Research Technical

U. S. Nuclear Regulatory Commission b. PERIOD COVERED I'Mclusivedes)

Washington, D. C. 20555

12 SUPPLEMENTARY NOTES

13. ABSTRAICT f)M•wr• . fo, .

il ABSTRACT (200 words or ress]

This report discusses work done in support of the resolution of Generic Issue-82,

"Beyond Design Basis Accidents in Spent Fuel Pools". Specifically the probability

of spent fuel pool failure due to earthquakes was determined for the pools at the

Vermont Yankee Nuclear Power Station (BWR) and the H. B. Robinson S. E. Plant,

Unit 2 (PWR). The dominant failure mode for each pool was gross structural failure

caused by seismic motion resulting in the loss of pool liner integrity. The resulting

sudden loss of water was then assumed to lead to a self-propagating cladding failure

and fission product inventory release from the spent fuel elements in the pool. The

mean annual frequency of failure due to this failure mode was found to be 6.7E-06 at

Vermont Yankee and 1.8E-06 at H. B. Robinson. Other earthquake induced failure modes

studied but found to be less important were loss of pool cooling and make-up

capability, fuel rack'damage and loss of liner integrity due to a cask drop accident.

14 DOCUMENT ANALYSIS - .. KEYWORDS/DESCRIPTORS

Spent Fuel Pools, Generic Safety Issue, Probabilistic Risk Assessment,

PWRs, BWRs Seismic Hazard, Fragility, Vermont Yankee Nuclear Power

Station, H. B. Robinson S. E. Plant

b. IDENTIFIERS/OPEN.ENDED TERMS

15. AVAILABILITY

STATEMENT

Unl imited

16. SECURITY CLASSIFICATION

(Thi. page)

Unclassified

IThss report)

Unclassified

17. NUMBER OF PAGES

18. PRICE

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