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Containment Performance Working Group Report.Draft Report for Comment
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Issue date:	05/31/1985
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	NUREG-1037, NUREG-1037-DRFT, NUREG-1037-DRFT-FC, NUDOCS 8506140588
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NUREG-1037 Containment Performance Working Group Report Draft Report for Comment

' U.S. Nuclear Regulatory Commission Offics of Nuclear Reactor Regulation p*"%q l

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I m

NUREG-1037 l

l Containment Performance Working Group Report Draft Report for Comment Manuscript Completed: May 1985 D:ta Published: May 1985 Divi; ion of Engineering Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission W2:hington, D.C. 20555 ig/ a

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ABSTRACT

' Containment buildings for power reactors have been studied to estimate their leak rate as a function of increasing internal pressure and temperature associ-ated with severe accident sequences involving significant core damage. Poten-tial leak paths through containment penetration assemblies (such as equipment hatches, airlocks', purge and vent valves, and electrical penetrations) have been identified and their contributions to leak area for the containment are incorporated into containment response analyses of selected severe accident sequences to predict the containment leak rate and pressure / temperature response as a function of time.

Becruse of lack of reliable experimental data on the leakage behavior of containment penetrations. and isolation barriers at pressures beyond their design conditions, an analytical approach has been used to estimate the leakage behavior of components found in specific reference plants that approximately characterize the various containment types.

NUREG-1037 iii

PREFACE The Reactor Safety Study as well as other industry studies have assumed the con-tainment to have failed after the containment internal pressure reaches its threshold capability value. The Senior Review Group (SRG) for the Severe Acci-dent Research Plan recognized that an improved model for containment behavior under important severe accident sequences was needed, and established the Con-tainment Performance Working Group (CPWG).

The CPWG was set up under the direction of Mr. Vincent S. Noonan, Office of Nuclear Reactor Regulation (NRR), to develop a realistic containment behavior model that takes into account the response of varicus containment openings as the containment structure deforms, as a result of internal pressure, and as elastomeric seals for various penetration closures are subjected to accident temperature.

This report is made up of two parts: (1) the main report and (2) appendices supporting information cited in the main report. The main report presents the CPWG findings with respect to leak area models and containment performance for the containment types studied. Supporting data included in the appendices are grouped into three major areas as follows: Appendix A deals with the effect of nonuniform thermal environment on degradation of containment penetrations during severe accidents; Appendix B deals with leak area estimates for power reactor containments during severe accident conditions; and Appendix C deals with prob-able leak areas in containment electrical penetration assemblies.

In addition to the staff of the Nuclear Regulatory Commission, Brookhaven National Laboratory (BNL), Idaho National Engineering Laboratory (INEL) and Sandia National Laboratories (SNL) are major participants in this study.

Dr. Krzysztof Parczewski of the Division of Engineering (NRR) studied the effect of nonuniform thermal loads on containment penetration seals for purge and vent valves and personnel airlocks. The results of his study are presented here as Appendix A.

The INEL staff under the direction of Mr. T. L. Bridges visited the six refer-ence plants and gathered the details of each containment structure and the var-ious penetrations. The INEL staff (EG&G) then analyzed each containment and 1

prepared estimates for leak areas. INEL staff findings are presented here as Appendix B.

The behavior of electrical penetration assemblies was studies by Dr. W. A. Sebrell of SNL and his work appears as Appendix C.

The behavior of containment purge and vent valves was studied by Mr. B. E.

Miller and Dr. A. Agrawal of BNL. Dr. L. Greimann of Ames National Laboratory performed a nonlinear finite element analysis of the Sequoyah equipment hatch penetration to determine its potential for leakage at high internal pressure.

Various leak area estimates were then grouped and integrated into pressure- and temperature-induced leak areas by Dr. C. H. Hofmayer of BNL. Dr. W. T. Pratt of BNL evaluated containment response incorporating the estimated containment leak NUREG-1037 y

area.

The technical direction for the containment response was provided by Mr. J. W. Shapaker and Mr. R. Palla of the Division of the Systems Integration

'(NRR). Valuable suggestions regarding BWR Mark II and Mark III containments were provided by Dr. J. E. Rosenthal of the Division of Systems Integration (NRR). Technical direction of the overall containment performance was provided by Mr. G. Bagchi of t'le Division of Engineering (NRR).

Under the direction of Dr. Themis P. Speis, Director, Division of Safety Tech-nology (NRR), the Containment Loads Working Group (CLWG) worked on the defini-tion of containment loading for important severe accident sequences. The pro-cess of estimating containment pressure / temperature response and its influence on the leak area and then modifying the leak area based on the latest under-standing of the CLWG has been an iterative process. The high degree of coopera-tion and personal commitment by Drs. Hofmayer and Pratt of BNL made it possible for the CPWG to develop an integrated containment performance model and to complete this report. Special thanks to Gladys Ordaz for typing this report.

Major findings of this report are summarized in the executive summary. The main purpose of this report is to present a first approximation of a leakage model for'six containment types (PWR large dry, subatmospheric, ice condenser and BWR Mark I, II, and III). It should be noted that the results presented here are based on engineering judgment and, therefore, are subject to large uncertainties.

However, as more data become available, containment behavior models can be supplemented by such data. The findings of this report are being presented for peer review and comment by interested parties. Comments should be addressed to:

Mr. Goutam Bagchi Division of Engineering U.S. Nuclear Regulatory Commission Washington, D.C. 20555 Signed by Richard H. Vollmer, Deputy Director ffice of I pection and Enforcement i

' /

. fA ky A. Arlotto, Director

[ Dtvision of Engineering Technology Office of Nuclear Regulatory Research Ja P.KYght Acting Director D ion of Engineering ;

O ce of Nuclear Reactor Regulation

{

i NUREG-1037 vi u _

CONTENTS Pages ABSTRACT.............................................................. iii PREFACE............................................................... v

. EXECUTIVE

SUMMARY

..................................................... xi 1 INTRODUCTION..................................................... 1-1 2 OVERVIEW OF CONTAINMENT STUDIES.................................. 2-1 2.1 Approach.................................................... 2-1

2. 2 Accident Sequences Considered............................... 2-2 2.2.1 PWR Large Dry Containment............................ 2-3 2.2.2 PWR Subatmospheric Containment....................... 2-5 2.2.3 PWR Ice Condenser Containment........................ 2-7 2.2.4 BWR Mark I Containment............................... 2-7 2.2.5 BWR Mark II Containment.............................. 2-8 2.2.6 BWR Mark III Containment............................. 2-9 2.3 Leak Areas.................................................. 2-10 2.3.1 Capability Pressure.................................. 2-11 2.3.2 Pre-Existing Leakage................................. 2-11 2.3.3 Large Opening Penetrations........................... 2-12 2.3.4 Purge and Vent System Isolation Valves................ 2-13 2.3.5 Piping Penetrations.................................. 2-14 2.3.6 Electrical Penetration Assemblies. . . . . . . . . . . . . . . . . . . . 2-15 2.3.7 Non-Uniform Thermal Effects.......................... 2-15 2.4 Containment Response........................................ 2-16 2.4.1 Containment Codes.................................... 2-16 2.4.2 Leakage Models....................................... 2-17 3 RESULTS.......................................................... 3-1 3.1 Leak Area Estimates......................................... 3-1 3.1.1 PWR Large Dry Containment............................ 3-1 3.1.2 PWR Subatmospheric Containment....................... 3-2 3.1.3 PWR Ice Condenser Containment........................ 3-5 3.1.4 BWR Mark I Containment............................... 3-7 3.1.5 BWR Mark II Containment.............................. 3-10 3.1.6 BWR Mark III Containment............................. 3-13 3.2 Impact of Leakage on Containment Response................... 3-15 3.2.1 PWR Large Dry Containment............................ 3-15 3.2.2 PWR Subatmospheric Containment....................... 3-16 3.2.3 FWR Ice Condenser Containment........................ 3-17 3.2.4 BWR Mark I Containment............................... 3-17 NUREG-1037 vii I --

CONTENTS (Continued)

P.89*

3.2.5 BWR Mark II 3.2.6 BWR Mark-III Containment.............................. 3-18 Containment............................. 3-18 3.3 Impact on Radiological Consequences ........................ 3-19 4

CONCLUSIONS ..................................................... 4-1 5'

REFERENCES....................................................... 5-1 APPENDIX A - EFFECT OF NON-UNIFORM THERMAL ENVIRONMENT ON DEGRAD CONTAINMENT PENETRATIONS'DURING SEVERE ACCIDENTS

-APPENDIX B - LEAK AREA ESTIMATES FOR POWER REACTOR CONTAINME ACCIDENT CONDITIONS

> APPENDIX C - PROBABLE LEAK AREAS IN CONTAINMENT ELECTRICAL PENETR ASSEMBLIES FOR SEVERE ACCIDENT ~ STUDIES FIGURES

-Figure Title Pg 1.1 Organization of the Containment Performance Working Group.... 1-3

2.1 PWR Large Dry- Containment Configuration. . . . . . . . . . . . . . . . . . . . . .

2.2 2-18 2.3 PWR Large Dry. Containment Response Without Leakage........... 2-19 2.4 PWR Subatmospheric Containment Configuration................. 2-20 2.5 PWR Subatmospheric Containment Response Without Leakage...... 2-21 PWR Ice Condenser Containment Configuration.................. 2-22 2.6. PWR Ice Ccndenser Containment Response Without Leaka 2.7 2-23 BWR Mark I Containment Configuration................ge....... 2-24 2.8 BWR Mark I Drywell Response Without Leakage. . . . . . . . . . . . . . . . . .

2.9 2-25 2.10-BWR Mark II Containment Configuration........................ 2-26 2.11 Variation in the Mark II Pedestal Configuration.............. 2-27 BWR Mark.II Drywell Response Without Leakage.................

2-28 2.12 BWR Mark III Containment Configuration.......................

2.13 BWR Mark III Drywell Response Without 2-29 Leakage................ 2-30 12.14 Seal Life as a Function of Time at Temperature...............

2.15 2-31 Criteria for Leak Area Estimates Based on Seal Degradation. . . 2-32 13 . 1

'PWR Large D ry Containment Response. . . . . . . . . . . . . . . . . . . . . . . . . . . 3-22 3.2 PWR Large Dry Containment Leak Rate Estimates. . .

3.3 3-23 BWR Ma rk I D rywe l l Re spo ns e . . . . . . . . . . . . . . . . . . . . . .............

3.4 BWR Mark I Leak Rate Estimates........................... 3-24 3.5 3-25 BWR Mark II Containment Response.............................

3.6 BWR Mark II Leak Rate Estimates..............................

.... 3-26 3-27 I

l i

NUREG-1037 viii

CONTENTS (CONTINUED)

TABLES Table Title Pg Summary of Leak Area Estimates............ .................... xiv 1

2 Containment Design Parameters.................................. xvi xvii 3- Summary of Containment Performance.............................

3.1 Pressure-Dependent Leak Area Estimate for PWR Large Dry Containment........................................... .... 3-3 3.2 Pressure-Dependent Leak Area Estimate for a Subatmospheric Containment....................................................

3-5 3.3 Pressure-Dependent Leak Area Estimate for a PWR Ice Condenser Containment....................................................

3-7 3.4 Pressure-Dependent Leak Area Estimate for a BWR Mark I .

~ Containment....................................................

3-10 3.5' Temperature-Dependent Leak Area Estimate for a BWR Mark I Containment....................................................

3-10 3.6 Pressure-Dependent Leak Area Estimate for a BWR Mark II Containment....................................................

3-12 l

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$ EXECUTIVE SUPMARY At.the-request of the. Severe Accident Research Plan (SARP) Senior Review Group, a Containment Performance Working Group (CPWG) was established to develop containment leakage models for use in severe accident source term work. These

~ leakage models quantify leakage areas as a function of containment pressure and temperature loading for specific accident sequences for various containment types. The leakage models have been incorporated into existing containment computer codes to permit a more realistic assessment of containment behavior f.or severe accidents, specifically, the consideration of containment. leakage as a function of time and the impact of containment pressure relief (due to leak-age) on the mode and timing of containment failure. This information can be used as input to a radiological consequence analysis.

The purpose of this report is to present a first approximation of a leakage model for six containment types (PWR large dry, subatmospheric, ice condenser and BWR Mark I, II, and III). It should be noted that the results presented

here are based on engineering judgment and, therefore, are subject.to large 4

uncertainties. <However, several.research programs are under way that should provide additional bases to supplement and modify the engineering judgment j as appropriate.

Numerous studies have reported containment shell capability pressures that have been utilized as the threshold pressure in-severe accident risk estimates.

s

These capability pressures generally correspond to the point at which the con-

.tainment first reaches an initial general yield state (or 1% tendon strain in a prestressed concrete containment). This study focused on identifying potential leakage paths that may occur before such capability pressures are reached. Con-sequently, it does not consider potential leak paths that may result from large containment deformations. For certain containments, no significant leak paths have been. identified up to the capability pressures considered in this report.

Before further studies are conducted to identify leak paths that may: result t' from higher containment strain conditions, the impact of these findings on If the the overall severe accident risk assessment needs to be assessed.

capability pressure (as defined above) is reached sufficiently late in the acci-dent sequence, the offsite consequences may not be appreciably different if the containment leaks or fails in a gross manner.

The approach taken in conducting this study involves a detailed review of con-tainment penetration designs and an analytical treatment of penetration perform-j ance to estimate leak area as a function of predicted pressure and temperature conditions. The leak area models are then used for selected accident sequences

-to calculate the containment response and associated leak rate. The containment loads considered in this study are based on the findings of the Containment Loads Working Group (CLWG).

)- , Reference containment designs of each type that were selected for this study are:

1 l

(1) PWR,.large dry - Zion (2) -PWR, subatmospheric - Surry I- , 'NUREG-1037 xi l

k (3) PWR, ice condenser - Sequoyah (4) ,BWR, Mark I - Peach Bottom / Browns Ferry (5) BWR, Mark II - Limerick ;

(6) BWR, Mark III - Grand Gulf Leak areas in this report are presented primarily as a function of pressure. A

' preliminary model to treat leak area as a function of temperature is also pre-i sented; however, it is recognized that this approach is very crude and needs more supporting test data before it can be fully utilized. A more refined leak-age model should treat leakage as a combined function of pressure and tempera-ture, as well as of other parameters such as aging and radiation. Although the leakage models are primarily presented as a function of pressure, they qualita- ,

tively recognize that the associated leakage will occur only if seal degradation i (from high temperatures, aging, or other effects) has occurred.

I i

Although specific plants have been selected for this study, it should be realized that many containments have similar details, even between different containment types. For example, the purge system valves or personnel and equipment hatches of one specific manufacturer can be found on both PWR and BWR plants.

Therefore, the evaluations of the six containments in this study probably have identified most unique features that could have a significant effect on containment leakage. However, further investigations are necessary to determine if such features are applicable to or could significantly affect the conclusions concerning the behavior of each containment type. Furthermore, a sensitivity study of the potential variation of containment leakage within the family of each containment type should be performed to better establish the uncertainty bounds.

For each containment type the major penetrations having the greatest potential' for leakage were identified and evaluated. These include:

j -

Large Opening Penetrations

- equipment hatch

- personnel airlock

-- drywell head (BWR)

-- fuel transfer tube

-- CRD removal hatch Purge and Vent System Isolation Valves Piping Penetrations Electrical Penetration Assemblies '

Originally it was intended that three leak area estimates (high, medium, low) as a function of containment pressure would be determined for each plant. The i

high leak area was to be one which has a small probability of being exceeded.

The medium area was to be a best estimate value. The low leak area was to be one which has a high probability of being exceeded. For penetrations with

elastomer-type seals, the leak areas were to be determined by first determining '

the structural response of the seal joints due to pressure loading and then using seal behavior assumptions based on engineering judgment or test results

to obtain the three desired leak areas. Well into the program a small-scale

NUREG-1037 xii

series of seal behavior tests was conducted to try to determine seal behavior A double 0-ring characteristics under simulated severe accident conditions.

configuration with neoprene seals and a double tongue-and groove configuration with silicon rubber seals were tested. The results of these tests suggest that seal leakage for flange separations below 0.030 in, will be nearly insignificant.

Because of the limited amount of testing performed, it was concluded that addi-

'tional testing must be performed before attempting to characterize the behavior of seal materials during severe accident conditions. Sandia National Laboratory is currently undertaking such a program for the Office of Nuclear Regulatory Research of.the NRC. Therefore, rather than determine three containment leak area estimates as originally planned, it was decided that only one leak area estimate would be determined for each containment type based on calculated structural response (flange separation or separation of sealed surfaces) of the penetration realizing that this is an upper bound value. In addition to the assumption that the seals have no resilience, it was also assumed that the per-sonnel airlocks would have only one door available to resist the containment pressure and furthermore, the thermal effects of the pressure unseating closures were neglected. Both of these assumptions are consistent with the goal of determining upper bound leak area estimates.

A summary of leak area estimates is presented in Table 1. As noted above, the upper bound leak area estimates are limited to the containment shell-capability pressures. These pressures and other containment design parameters are listed for each plant in Table 2. The CPWG believes that it is not likely, for the severe accident conditions described herein, that the leak area estimates for the specific plants studied in this report will be exceeded. However, smaller leak areas may also be justified. In the following paragraphs the potential effects of the predicted upper bound leak area estimates on containment perfor-mance and offsite consequences are briefly described. This information is summarized in Table 3.

The leakage estimated for the PWR large dry containment could delay or possibly prevent the containment from reaching its capability pressure. However, the leakage will not occur until relatively late in the accident sequence so that it is expected to have a minor influence on offsite consequences (refer to Table 3).

The leakage areas estimated for the PWR subatmospheric and ice condenser con-tainments were found to be relatively small and did not affect containment.

response; they were also expected to have little impact on offsite consequences.

The leakage estimated for the BWR Mark I and II containments does alter contain-

' ment response and it has both positive and negative effects (refer to Table 3).

However, changes to the BWR operating procedures involve the use of wetwell venting. If the wetwell is vented, then there will be no driving force to pro-duce significant drywell leakage (even with high drywell temperatures), so that-uncertainties associated with drywell integrity will become less important as all of the fission products will be directed through the suppression pool.

For BWR Mark III containments it is concluded that the drywell personnel airlocks will maintain their integrity even in the presence of diffusion flames in the wetwell. Because suppression pool bypass early in the accident is important, tests should be conducted to confirm this finding. The drywell personnel airlock is predicted to result in significant bypass leakageHowever, from high temperatures in the drywell during core / concrete interactions.

NUREG-1037 xiii

e Table 1 -Summary of leak area estimates

=

55 Leak source before s

$5 reaching capability Upper bound-gg Plant / type pressure leak area Comments Zion /PWR large Personnel airlock 5 in.2 at 134 psig dry Initial yield of airlock door frame at 107 psig No credit for redundant airlock doors Purge valve leakage eliminated based on tests of valves at 400*F for 11 hours1.273148e-4 days 0.00306 hours 1.818783e-5 weeks 4.1855e-6 months -

Surry/PWR Airlock barrel subatmospheric flange 0.4 in.2 at 119 psig Purge valve leakage eliminated since CLWG predicts. temperatures less than 300'F Personnel airlock Airlock barrel flange unseats at 81 psig No credit for redundant airlock doors Sequoyah/PWR Personnel airlock

>L

< ice condenser 0.3 in.2 at 50 psig No credit for redundant airlock doors (2) i

Nonlinear analysis of pressure seating hatch indicates no leakage

Peach Bottom / Drywell head 35 in.2 at 117 psig BWR Mark I Drywell head unseats at 27 psig Equipment hatch Equipment hatch unseats at 82 psig Personnel airlock Personnel airlock door frame yields at 94 psig No credit for redundant airlock doors CLWG predicts temperature ranges of 500 F for 2 hours2.314815e-5 days 5.555556e-4 hours 3.306878e-6 weeks 7.61e-7 months to 700 F for 1 hour1.157407e-5 days 2.777778e-4 hours 1.653439e-6 weeks 3.805e-7 months j If temperature is >500 F for 40 minutes, purge valve leakage should be considered. Maximum leakage of 14 in.2 (18-in.-diameter valves)

Table 1 (Centinued) i z t Ei Leak source before 2]

reaching capability Upper bound Plant / type pressure leak area Comments ;

i J.

53 Drywell head 42 in.2 at 140 psig~ Drywell head unseats at 85 psig

'd Limerick /

BWR Mark II Equipment hatch unseats at.75 psig Equipment hatch CLWG predicts temperatures > 600*F for'l-1/2 hours. One case with limestone concrete temperature >750*F for 40 minutes. Two cases with basalt concrete temperature >650*F '

for 60 to 80 minutes i

See comments Main concern is bypass leakage of drywell Grand Gulf / Drywell personnel i BWR Mark III airlock Drywell leakage estimated equivalent to be i

9 in.2 at a pressure of 4 psid estimated on the basis of observed leakage from an integrated i

x

< 1eak rate test i Approximately 4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months after start of accident, l temperatures in drywell could reach 900*F Drywell airlock seals fail 5 to 6 hours6.944444e-5 days 0.00167 hours 9.920635e-6 weeks 2.283e-6 months after accident resulting in maximum bypass leakage of 135 in.2 !

Standing flames in wetwell predicted not to t affect integrity of drywell and containment ,

personnel airlocks i

t I

Table 2 Containment design parameters Total con-Allowable tainment Plant / type leak rate free volume, Design pres- Capability vol %/ day 103 ft3 sure, psig pressure, psig Zion /PWR large dry- 0.1 2,600 47 134 Surry/PWR subatmospheric 0.1 1,800 45 119 Sequoyah/PWR fce condenser 0.25 1,200 12 50 Peach Bottom /BWR Mark I 0.5 280 62 117*

Limerick /BWR Mark II 0.5 410 55 140 Grand Gulf /BWR Mark III 0. 4 1,670 15 60

CPWG utilized the capability pressure predicted for Browns _ Ferry.

NUREG-1037. xvi

I' ]

Table 3 Summary of containment performance E Impact of leakage on offsite' consequences-A Plant / type Containment performance findings.

? Without leakage the containment would reach its capa- The impact of leakage on offsite consequences g Zion /PWR bility pressure between 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months to several days after~ would be low for sequences in which containment y large dry integrity is threatened by'long-term pressure /

the start of the accident depending on conditions in the reactor cavity.

temperature buildup. This is because significant leakage is predicted only if the reactor cavity is flooded. A flooded cavity. implies a coolable Leakage before the capability pressure would not exceed 5 in.2 Such leakage would be sufficient debris bed and hence limited core / concrete inter-actions, which in turn results in minimum fission to delay or possibly prevent the containment from product release via this mechanism. Hence, as reaching its capability pressure. most of the fission products are released to the containment.during invessel degradation of the Significant leakage could occur at 8 hours9.259259e-5 days 0.00222 hours 1.322751e-5 weeks 3.044e-6 months after the reactor core for this sequence, aerosol agglom-accident, but only in the case of a flooded cavity, eration and settling will reduce the fission pro-Without leakage the containment would not approachi ducts suspended in the containment atmosphere Surry/PWR before the beginning of significant leakage, subatmos - its capability pressure until several days after cheric~ the start of the accident. Other mechanisms that result in early containment x failure, such as' hydrogen burns, isolation failure,

'l etc., have potentially a high impact on offsite consequences but would be unaffected by leakage Leakage before the capability pressure would not ex- induced by the. severe accident conditions ceed 0.4 in.2 Such leakage would have little effect considered in this report.

on containment performance.

Sequoyah/PWR Without leakage the containment would reach its capa-ice con- bility pressure after about 8 hours9.259259e-5 days 0.00222 hours 1.322751e-5 weeks 3.044e-6 months following the denser start of the accident assuming no early failure via hydrogen burns.

Leakage before the capability pressure would not ex-ceed 0.3 in.2 Such leakage would have little effect on containment performance.

z g Table 3 (Continued)

S

,L Plant / type Containment performance findings 8 Impact of leakage on offsite consequences N

Peach Bottom /Without leakage the containment would reach its The leakage in Mark I and II containments can be BWR Mark I capability pressure between 4 and Sh hours af ter the start of the accident. large and has both positive and negative effects.

A negative effect is the potential for signifi-cant release of the fission products earlier than A leakage of approximately 10 in.2 to 12 in.2 would be would have been predicted based on threshold models.

sufficient to prevent the containment from reaching This earlier release when coupled with the' potential its capability pressure. for greater pool bypass (less fission product scrub-Leakage before the capability pressure could be as bing) could result in increased offsite consequences.

high as 35 in.2 However, a positive effect of leakage vs. gross fail-ure is that it could result in significant aerosol Limerick / agglomeration and settling in the reactor aerosol Without leakage the containment could reach its capa-BWR Mark II bility pressure as early as 5 hours5.787037e-5 days 0.00139 hours 8.267196e-6 weeks 1.9025e-6 months but not later than agglomeration In addition, theand settling standby gasin the reactor treatment building.

system could 1 day following the start of the accident. be utilized to scrub the fission products under x these circumstances. The above uncertainties asso-A leakage of approximately 5 in.2 would be sufficient ciated 1 with leakage vs. gross failure could be elimi-to prevent the containment from reaching its capabil- nated by using current operating procedures which-3 ity pressure.

call for wet well venting.

aa bef e the capability pressure could be as

~ The impact of drywell leakage is that it pro-vides a mechanism for fission products to bypass Grand Gulf / Without leakage the containment would reach its capa- the suppression pool. However, significant pool BWR Mark III bility pressure after about 13 hours1.50463e-4 days 0.00361 hours 2.149471e-5 weeks 4.9465e-6 months following the bypass is expected to be late in the accident start of the accident, assuming no early failure via after most of the fission products have already hydrogen burns.

been scrubbed by the pool. Hence, the inpact on the offsite consequences of loss-of-drywell Drvwell leakage equhalent to 9 in.2 at a pressure integrity late in the accident sequence is not of 4 psid was determined on the basis of observed expected to be significant. The leakage esti-leakage from an integrated leak rate test. mate at 4 psid is important because of the poten-tial for suppression pool bypass early in the Drywell leakage sufficiently large to result in pool accident sequence and should therefore be care-bypass could occur 5 to 6 hours6.944444e-5 days 0.00167 hours 9.920635e-6 weeks 2.283e-6 months af% r the start of fully assessed in future studies. However, the the accident.

leakage paths at this pressure would be tortuous, with the potential for significant retention of aerosol fission products.

this occurs late in the accident sequence and is expected to have little impact on offsite consequences.

The preceding results depart significantly from the assessment of containment performance in the Reactor Safety Study (RSS) (NUREG-75/014, formerly WASH-1400) and in subsequent studies based on RSS methods (Reactor Safety Study Methodology Application Program, or RSSMAP) (NUREG/CR-1659). For accidents with total loss of all active containment heat removal (which is the focus of the present CPWG activities), the RSS predicted that the containment pressure would exceed the structural capability pressure very close to the time that the core materials melt through the reactor pressure vessel. The CPWG found that for risk-dominant sequences the accident environment (pressure / temperature) inside the contain-ment does not challenge its integrity in most cases until several hours after the reactor vessel failure. Major reasons for this development are (1) more detailed accident progression modeling following the reactor vessel failure and (2) industrywide studies of containment capability pressures. Containment capa-bility pressures used by the CPWG are presented in Table 2. It should be noted that the CPWG did not perform any analysis to determine the containment capabil-ity pressures but used the information generally available from other studies and Industry Degraded Core Rulemaking Program (IDCOR) work. In addition, the detailed assessment of the potential for containment penetration leakage per-formed by the CPWG goes far beyond anything considered in the RSS. In the RSS, leakage was assumed to have no impact on containment response or offsite conse-quences. However, the present CPWG assessment has demonstrated the importance of containment performance during severe accidents and the potential influence of significant leakage before reaching the structural capability pressure of the containment. NRC-sponsored research on containment capacity and failure modes has already demonstrated through a number of small-scale (1/32) steel models and a large-scale (1/8) steel model that containment function is expected to be available for membrane strains between 2% and 3%. The large-scale model incorporated representative openings such as piping penetrations, equipment hatch, personnel airlock, etc., and was tested with nitrogen at ambient tempera-ture. Results of the large-scale model test show that for membrane strains between 2% and 3% the total observed leak rate remained below 2% of volume per day. The large-scale model experienced a bursting type of failure at approxi-mately 5 times its design pressure. The capability pressure of a containment and its behavior under a severe accident challenge is very much dependent on specific design details, and have such important implications for offsite conse-quences that future studies on severe accident risk must incorporate realistic containment behavior models, especially in the light of recent research findings.

It should be noted that the above comments are qualitative and are intended to indicate potential trends. The leakage estimates must be incorporated into containment failure mo4 and fission product release analyses to determine the quantity and characteristics of the radionuclide release. It should also be noted that the point of release of the radionuclides is important (auxiliary building vs. directly to the atmosphere) and will certainly influence the off-site consequences.

In addition to the preceding findings, this study has led to the following general conclusions:

NUREG-1037 xix

s The potential for containment leakage through penetrations before reaching currently reported containment capability pressures should be considered '

.in severe accident risk estimates.

The potential for significant leakage before reaching currently reported containment capability pressures appears to be greater for BWRs than for PWRs.

-Leakage before reaching containment capability pressures can also occur with PWRs, but such leakage is much more plant' specific.

- Failure of nonmetallic seals for containment penetrations -(primarily pressure unseating equipment hatches, personnel airlocks, drywell heads and purge valves) are the most significant sources of containment leakage.

Although generic studies of containment types are useful in identifying sources of containment leakage, final conclusions may need to be plant specific.

Current efforts rely on analysis and engineering judgment. Additional test data are needed to better quantify the leak tightness of containment penetrations when subjected to severe accident conditions.

On the basis of the results to date, both analytical and experimental studies should continue to better characterize containment leakage before reaching con-tainment capability pressures as defined above. Furthermore, efforts should be made to better define the confidence levels associated with these capability pressures. Further studies should include the following:

Tests to fully assess the behavior of penetration seals under severe accident pressure and temperature conditions, including the effects of aging and radiation.

Sensitivity studies to assess the potential variation of containment leakage within the family of each containment type.

Sensitivity studies to determine the magnitude and timing of containment leakage which can have a significant effect on radiological consequences.

An assessment of the potential for plugging of identified leak paths.

An assessment of the survivability of equipment inside containment during important severe accident sequences.

Identification of leakage paths after release from containment and an' assessment of the effect of holdup of releases to the auxiliary or reactor buildings.

NUREG-1037 xx

-CONTAINMENT PERFORMANCE WORKING GROUP REPORT 1--INTRODUCTION The'.U.S.' Nuclear Regulatory Commission (NRC) is reassessing accident source term evaluation methods and assumptions. The overall-NRC effort will address important severe accident sequences by grouping severe accidents into cate-gories based-on core melt progression and containment performance. It will consider degraded core accidents like the accident at Three Mile Island Unit 2, core melting within the vessel, vessel melt-through, and basemat melt-through.

The containment performance categories will include small containment leakage, large containment leakage, containment bypass, and early and late containment failures. Severe accident sequences that could yield the various categories will be identified on the basis of current knowledge and engineering judgment.

Emphasis will.be placed on identifying a complete set of qualitatively distinct sequences. Special attention will be pai'd to identifying sequences that can potentially lead to very large source terms. The results of this effort will be described in NUREG-0956, "A Reassessment of Accident Source Term Evaluation Methods and Assumptions." The containment studies described in this report are being utilized in support of this larger NRC effort.

Two basic models can be used in severe accident risk estimation to characterize the loss of containment integrity; namely, the " threshold" model and the

" leakage before failure" model. Virtually all risk assessments performed to date have used the' threshold model, which defines a threshold pressure, with some associated encertainty, at which the containment will suffer a loss of holding capability with the potential for significant and rapid release to the environment of the containment atmosphere (which may contain a large amount of fission. products). If the containment pressure loading is calculated to be below the threshold pressure, the containment _is considered to be im,act and

'the.offsite consequences are therefore quite low. Some recent analyses have-

_ pointed out that loss of containment holding capability is not-the only pathway

'to substantial offsite consequences. Significant leakage, on the order of 100 volume percent per day (vol %/ day) or more, may result (if it occurs sufficiently early in the accident sequence) in substantial offsite releases of radioactive

, fission products.' The. leakage-before-failure model provides a means of accounting for this condition when performing risk assessment analyses.

At the request of the Severe Accident Research Plan (SARP) Senior Review Group, a Containment Performance Working Group (CPWG) was established to develop con-tainment leakage models for use in severe accident source term work. These leakage models quantify leakage areas as a function of containment pressure and temperature loading for various containment types. The leakage models have

, been incorporated into existing containment computer codes to permit a more realistic assessment of containment behavior for severe accidents; specifically, the consideration of containment leakage as a function of time and the impact

-of_ containment pressure relief (due to leakage) on the mode and timing of containment failure. This information can be used as input to a radiological conse_quence analysis.

NUREG-1037 1-1

The purpose of this report is to present a first approximation of a leakage model for six containment types (PWR: large dry, subatmospheric, and ice con-denser; and BWR: Mark I, II, and III).

It should be noted that the results pre-sented herein are based on engineering judgment and, therefore, are anticipated to be controversial. However, several research programs are under way that should provide the required bases to supplement and modify the engineering judgment as appropriate.

Numerous studies have reported containment shell capability pressures that have been utilized as the threshold pressure in severe accident risk estimates (TEC, IDCOR Technical Report 10.1, SWECO Report TP-84-13; NRC reports NUREG/CP-0033, NUREG/CR-2442, NUREG/CR-1891, and NUREG/CR-1967). These capability pressures generally correspond to the point where the containment first reaches an ini-tial general yield state (or 1% tendon strain in a prestressed concrete con-tainment). This study focused on identifying potential leakage paths that may occur before such capability pressures are reached. Consequently, it does not consider potential leak paths that may result from large containment deformations.

For certain containments, no significant leak paths have been identified up to the capability pressures considered in this report. Before further studies are conducted to identify leak paths that may result from higher containment strain conditions, the impact of these findings on the overall severe accident risk assessment needs to be evaluated. If the capability pressure (as defined above) is reached sufficiently late in the accident sequence, the offsite consequences may manner.

not be appreciably different if the containment leaks or fails in a gross Initially, this study attempted to provide three leakage area estimates (high, medium, and low) for each containment type. However, because of the lack of definitive test data to fully characterize the leakage behavior of containment penetrations, the CPWG modified its approach and focused on identifying upper bound leak area estimates for each containment type. The CPWG believes that for the severe accident conditions described here, the leak area estimates in this report are not likely to be exceeded. On the other hand, smaller leak areas may also be justified. The results presented should be utilized to assess the impact of containment leakage on the radiological consequences of an acci-dent. However, until more test data are available, these results should be coupled with the results obtained utilizing threshold models.

The CPWG consists of NRC staff members and consultants. The organizational makeup of this group and their primary responsibilities are shown in Figure 1.1.

NUREG-1037 1-2

ASTPO SARP SEVERE ACCIDENT E

x SOURCE TERM SENIOR REVIEW GROUP POLICY DECISIONS T D.ROSS RESfrNRR R.BERNERO O

~

I E CONTAINMENT LOADS FA URE/ LEA AGE T. SPEIS R.BERNERO I I

! CONTAINMENT LOADS CONTAINMENT CONTAINMENT -

i GROUP m BEHAVIOR GROUP _

PERFORMANCE GROUP 1 -

M. SILBERBERG M. SILBERBERG V.NOONAN j I G. BAGCHI J. COSTELLO l

' ' "AM STUDY GROUP LEADERS STUDY GROUP LEADER j ROSENTHAL J.SHAPAKER G.BAGCHI 7

w J. TELFORD R. WRIGHT P. NIYOGI P. N1 YOGI J. TELFORD l

CONTAINMENT Y

' LOADS INPUT TO CONSULTANTS NU REG-0956 I I I CONTAINMENT CONTAINMENT NON-UNIFORM PERFORMANCE RESPONSE STUDIES THERMAL EFFECTS CONSULTANTS W. BUTLER K. PARCZEWSKI BNL C. HOFMAYER J.SHAPAKER BNL B. MILLER J. HUANG BNL W.T.PRATT R. PALLA INEL T. PRIDGES SANDIA W. SEBRELL AMES L. GREIMANN Figure 1.1 Organization of the Containment Performance Working Group i

N yn

~

2 OVERVIEW OF CONTAINMENT STUDIES . ,,

2.1[ Approach-The. approach.~taken in conducting this study involves'a detaile'd review of con-

.tainment penetration designs and an analytical treatment of. penetration.per-

-formance to estimate. leak area as a. function of predicted pressure and tem-

~

'I.

p'erature conditions. The composite-leak-area'models are then used for selected '

accident sequences to calculate theicontainment' response;and, associated leak crate. Referenc~e containment designs of_each type that were selected for this study.are:

'(1) 'PWR,.large dry - Zion >

(2) PWR,;subatmospheric. ,Surry

-(3) .-PWR,. ice condenser - Sequoyah-(4) ' BWR, Mark I - Peach . Bottom / Browns Ferry

,(5) BWR, Mark II'- Limerick

'(6) BWR,. Mark III - Grand Gulf These. plants were.. selected for the containment leakage study on the basis that

.probabilistic risk assessment (PRA) studies have.already been performed for them and also for consistency with other aspects of the NRC. source term reassess- ;

ment, namely the methods development (Battelle, BMI-2104) and the Containment

~ '

. Loads Working Group (NUREG-1079).

l- Leak area estimates for different components.and penetrations of each contain-ment type were developed by.the following three organizations:

(1) . Idaho National Engineering Laboratory (INEL) ~1esk areas for large opening penetrations (2) Orookhaven National Laboratory ~~(BNL) ' pre-exisEing leak areas and leak areas for purge and vent. valves.

-(3) Sandia National Laboratory (SNL') - leak' areas for electrical penetrations i BNL also assisted the NRC staff in integrating the. leak area estimates from each' separate orgacization to develop a total estimate of containment leakage.

In addition, BNL provided an overview regarding the limitations and signifi-cance of the results.

. Containment? response studies were conducted by BNL to' incorporate containment.

leakage models into existing containment computer codes and to perfctm contain-

. ment response analyses for certain important severe accident sequences for a ;

representative plant of each containment type. ,

As, discussed'in Section.1, the CPWG attempted to identify upper bound leak j area estimates'for each containment type. The approach used in determining these leak areas is discussed in Section 2.3. The leak areas in this report

! are presented primarily as a function of pressure. A preliminary modal to i

NUREG-1037 2-1

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_._ _ y I

5 treat leak area as a function of temperature is also presented; however, it is recognized that.this approach is.very crude and needs more supporting test data i before it can be fully utilized. A more realistic model should treat leakage as a combined _ functi_on of. pressure and temperature, as well as other parameters such as aging and radiation. Although the leakage models are primarily pre-sented as a function of pressure, they qualitatively recognize that the asso-ciated leakage will occur only if seal degradation has occurred as a result

. of high temperatures, aging, or other effects.

Although specific plants have been selected for this stu'y, d it should be noted that many containments ^have similar details, even between different containment types. For example, the purge system valves of one specific manufacturer can j be found on both PWR and BWR plants. Therefore, the evaluations of the six con-

, tainments in this study probably have identified most unique features that could t

have a significant effect on containment leakage. However, further investiga-tions are necessary to determine if such features are applicable to or could l, significantly affect the conclusions concerning the behavior of each contain-ment type. Furthermore,.a sensitivity study of the potential variation of con-tainment leakage within the family of each containment type should be performed to better establish the. uncertainty bounds.

! 2.2 Accident Sequences Considered In this'section we discuss those severe accident sequences that are to be con- !

sidered as part of the containment leakage study. In the previous section six !

containment types were. identified for individual consideration. Each of these

, containment types"has iinique contributors to core melt which can present an d

early challenge to containment integrity. In the'following sections the various

severe accident sequences pertinent to each of the containment types are des-cribed. In addition, we. attempt to define the severe containment loads (pres-

sure/ temperature histories) appropriate to each of these accident sequences. ,

The containment loads were calculated assuming minimum leakage and used to estimate potential leakage areas (refer to Section 2.3). This activity repre-sents the link between the activities of the Containment Performance Working

, Group (CPWG) and the Containment Loads Working Group (CLWG) (NRC, NUREG-1079).

The Containment Loads Working Group consists of NRC staff members and con-

j. sultants. The organizational makeup of this group and their relationship to

( the CPWG are shown~in Figure 1.li The CLWG has-developed a number of standard L problems to test containment integrity. These standard problems were developed

( specifically to serve as a basis for discussions regarding uncertainties and different approaches to the analyses of severe containment loads during core melt accidents. The' standard problems were designed to address some of the l

most severe containment loads expected during core meltdown accident sequences for each of the six containment types under consideration. Obviously, one i ' standard problem per containment type cannot cover all possible accident i sequences and containment failure modes. However, from the information provided

, by.the CLWG it was possible to select the most severe pressure / temperature i histories that could be expected to occur over an extended period of time in the containment buildings during postulated core meltdown accidents. In the following sections we will, factor into our discussions of the various accident j sequences, the findings of the.CLWG. However, it was not always possible to ,

[ consider the most updated containment loads.from the CLWG. In addition, it was ;

not obvious in advance which containment loading characteristics represent the ;

1 I

.NUREG-1037. 2-2

greatest challenge to containment integrity. Consequently, this link between the two groups was highly iterative. The CPWG identified those containment leads that pose the greatest threat to containment integrity. This criterion focused the activities of the CLWG to more accurately characterize the loads.

The results of the CLWG are reported in NUREG-1079 (NRC).

2.2.1 PWR Large Dry Containment The response of this particular containment type to severe accidents has been the subject of ex'.ensive analyses. This is because the Zion and Indian Point plants received particular attention because such large and dense populations surround the two sites. The Zion and Indian Point facilities are PWRs with large dry containments. The Zion containment configuration is shown in Figure 2.1. The utilities and the NRC staff have conducted parallel studies of severe accident sequences for the Zion and Indian Point plants. A preliminary report, which represented the NRC staff's initial contribution to the parallel staff / utility study, was published in October 1981 (NRC, NUREG-0850, Vol.1).

On September 17, 1981, the NRC received the Zion Probabilistic Safety Study l (ZPSS) from Commonwealth Edison Company (CECO), which represented Ceco's contri- I bution to the parallel staff / utility study. In March 1982 the Power Authority of the State of New York and the Consolidated Edison Company submitted the Indian Point Probabilistic Safety Study (IPPSS). The NRC and its contractors have reviewed both of these submittals in great detail. The IPPSS was reviewed cs part of the NRC testimony presented at the Indian Point hearings. Volume 1 of the ZPSS review (NRC, NUREG/CR-3300) has been published and Volume 2 is cur-rently under peer review and will be published within the summer of 1985.

The above independent and parallel studies performed by the utilities and the NRC and its contractors provide a very detailed assessment of severe accidents in PWRs with large dry containment buildings. Both the NRC and utilities' studies concluded that for a wide range of core melt accident sequences, the containment pressure / temperature loads would not exceed the predicted ultimate capacity of the containment building. This is a major change from the conclus-ions offered in the Reactor Safety Study (RSS; also referred to as WASH-1400; cvailable as NUREG-75/014), which indicated a significantly higher potential for all core melt sequences to result in exceeding the ultimate capacity of the containment building.

More recent studies (NRC, NUREG-0850, Vol. 1; CECO, ZPSS; PASNY, IPPSS; NRC, Testimony for Indian Point; (NRC, NUREG/CR-3300, Vols.1 and 2) indicate a low potential for the containment ultimate capacity to be exceeded if the contain-ment heat removal systems (CHRSs) are operating. Without CHRS operation, the containment ultimate capacity may eventually be exceeded but only very gradually and several hours after reactor vessel failure. As a result, the potential for significant containment leakage before the ultimate capacity of the containment is exceeded should be reviewed carefully.

The above studies have focused our attention on those accident sequences in which the emergency core cooling systems (ECCSs) have failed in injection (lead-ing to early core melt) and in which the CHRSs are also not operating. Accident sequences initiated by a loss of all alternating current (ac) power and coupled with failure of the auxiliary feedwater system have been shown (NRC, NUREG-0850, Vol. 1; CECO, ZPSS; NRC, Testimony for Indian Point) to be the major contributors to the class of accidents without ECCS and CHRS operation. These accident NUREG-1037 2-3

~ _ __ _ _ _ _ _ _ _

sequences were designated as TMLB' in the Reactor Safety Study and as TE in the probabilistic safety studies for Zion and Indian Point and in NUREG/CR-3300 (NRC)

(Vols. 1 and 2). Other accidents with CHRS operating would not be expected to result in significant containment leakage because the atmospheric pressure and temperature remain at relatively low levels during these sequences.

A typical pressure / temperature history for a TMLB' or TE sequence is shown in Figure 2.2 for the Zion containment. This sequence results in failure of the heat removal capability of the secondary system, which eventually results in overheating of the primary system. The primary system water boils at the set-point of the pressurizer safety / relief valves. Without CHRS operation, the containment building pressure will gradually increase from relief of primary system steam through the relief valves. Without ECCS operation, the reactor core will eventually uncover, melt, and slump into the bottom of the reactor vessel.

When the molten core contacts the remaining water in the bottom head of the reactor vessel a large quantity of steam could be produced. This steam would be released through the relief valves and would increase the containment pressure. With hot core materials in the bottom vessel head and a high primary system pressure, the head is predicted to fail rapidly. The depressurization of the primary system results in a further rapid increase in containment pres-sure. As the core materials are released into the reactor cavity, they could contact water. Even if the cavity is dry before the vessel fails, the accumu-lators would dump water from the depressurized primary system onto the core debris in the cavity for this accident sequence. The interactions of core de-bris with water in the cavity can, depending on the mode of contact, generate significant steam, hence containment pressurization. The pressure / temperature history in Figure 2.2 reflects rapid mixing of the water and core debris, which results in the very severe pressure spike shortly after reactor vessel failure.

Note that even with this very conservative assumption, the pressure spike at reactor vessel failure does not exceed the predicted capability pressure of the Zion containment (149 psia).

The CLWG analyzed the response of the Zion containment to a TMLB' or TE sequence by reference to the large dry PWR standard problem (memorandum, February 15, 1984).

The consensus of the group was that (for sequences with high pressure discharge of core material from the reactor vessel) rapid quenching of 100% of the released core materials should be assumed. Consequently, the calculations in Figure 2.2 which reflect rapid quenching of 100% of the core, are limiting calculations and consistent with a high pressure primary system failure. A low pressure release of core debris would result in slower quenching of the core debris and hence more gradual pressure rises than shown in Figure 2.2.

One remaining issue to be resolved by the CLWG relates to high primary system pressure failure sequences. Recent SNL experiments have indicated the poten-tial for sweeping fine particles of core debris out of the reactor cavity.

This core debris might then directly heat the containment atmosphere resulting in a rapid pressure / temperature pulse. The direct heating of the atmosphere (as opposed to boiling water and pressurizing containment) would not of itself pose'a direct threat to containment integrity. However, if the direct heating is coupled with exothermic chemical reactions (such as zirconium oxidation),

then a direct pressure threat may exist close to the time of reactor vessel failure. This issue is being pursued by the CLWG. It should, however, be NUREG-1037 2-4

. . - .~ . . . - - -

,e T

5 lnoted that this would be'a narrow pressure / temperature spike, which would be un-likely to cause' seal: degradation (refer to Appendix A) but may be a potentially important1 direct threat to containment integrity.

Ifthecoredebrisandwaterarewelljaixed,thetemperatureofthecore materials will-decrease rapidly. After.the core debris is cool, the subse-

-quent behavior of..the resulting debris bed depends on the supply of water to the reactor cavity-and whether or not a "coolable-debris bed" is formed. .The term coolable'< implies that the decay heat in the core materials can be contin-

-uously removed by boiling water. This will maintain the core debris at a relatively low temperature and the supply of water will be adequate. If this' state can be established, interaction..f.-the core materials and the concrete reactor cavity will be _ limited and subsequent containment pressurization ~ will result from boiling water.

For the Zion and Indian Point plants, a continuous supply of water to the reac--

tor. cavity cannot be guaranteed for TMLB'-or TE sequence. Eventually, as the

pressure.in' containment rises, the water supply to the reactor cavity will.be

-interrupted. The point at which the water supply to the cavity stops. depends on.the, amount of water retained in sumps and on floors. The containment re-'

e .sponse;in Figure 2.2 assumes that sufficient water.is supplied to the reactor cavity.so that the core debris remains. flooded up to the point of-containment

. failure. .From an inspection of Figure 2.2 it is clear that the atmospheric ,

_ ' temperature corresponds closely to saturated conditions appropriate to the containment pressure. This is to be_ expected because boiling water is the primary source'of pressurization.

i -If'the water supply-to the reactor cavity is-limited, the core materials will-dr)out,heatup,andbegintoattackconcrete. During extensive core / concrete interactions, high temperature gases will_be released from the decomposing con-

-crete. Under.some. circumstances these gases could superheat the containment atmosphere. Thus,.a dry cavity could represent a greater threat to seal integ- I rity:because of'the high temperature environment in containment. The extent of core / concrete interactions depends on-the time after scram (decay heat level),,the temperature and composition of the core materials, concrete type, core / concrete. interfacial area',. upward heat transfer from the core materials,.

Cnd communicaton between the reactor cavity and containment. However, sensi -

tivity calculations performed for the CLWG, which considered all of the above

-parameters, indicated bulk. containment atmospheric temperatures below 300*F for a large dry containment. These atmospheric temperatures were not found.to threaten seal integrity. Thus, seal integrity during extensive core / concrete Linteractions was found to be.less of a concern for large volume containments

.than for smaller volume containments (refer to Sections 3.1.4 and 3.1.5).

i ;2.2.21 PWR Subatmospheric Containment

! -The'PWR analyzed in the Reactor Safety Study (NRC, NUREG-75/014) was the Surry i plant, which has a'subatmospheric containment as shown in Figure 2.3. The RSS

'showed that TMLB' accident sequences can provide an early challenge to the inte-

.grity of the Surry containments. Consequently, the RSS also directs our atten-tion to. accident sequences similar to those considered for PWRs with large dry i containments. . However, in the RSS it was concluded that the containment would fail close.to the. time of vessel failure with a very high conditional probabil,ity 3

(CP = 0.8,. refer to-Appendix V'of the Reactor Safety Study).

NUREG-1037 2-5 .

+

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Recent assessments.of the ultimate capacity of the Surry containment have indicated that the RSS may have underestimated the capacity of the containment building. Consequently, a reanalysis of the Surry plant may also indicate a lower potential for exceeding the ultimate capacity of the containment at the time of. vessel failure (as for the Zion and Indian Point containments). In addition the'Mi11 stone Unit 3 Probabilistic Safety Study (Northeast Utilities) has.recently been published and has,been reviewed (NUREG/CR-4143) by NRC staff.

Millstone Unit'3 is also a PWR with a subatmospheric containment. The study indicates a high ultimate capacity containment, which results in containment failure for those accidents without CHRS operation many hours after vessel failure.

The Millstone Unit 3 and the Surry reactor cavity configurations are different from the Zion and Indian Point reactor cavity designs. In the Millstone Unit 3 and Surry designs, the reactor cavity will be dry for a large number of accident sequences. Hence, extensive core / concrete interactions will occur, which could result-in significant generation of combustible gas. Therefore, the potential for a hydrogen burn failure mode is higher at Millstone Unit 3 and Surry than at the Zion and Indian Point facilities. However, we have shown in Appendix A, that short-term hydrogen burns are unlikely to significantly degrade seals.

. Consequent,1y, accident sequence that,invoive gradual pressure and temperature buildup continue to be the focus of the Containment Performance Working Group.

The higher hydrogen probabilities at Millstone Unit 3 and Surry (relative to the Zion and Indian Point plants) represent a direct early threat to containment integrity, which must of course be considered in any evaluation of overall risk at these plants.

A. typical pressure / temperature history for a TMLB' or TE sequence is shown in Figure 2.4 for the Surry containment. These sequences assume no depressuri-zation of the primary system before vessel failure and no supply of water to the reactor cavity other than accumulator discharge. The pressure spike in Figure 2.4 again reflects the rapid quenching of 100% of the core and shows a similar response to the equivalent Zion sequence (refer to Figure 2.2).

It was noted in Section 2.2.1 that containment loads during core / concrete inter-actions depend on a number of parameters and also on the core / concrete inter-action model. The original core / concrete interaction model (INTER) (see Sandia report SAND 77-0370) in MARCH 1.1 (NRC, NUREG/CR-1711) was used in earlier drafts of this report to estimate seal leakage. However, it was found by the CLWG (see NRC, NUREG-1079) to overpredict gas release rates and atmospheric tempera-ture during concrete decomposition. Thus an improved core / concrete interaction model, namely CORCON (SNL, SANC 80-2415) was used to replace INTER in our latest analysis. The resulting pressure / temperature response using CORCON is given in Figure 2.4. This containment response was used to assess the potential for seal leakage in Section 3.1.2.

The direct heating problem for high primary system pressure accident sequences, which was first discussed in Section 2.2.1 (for large dry containments), is also a concern for subatmospheric containments. This issue is also being pursued by the CLWG specifically for applicability to subatmospheric contain-ments. However, it was noted in Section 2.2.1 that the phenomenon results in a narrow pressure / temperature spike, which may pose an early direct threat to containment integrity but should not influence long-term seal degradation.

NUREG-1037 2-6 b

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l L

2.223 PWR Ice Condenser Containment

Unlike PWRs'with large dry or subatmospheric containments, no major industry PRAs for PWRs with ice condenser containments have been submitted to the NRC.

However, in'one part of the Reactor Safety Study Methodology Application Pro-

, gram (RSSMAP) sponsored by the NRC, the Sequoyah plant was analyzed (NRC, NUREG/CR-1659). Sequoyah is a PWR with an ice condenser containment as shown

, in Figure 2.5.

4 r0ne of.the products of the RSSMAP was a characterization (Cybulskis et al.) of

the' physical processes. associated with a variety of core melt accidents. This consisted of' assessing containment failure modes and their potential probabili-l l ties. In the Sequoyah RSSMAP, hydrogen burn failure modes were found to be l-important contributors to early containment failure. . However, we have shown

,(refer to Appendix A) that short-duration hydrogen burns =re unlikely to degrade seals. _So that although hydrogen burns are an important potential failure mode,

'they are of secondary importance in this report because the CPWG is attempting at:this-time to characterize leak paths that do not result from short= duration pressure spikes that exceed the containment shell capability pressure.

Since the RSSMAP. study, a deliberate ignition system using thermal ignitors has ,

~been installed at Sequoyah. The ability of ignitors to control the release of

, combustible gases to containment.has been the subject of extensive analyses and L

experiments. A recent publication-(NRC, NUREG/CR-3278) gives an analysis of

. hydrogen burning in an ice condenser containment with ignitors installed. It ,

cppears- from these studies that the potential for early containment failure

, caused by hydrogen burning will be less with the hydrogen control system instal-1 led than predicted in NUREG/CR-1659 (NRC), which assumed random ignition sources.

However, for the purposes of assessing the performance of the ice condenser containment we will focus our attention on the TMLB' sequence for which the

. " ignitors are not available.

Typical pressure / temperature histories (without hydrogen burns) for a TMLB'

-sequence are shown in Figure 2.6. Without the ignitors,-hydrogen burns will cccur randomly when the atmosphere is flammable. We have noted in earlier sec-1- tions that short-term hydrogen burns are a threat to containment structural i integrity but do not result in seal degradation. The long-term pressure / temper-

' .ature buildup-in Figure 2.6 was therefore used to assess the potential for fleakage in an ice condenser plant (refer.to Section 3.1.3).

l

~

l 2.2.4- BWR Mark I Containment Peach'Bottoc was the BWR_ana'lyzed_in the Reactor Safety Study having a Mark I containment, as'shown in Figure 2.7. The drywell and wetwell atmospheres are continuously inerted:in the Mark'I. design so that hydrogen combustion is not a

' threat to containment integrity. The RSS found that failure via overpressuriza-tion was an important challenge to containment integrity. However, the NRC-o . sponsored severe accident sequence analysis (SASA) program has analyzed in de-l ' tail >(NRC reports NUREG/CR-2825~, NUREG/CR-2182, and NUREG/CR-2973) the response

_of.a.BWR with a Mark I containment (Browns Ferry was used as the model) to a __

number of severe core melt accident sequences. This work indicated that high

. temperatures in the drywell~during ex-vessel core / concrete interactions could

( result in loss of containment integrity because of seal failure before the

l. ~ pressure exceeded the ultimate capacity of the containment. The latter work I

! .NUREG-1037 2-7 i

V

j provided the impetus for the CPWG.to analyze the response of seals in the drywell to the high temperatures resulting from core / concrete interactions.

The-CLWG developed a BWR Mark I standard problem specifically designed to.

, address temperature predictions in the drywell.

It-was noted in Section 2.2.1 that predicted pressure / temperature histories during core / concrete interactions are sensitive to model assumptions and several parameters such as time after scram, temperature and composition of the core materials, concrete type, core / concrete interfacial area, and upward heat trans-fer from the core materials. Some of these parameters are accident sequence dependent-(e.g., time after scram and initial temperature and composition of' the core materials) and others are plant specific (e.g., concrete type and core /

concrete interfacial area). A full range of sensitivity studies-was performed by the CLWG covering all of the above parameters using different model assump-1 tions (refer to NRC, NUREG-1079). In Figure 2.8 we provide the CLWG predicted l pressure / temperature histories using various model assumptions. Clearly, the drywell temperature can vary widely depending on model assumptions, particu-larly the initial molten core materials (corium) temperature, the interfacial area, and the concrete type. The reader is referred to NUREG-1079 for the de-tails of the sensitivity studies.

2.2.5 BWR Mark II Containment Because such a large population surrounds the Limerick site, extensive analyses of severe accidents have been performed for the facility. The Limerick plant is a BWR with a Mark II containment as shown in Figure 2.9. The Philadelphia Electric Company published the Limerick Generating Station Probabilistic Risk Assessment (PRA). The Limerick PRA assessed the risk from core melt accidents initiated by internal events. The risk associated with external events was reported in the Limerick Severe Accident Risk Assessment (SARA).

Both of the above documents have been reviewed by the NRC staff and their con-tractors at BNL. The Limerick PRA was reviewed in NUREG/CR-3028-(NRC) and the Limerick SARA in a BNL draft-report, August 15, 1983. A containment failure mode and fission product release analysis has been performed (BNL-NUREG-33835)

. at BNL for input to the NRC staff's Draft Environmental Statement (DES), (NRC, NUREG-0974, June 1983) and Final Environmental Statement (FES) (NRC, NUREG-0974, April 1984)~ for Limerick.

The Limerick PRA and the BNL review both found that transients with loss of all coolant makeup (TQUV) were the highest probability sequence that led to contain-ment failure at Limerick. However, these sequences were found not to result in early fatalities because of the long time between the start of core damage (and fission product release from the fuel) and eventual containment failure result-ing from overpressurization. It therefore becomes important at Limerick to 4

- determine if the timing of containment failure is late, via overpressurization, or perhaps early, because of high temperature degradation of seals and hence leakage. Therefore the problem for the BWR Mark II containments is similar to the Mark I problem (refer to Section 2.2.4). Consequently, the CLWG sample problem for the Mark II containment was formulated to focus on drywell temper-ature-estimates in a similar manner to the Mark I problem. The only additional concern.(relative to Mark I) relates to the fact that the wetwell and drywell are directly above cach other and separated by the diaphragm floor in the Mark II NUREG-1037 2-8

(see Figure 2.9) configuration. Water cannot accumulate to significant depths on the diaphragm floor so that-as core materials are released from the vessel, the initial interactions will be limited. Subsequent core debris / water inter-cctions depend on how rapidly the core debris passes through the diaphragm floor into the suppression pool. It is the manner in which the core materials pass through the diaphragm floor that depends heavily on plant-specific features.

In some Mark II designs (see Figure 2.10), several downcomers directly underneath the vessel or drains in the floor leading to metal holding tanks beneath the diaphragm floor ensure rapid movement of core materials from the drywell into the suppression pool, whereas in other Mark II designs the core materials must melt through the concrete of the diaphragm floor before reaching the suppression pool. In other designs the region within the pedestal wall in the wetwell is dry so that core debris moving directly down from the vessel will not contact significant quantities of water. For the Limerick design we predict most of the core debris will initially remain on the diaphragm floor so that the above is not a concern. However, the above discussion should be considered when applying this work to other Mark II containment designs.

The CLWG predicted pressure / temperature histories for a Mark II containment using various model assumptions are shown in Figure 2.11. Because the Mark II containment volume is larger than the Mark I containment (249,000 cubic feet for Mark II and 159,000 cubic feet for Mark I drywell volume), differences in model assumptions do not result in as large a variation in the Mark II results as was found for the Mark I results (refer to Figure 2.8). Again a full range of sensitivity studies was performed by the CLWG covering all of the parameters noted in earlier sections (refer to NUREG-1079). However, because of the larger volume of the Mark II containment, the predicted drywell pressure / temperature

' histories were less sensitive to the various parameters.

2.2.6 BWR Mark III Containment The response of this particular containment type to severe accidents has been the subject of extensive analyses. This is because General Electric Company (GE) submitted to the NRC a PRA for GE's proposed standard plant (February 1982).

The standard plant contains a BWR with a Mark III containment. The PRA has been subjected to extensive review and reevaluation by the NRC'(NUREG-0979) and its contractors (NUREG/CR-4135). In addition, the Grand Gulf plant (also a BWR with a Mark III containment) was analyzed as part of RSSMAP (NUREG/CR-1659). A typical BWR with a Mark III containment is shown in Figure 2.12.

The above studies found that hydrogen combustion phenomena pose an early pres-sure threat to containment integrity. However, BWR owners are now required to comply with the hydrogen rule. This rule has usually been met by installing deliberate ignition devices, which are designed to burn hydrogen in a controlled manner for a range of postulated degraded core (not full core meltdown) acci-dents. As part of the staff's review of deliberate ignition systems for degraded core hydrogen control for Mark III containments, the potential for developing standing flames in the wetwell above the suppression pool was identified as an equipment survival concern. This phenomenon led the CPWG to analyze the response of seals to such diffusion flames.

The CLWG analyzed the local wetwell pressure / temperature histories during the existence of standing flames. A range of in-vessel hydrogen generation rates was postulated and it was assumed that the hydrogen was discharged from the NUREG-1037 2-9 -

F primary system via safety relief lines to the suppression pool. The hydrogen

'then bubbled through the pool _into the wetwell airspace where it was assumed to ignite and produce standing flames above the pool surface. Local heat fluxes on the drywell and containment walls were calculated and the impact on seals was assessed (refer to Appendix A).

Any potential degradation of the drywell wall while the core debris is still in the pressure-vessel could have serious consequences, as any fission products released to the drywell _ could bypass the suppression pool and reach the wetwell without the' benefit of pool scrubbing. The local heat fluxes were found not to degrade seals.during the in vessel hydrogen generation period (refer to Sec-tion 3.1.6.4 and Appendix A).

However, after the core materials are predicted to fail the reactor vessel and fall'into the pedestal region, extensive core / concrete interactions could-occur.- Again,.high superheated temperatures, which are sensitive to model and input parameters.(as noted for the Mark I and II assessments), would occur in the drywell. Figure 2.13. indicates _ a range of drywell pressure / temperature histories, which were used to asse'ss the potential for drywel1 leakage (refer to Section 3.1.6.6 and Appendix A).

~ 2.3 Leak Areas In this section we discuss the criteria used to estimate leakage areas that may result from the severe accident pressure and temperature conditions presented in Section 2.2. The leak area estimates are limited to the pressures corres-ponding to_the capability pressure of each reference plant. This approach is

' discussed further in Section 2.3.1. Pre-existing leak areas are identified for each reference plant as discussed in Section 2.3.2.

.For each containment type, the major penetrations having the greatest potential for leakage are identified and evaluated. These include

-Large Opening Penetrations-

- equipment hatch

- personnel airlock

-- drywell head (BWR)

-- fuel transfer tube

-- CRD removal hatch Purge and Vent System Isolation Valves

' Piping Penetrations Electrical Penetration Assemblies The methods used to estimate leak areas for these items are discussed in

' Sections 2.3.3 through 2.3.6.

Finally, the effect of non-uniform thermal loads on the behavior of elastomers used in penetration seals is discussed in Section 2.3.7.

NUREG-1037' 2-10

2.3.1 Capability Pressure

This study concentrated on identifying potential leak paths through containment l penetrations and did not investigate the overall containment shell capability.

l In most cases, the analyses of the containment penetrations did not consider the effects of containment shell deformations and deflections. The neglect of L such interaction effects is not considered significant if the overall behavior of the containment shell remains close to its elastic limit. Consequently, the capability pressures defined in this study for each containment type are based on the point at which the containment first reaches an initial general yield state (or 1% tendon strain in a prestressed concrete containment). Leak area estimates are provided up to the capability pressure. Beyond this pressure the containment performance is uncertain. If containments do not leak before reach-ing the above defined capability pressure, many investigators believe that the containment will leak at higher strain conditions. On the other hand under NRC-sponsored research at Sandia National Laboratories, tests have been conducted on a number of small-scale (1/32) steel models and a large-scale (1/8) steel model to obtain data on containment capacity and failure modes. Results of these. tests clearly indicate that containment' function is maintained for mem-brane strains between 2% and 3%. The large-scale model incorporated representa-tive openings such as piping penetrations, equipment hatch, personnel airlock, etc., and was tested with nitrogen at ambient temperature. For the large-scale test at membrane strains between 2% and 3%, the total observed leak rate remained below 2% of volume per day. The large-scale model experienced a bursting type of failure at approximately five times its design pressure. Thus, the determina-tion as to whether containments will leak or rupture under high strain condi-tions may be very plant specific. Before further studies are conducted to identify leak paths that may result from large containment strain conditions, the impact of the findings in this report on the overall severe accident risk assessment needs to be assessed. If the capability pressure is reached late in the accident sequence, the offsite consequences may not be appreciably dif-ferent if the containment leaks or fails in a gross manner.

The specific capability pressures used for each reference plant are discussed in Section 3.1. These pressures are based on studies reported elsewhere (TEC, IOCOR Technical Report 10.1; SWEC0 Report TP-84-13; NRC reports NUREG/CP-0033, NUREG/CR-2442, NUREG/CR-1891, and NUREG/CR-1967).

2.3.2 Pre-Existing teakage Since a number of operational incidents, such as those reported by M. B. Weinstein, have resulted in degraded containment integrity, it was felt that the likelihood of pre-existing leak paths should be considered in this study. Initially, some rather large pre-existing leak areas were postulated on the basis of information reported by M. B. Weinstein. However, the likelihood of these leak areas occurring was not well defined. Furthermore, it was decided that large pre-existing leakages are more appropriately handled by accident scenarios that censider failure of the containment to isolate; therefore, such leakages are not addressed in this report.

For completeness, it was decided to include some pre-existing leakage effects in these studies. These leakages ara based on the allowable leakage rate for each reference plant that ranges between 0.1 and 0.5% of the containment volume per day at the calculated peak containment pressure. The pre-existing leak area used for each reference plant is described in Section 3.1. These NUREG-1037 2-11

areas are based on the calculations and procedures described in the BNL letter dated June 29, 1984.

Because definitive estimates of the extent and frequency of pre-existing con-tainment leakage at operating reactors are not presently available, the NRC sponsored a program at Pacific Northwest Laboratory (PNL) to consolidate ex-perience data .in the areas of containment integrity and leaktightness so that a realistic evaluation of the containment performance under various accident-conditions can be performed. As part of this program, PNL will identify poten-tial parameters that may affect the isolation system function or reliability, including containment isolation valve failures, deterioration of penetration seal materials, and electrical penetration failures. Information related to the estimated leakage characteristics of these failures, the failure frequen-cies, the failure duration, and the failure probabilities associated with con-tainment leak paths will be developed as part of this effort. The results of these studies will be assessed to determine whether any additional consideration of pre-existing leakage is required.

2.3.3 Large Opening Penetrations Containment leak area estimates for large opening penetrations were determined for each of the six reference plants for severe accident pressure and tempera-ture conditions. Originally, it was intended that three leak area estimates (high, medium, low) as a function of containment pressure would be determined for each plant. The high leak area was to be one that had a small probability of being exceeded. The medium area was to be a best estimate value. The low leak area was to be one that had a high probability of being exceeded. For penetrations with elastomer-type seals, the leak areas were to be determined by first determining the structural response of the seal joints caused by pressure loading and then using seal behavior assumptions based on engineering judgment or test results to obtain the three desired leak areas. Well into the program a small-scale series of seal behavior tests were conducted to try to determine the seal's behavior characteristics under these conditions. A double 0-ring configuration with neoprene seals and a double tongue-and groove configuration with silicon rubber seals were tested. The results of these tests, contained in Appendix B, suggest that seal leakage for flange separations below 0.030 in.

will be nearly insignificant. Because the testing was limited, additional seal testing was required before attempting to characterize the behavior of seal materials during severe accident conditions. Sandia National Laboratory is currently undertaking such a program for the Office of Nuclear Regulatory Re-search of the NRC. Therefore, rather than determine three containment leak area estimates as had been planned originally, it was decided that only one leak area estimate would be determined for each containment type based on the penetrations' calculated structural response (flange separation) only, realizing that this is an upper bound value. In addition to the assumption that the seals have no re-silience, it was also assumed that the personnel airlocks would have only one door available to resist the containment pressure and furthermore, the thermal effects of the pressure unseating closures were neglected. Both .of these assum-ptions are consistent with the goal of determining upper bound L rk area estimates.

I Section 2 of Appendix B describes the analyses performed to determine the leak '

area estimates for large opening penetrations. In some cases, the above assump-tions resulted in significant leakage at the containment design pressure.

NUREG-1037 2-12

Because containments have been shown to be leak tight for pressures up to their design pressure, it is judged that some credit for seal resilience should be given. Therefore, the total estimate of pressure-dependent leakage assumes that only pre-existing leakage is present up to the containment design pressure.

Beyond this pressure, it is assumed that the seals have no further resilience.

In effect, the separation areas discussed in Appendix B are reduced by the sep-aration area predicted to occur at the containment design pressure.

2.3.4 Purge and Vent System Isolation Valves With regard to containment isolation valves, the large-diameter butterfly valves associated with the purge and vent system are considered to have the greatest potential for containment leakage. This is because these valves generally use a non-metallic seal, whereas other isolation valves, such as gate, globe, and check valves, use a metal-to-metal seal. For most plants, including the six reference plants, the purge and vent valves are designed to the American Society of Mechanical Engineers (ASME) Code and are capable of maintaining their struc-tural integrity under severe accident conditions. For some plant-specific cases, where the valves are designed to the American Water Works Association (AWWA) standard, further structural evaluations may be required. The main concern is that the non-metallic seals between the valve body and disc will become degraded when subjected to the combination of high pressures and temperatures associated with severe accident conditions. For the reference plants investigated, these valves ranged in size from 20 to 42 in. in diameter; there were some smaller bypass valves. For other plants these valves have even larger diameters. The metal-to-metal clearance between the valve disc and body is normally between 1/16 and 1/8 in. (letter, June 29, 1984). Therefore, for a 42-in.-diameter valve an upper bound estimate of the leak area (if the seal totally failed) would be approximately 16 in.2 The potential leak area could be even greater because there are normally at least two containment penetrations containing these valves. Using this approach, and considering the number of containment penetrations with these valves, the upper bound leak area attributable to purge and vent valve leakage is determined for each reference plant. These leak areas are discussed in Section 3 and are based on the calculations included in the BNL letter dated June 29, 1984. These calculations also neglect the presence of two redundant isolation valves in series. This assumption is consistent with the goal of determining upper bound leak area estimates.

Most of the reference plants investigated utilized ethylene propylene as the seal material in their purge and vent valves. Data depicting seal life as a function of time at temperature for basic elastomer compounds are shown in Figure 2.14 (Parker, 0-Ring Handbook). As indicated by Figure 2.14, significant purge valve leakage would not be expected in PWR plants that use this material because the temperatures resulting from severe accident conditions generally do not exceed 400 F for a long enough time to exceed the seal life of the material.

This conclusion is consistent with the test results discussed in Section 3.1.1.5 and by BNL (letter, June 6,1984). On the other hand, for BWRs, which can experi-ence temperatures in excess of 600 F for extended periods of time, significant seal degradation could occur.

To consider thermal degradation of penetration seals, such as those associated with the purge valves, a temperature-dependent leakage model was developed.

The criteria used to define the temperature-dependent leakage model are illus-trated in Figure 2.15. The terms used in this figure are defined as follows:

NUREG-1037 2-13

Soak time - the elapsad time at or above a reference temperature for seal degradation to begin.

Ramp time - the elapsed time at or above a reference temperature following initial seal degradation, for total seal failure to occur.

A1 pre-existing leak area.

A2 - upper bound leak area associated with total seal failure.

For an upper bound high leak area estimate, it is assumed that when the soak time for a given seal material exceeds the exposure time shown in Figure 2.14, the seal will begin to leak. When the ramp time for a given seal material exceeds the exposure time shown in Figure 2.14, it is assumed that the seal will be totally failed.

This model is based solely on engineering judgment and is subject to change when better information becomes available. It is utilized in this study only to give a rough approximation of the containment response to potential seal degradation issues. As noted in Section 2.3.3, Sandia is conducting a testing program that should provide more information on the behavior of seals under high-temperature. conditions.

The above model is useful only for penetrations where leak area does not depend on pressure as well. Ideally, the model should be capable of handling leakage caused by the effects of both pressure and temperature at the same time. Because the information is limited, such a refinement is not warranted. For the present, pressure-dependent and temperature-dependent leakage must be considered separately and appropriate judgments need to be made to consider the consequences if both effects are present at the same time.

2.3.5 Piping Penetrations Data for piping penetrations from the six reference plants were collected and classified into eight different types. These are (1) embedded, (2) flanged, (3) flued head, (4) single bellows, (5) double flued head, (6) flued head /

bellows, (7) double bellows, and (8) process pipe bellows. Two types of load-ing were considered: (1) differential pressure across the penetration, and (2) differential support displacements resulting from the deflections of the containment wall that result from internal pressurization and thermal expansion.

The effect of the differential pressure was evaluated using simple calculations.

The evaluation of the displacement loading was more complex and involved the use of a screening criteria. The details of the piping penetrations considered and the analyses that were performed are described in Section 3 of Appendix B.

Stresses in the piping resulting from the differential pressures across the penetration were found to be negligible. Results of the differential support displacement analyses indicate that for the capability pressures considered in this stuoy (refer to Section 2.3.1), piping penetrations will be in a state ranging from fully clastic in the best case to having plastic hinges formed in the pressure boundary that undergo plastic rotations in the range from 0 to 1 degree. Plastic hinges are predicted to occur exclusively in the piping, NUREG-1037 2-14

primarily near the junction of the penetration to the structure, and also at albows and tees in the pressure boundary.

On the basis of the results of the above studies, it was concluded that the piping penetrations and associated piping for the six reference plants are not likely to contribute to containment leakage before reaching the capability pressures as defined in Section 2.3.1.

2.3.6 Electrical Penetration Assemblies The potential leakage through electrical penetration assemblies (EPAs) for Zion, Surry, Grand Gulf, and Sequoyah were evaluated as described in Appendix C.

These evaluations indicated that the EPAs for these plants would not result in cny significant containment leakage under severe accident conditions. On the basis of these studies, it was judged that the same conclusion would apply to cll six reference plants.

Currently, Sandia is conducting a test program to improve definition of the leakage characteristics of EPAs under severe accident conditions. When com-pleted, the results of this test program will be assessed to determine if addi-tional consideration of EPA leakage is required.

2.3.7 Non-Uniform Thermal Effects In the analysis of postaccident conditions, the survival of containment pene-trations is evaluated for the average environment that is predicted to exist in the containment at different times after an accident. However, in many cases, some phenomena may occur that would produce local releases of thermal energy and, as a result, some of the penetrations may be exposed to considerably higher local heat fluxes than are predicted to exist in the containment atmos-phere. These effects were investigated for PWRs with large dry containments (Zion) and BWRs with Mark III containments. The details of these studies are described in Appendix A.

The review of thermal characteristics of containment penetrations in PWR plants has indicated that the only non-uniformity of thermal environment to which these penetrations would be exposed can be caused by burning of flammable gases (hy-drogen and carbon monoxide) generated in the plant during the transient. These burns may cause local hot spots from the high energy release. These localized energy releases will be of short duration and only the components with relatively small thermal inertia will reach high temperatures. Thermal responses were analytically determined for two components most likely to fail when exposed to elevated temperatures: the seal in the equipment hatch and the gasket in the purge valve. In both cases, the temperature rise, assuming a single glcbal burn in the containment, did not exceed 3F . It was concluded, therefore, that the effect of thermal non-uniformities on containment penetrations in PWR plants is negligible, and would not cause their failure.

In BWR plants the inflatable seals in the personnel airlock may be exposed dur-ing severe accidents to thermal energy fluxes coming from a standing hydrogen flame on the surface of the suppression pool. Although these fluxes may signi-ficantly raise the seal temperature, it still will remain below the value at which the seal may fail. However, later in the accident the high drywell tem-peratures during extensive corium/ concrete interactions may cause the seals to NUREG-1037 2-15

exceed the failure limit and a leak path may develop between the drywell and the containment (wetwell). The results of these analyses are discussed further in Sections 3.1.6.4 and 3.1.6.6.

2.4 Containment Response l

Containment response studies were conducted to incorporate containment leakage models into existing containment computer codes and perform containment response analyses for a number of severe accident sequences for the representative plants.

Specifically, these studies (1) incorporated the containment leakage models into existing containment computer codes, e.g., MARCH (NRC, NUREG/CR-1711) and CONTEMPT-4 (EG&G Report TREE-1979); (2) performed engineering analysis of con-tainment response for important accident sequences for a representative plant of each containment type; and (3) performed limited benchmark and confirmatory analyses using the CONTEMPT-4 code. This effort focused on determining the containment leakage rate resulting from accident-induced loads and determining the effect of the leakage on containment response and hence structural failure i predictions.

2.4.1 Containment Codes Various containment analysis codes have been reviewed, including versions of MARCH (NRC, NUREG/CR-1711), CONTEMPT-4 (EG&G Report TREE-1979), and COMPARE (NRC, NUREG/CR-1185), with regard to their capability to model: (1) containment leakage through a time-varying pressure- or temperature-dependent leakage area; and (2) severe accident phenomena such as core / concrete interactions, and carbon monoxide and carbon dioxide generation. On the basis of this review and on discussions with other NRC staff members as well as personnel from several DOE laboratories, it was concluded that an effective approach toward better charac- j terizing containment leakage during severe accidents was to modify two different '

versions of the MARCH computer code to accept the containment leakage models developed and to use the modified codes exclusively for the containment analy-sis. The modified MARCH codes treat the containment leakage as flow through a i single orifice of pressure or temperature-dependent areas; thus, leakage area '

l input represents a composite leakage area for the various penetrations.

Incorporation of the leakage models into the MARCH codes allowed evaluation of containment leakage simultaneously with the calculation of containment response, thus coupling the two effects and eliminating the need to iteratively determine the effect of containment leakage on containment pressure and temperature.

Two different versions of the MARCH computer code were utilized in the contain-i ment leakage study. For PWR containmerts, a version of MARCH 1.1 (NRC, NUREG/CR-1711) in use at Brookhaven National Laboratory (BNL) was adopted. For BWR containments, an Oak Ridge National Laboratory (0RNL) version, MARCH 1.1B (NRC, NUREG/CR-3179) was used. The ORNL version contains significant code improvements regarding channel box models for BWRs and other features.

MARCH 1.1 was used for this study rather than MARCH 2 because MARCH 2 was not available during the early stages of the work. We therefore continued to use MARCH 1.1 (even after MARCH 2 became available) for consistency. We do not consider the use of MARCH 1.1 to be a problem because this study uses the con-tainment response model in MARCH, which does not differ significantly between version 1.1 and version 2. In addition, the recommended containment loads from NUREG-1037 2-16

the CLWG were input directly to the containment response model in MARCH by overriding existing models (in MARCH) that made differences between the two versions even.less relevant to this study.

2.4.2 Leakage Models In the sections that follow, we describe the modifications made to the MARCH 1.1 and MARCH 1.1B computer codes to model containment leakage. The modifications are described in more detail in Brookhaven's Technical Report A-3736R. In Section 2.3 accident-induced leakage paths are described as being functions of either pressure or time at temperature. The pressure-dependent leakage model is described in Section-2.4.2.1 and the temperature-induced model in Section 2.4.2.2.

2.4.2.1 Pressure-Dependent Leakage Model During the early stages of this study, certain accident-induced leakage paths were found to be functions of containment pressure. When the internal contain-ment pressure exceeded a certain level, a corresponding leak area was assumed to open based on calculated response of containment closures to the accident-induced pressure. This area was assumed to remain constant until a higher con-tainment pressure resulted in a larger leak area. For some accident sequences and containment leakage estimates, the leak area eventually became sufficiently

-large to prevent further containment pressurization. This stepwise pressure-dependent leakage model was incorporated into earlier drafts of this report.

The present pressure-dependent leakage model assumes a continuously increasing leakage area vs. pressure characteristic. Leak areas were calculated at se-lected containment pressures (refer to Section 2.3). The present model assumes that the leak area varies linearly between the selected containment pressures.

In addition, for certain applications the model allows for the leak area to close if the pressure in containment decreases.

2.4.2.2 Temperature-Dependent Leakage Model In this section we describe a model (BNL Technical Report A-3736R) in which leak flow rates can be estimated on the basis of containment temperature and

" time at temperature."

The model is described in Section 2.3.4 and is based on the criteria shown in Figure 2.15. Once the code predicts a containment temperature of the atmos-

~

phere above the reference temperature a timer is set. After the elapsed time exceeds the soak time, the leak area is assumed to increase linearly until the ramp time elapses and the upper bound leakage area is reached.

NUREG-1037 2-17

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NUREG-1037 2-19

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NUREG-1037 2-23

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L

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.00, GENERAL TEMPERATURE LIMITS OF SASic 800 800l SILICON RUBBER ELASTOMER COMPOUNDS 700 w 700 FLUOROELASTOMER ETHYLENE PROPYLENE Er NEOPRENE

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NUREG-1037 2-32 h r -

3 RESULTS The results of leak area and leakage rate estimates for the six reference plants cre summarized below.=

3.1 Leak Area Estimates 3.1.1 PWR Large Dry Containment 3.1.1.1 General Description Lcak area-estimates for the PWR large dry containment are based on design data p;rtaining to the Zion plant. The containment vessel is a steel-lined pre-stressed concrete structure with a vertical cylindrical wall and shallow domed roof supported on a flat foundation slab. The containment free volume is 2,600,000 cubic feet.

During normal operation, the Zion containment operates at atmospheric pressure and at temperatures up to 130 F. The design accident pressure is 47 psig, and the design accident temperature is 271 F.

Leak area estimates are based on the accident conditions discussed in Sec-tion 2.2.1.

3.1.1.2 Capability Pressure Sargent and Lundy, the architect-engineer for Zion, has performed a recent anal-ysis of this containment (TEC,-IDCOR Technical Report 10.1). In this analysis the liner was considered to participate as a load-carrying member and the accep-tince criterion for failure was taken as 1% strain in the prestressing steel t2ndons.

The. failure mode was by hoop tendon yielding. This analysis predicted an ulti-mate internal pressure capacity of 134 psig which is used as the capability pressure for Zion in this study.

3.1.1.3 Pre-Existing Leak Area The pre-existing leak area is based on the maximum allowable leakage rate. For Zion the allowable leakage rate is 0.1% of the containment volume per day at the calculated peak containment pressure. This is equivalent to an allowable leak-age rate of 455 scfh (standard cubic feet per hour of gas flow). Assuming choked flow, the corresponding leak area is calculated to be 0.01 in. (letter, June 29, 1984).

3.1.1.4 Large Opening Penetration Leak Areas The large opening penetrations in the Zion containment that were investigated are (1) the equipment hatch with an integral personnel airlock, (2) the emer-g;ncy escape hatch, and-(3) the fuel transfer tube. The details of these NUREG-1037 3-1

l l

-t I^ penetrations and the calculations performed to estimate' potential leak areas l i are described in Section 2.3.1 of Appendix B.

As reported in. Appendix B, the only significant potential leak source is from the personnel airlock. The personnel airlock door frame is predicted to yield

- at'a pressure of 107 psig and is predicted to have a separation area of 6 in.2

' at a pressure of 134 psig. If the airlock door seals are assumed to have no resilience, then the reported separation area would result in an equivalent leak area. This also assumes that only one of the two redundant air'.ock doors '

is available to resist the cordainment pressure.

3.1.1. 5 Purge and Vent _ Valve Leak Areas .

The Zion containment purge and vent system is equipped with 42-in. butterfly I isolation valves. There are two penetrations for Zion; each penetration has

, two valves installed'in series. These valves are manufactured by Henry Pratt Company and use ethylene propylene for the seat material. 'The maximum leak -

L area associated'with these. valves, assuming total seat failure, would be i 32 in.2 based on.a maximum-metal-to-metal clearance between the valve disc and 4

body of 1/8 in. (letter, June 29, 1984). t j

L' To assess the behavior of these butterfly valves under severe accident condi-tions, a scoping test using a Henry Pratt valve with ethylene propylene. seat '

i material was conducted. Three. separate test sequences were performed during which the test assembly was monitored for leakage. The valve was subjected to i

high-temperature steam at 350*F and 120 psig and 400*F and 232 psig for periods

.. of.up to 11 hours1.273148e-4 days 0.00306 hours 1.818783e-5 weeks 4.1855e-6 months . The results of these tests revealed no seal leakage; however,

! seal degradation was evident (letter, June 6, 1984).

! On the basis of the results of the above tests and the fact that there are two

!. valves in series, it is not expected that the Zion purge and vent valves will' l result in any significant leakage if subjected to the accident conditions discussed in.Section 2.2.1.

l- 3.1.1.6% Total Leak Area Estimate j Sources of pressure-dependent leakage are identified in Section 3.1.1.4. As noted, the reported separation area would result in an equivalent leak area if

[ it is assumed that the seals have no resilience. This assumption would result in a leak area of 0.-7 in.2 at the containment design pressure of 47 psig.

~ Since containments have been shown to be leak tight for pressures up to their i s design pressures, it is judged that some credit-for seal resilience should be l ~ given. Therefore, the total estimate of pressure-dependent leakage assumed that only pre-existing leakage is present up to the containment design pressure (47'psig). Beyond this pressure, it was assumed that the seals have no further resilience. In'effect; the separation area discussed in Section 3.1.1.4 was

' reduced by 0.7 in.2 The total estimate of pressure-dependent leakage which is attributed to the personnel airlock is shown in Table. 3.1.

0

< NUREG-1037 3-2 1

l

3.1.2 PWR Subatmospheric Containment 3.1.2.1 General Description Leak area estimates for the subatmospheric containment are based on design data p rtaining to the Surry plant. The containment vessel is a steel-lined rein-forced concrete structure with a vertical cylindrical wall and hemispherical dome supported on a flat base mat. The containment-free volume is 1,800,000 cubic feet.

The Surry containment normally operates at subatmospheric pressure between 9.0 and 11.0 psia. The temperature of the containment air is between 60 F and 105 F during normal operation and 60 F during shutdown. In addition to the normal subatmospheric conditions, the containment is designed for design-basis acci-d:nt (DBA) conditions corresponding to a design pressure of 45 psig and a design temperature of 150 F.

Leak area estimates are based on the accident conditions discussed in Section 2.2.2.

Table 3.1 Pressure-dependent leak area estimate for PWR large dry containment Pressure, Leak area, psig in.2 0 0.01 47 0.01 107 1. 0 134 5.0 3.1.2.2 Capability Pressure The Surry containment was evaluated as part of the Reactor Safety Study (NRC, NUREG-75/014). In that study it was concluded that the containment, when sub-jected to slow overpressurization by steam, would fail at a pressure between 70 and 100 psig. However, this estimate was very conservative since it did not consider factors such as the liner strength, seismic rebar, and design conserva-tism.

Stone and Webster, the architect-engineer for Surry, performed an analysis of the containment and estimated a containment capacity of 119 psig (SWECO Report TP-84-13). This capacity corresponds to the initiation of a general yield state in the main reinforcement and is a factor of approximately 2.6 times the design differential pressure. Other studies (NRC, NUREG/CP-0033) indicate factors of 2.7 for Indian Point and 2.9 for Zion. Consequently, the Stone and Webster pres-sure of 119 psig appears reasonable and will be used as the capability pressure fcr Surry in this study.

Some investigators believe that concrete cracking will cause the liner to stretch in a non-uniform manner resulting in large liner strains at major crack locations (Rashid, 1983). They predict that the liner will rupture at the early NUREG-1037 3-3

l stages of reinforcement yielding. The potential for this failure mode is being investigated further by the Electric Power Research Institute (EPRI). However, for this study, it is assumed that the liner will not leak before reaching the above calculated containment capability pressure.

3.1.2.3 Pre-Existing Leak Area The pre-existing leak area is based on the maximum allowable leakage rate. For Surry the allowable leakage rate is 0.1% of the contained volume per day at the calculated peak containment pressure. This is equivalent to an allowable leak-age rate of 305 scfh. Assuming choked flow, the corresponding leak area is calculated to be 0.004 in.2 (letter, June 29, 1984).

3.1.2.4 Large Opening Penetration Leak Areas The'large opening penetrations in the Surry containment that were investigated are (1) the equipment hatch with an integral emergency escape hatch, (2) the personnel airlock, (3) the fuel transfer tube, and (4) the construction access hatch at the apex of the containment. The details of these penetrations and the calculations performed to estimate potential leak areas are described in Section 2.3.2 of Appendix 8.

As reported in Appendix B, the potential leak sources are the equipment hatch /

emergency escape hatch flange seal and the personnel airlock. The equipment.

hatch / emergency escape hatch _ flanges are predicted to unseat at a pressure of 81 psig and have a flange separation area of 0.3 in.2 at a pressure of 119 psig.

The personnel airlock is predicted to have a separation area of 0.2 in.2 at a pressure of 119 psig.

The seals for the equipment hatch / emergency escape hatch flange are double silicon rubber 0-rings; the personnel airlock door seals are double neoprene 0-rings. These seals are not expected to become degraded as a result of acci-dent temperature conditions described in Section 2.2.2. However, if it is assumed that these seals have no resilience, then the reported separation areas would result in equivalent leak areas. This also assumes that the. inner air-lock door is open or failed and the full containment pressure is placed on the outer airlock door.

7 3.1.2.5 Purge and Vent Valve Leak Areas The Surry containment purge system supply and exhaust ducts are equipped with two 36-in. butterfly isolation valves in series. One valve is inside the con-l tainment and one valve is outside the containment. The outer exhaust valve is I..

fitted with an 8-in. bypass butterfly valve. In addition, an 18-in. pressure-equalizing butterfly valve is installed on the outside of the containment bet-ween the two supply system isolation valves. These valves are manufactured by Allis Chalmers and the specifications for the valves indicate that the seat material is Hycar. Hycar is a- trade name for nitrile or bunan (NBR), an elas-tomer. The maximum leak area associated with these valves, assuming total seat failure, would be 38 in.2 based on a maximum metal-to-metal clearance between the valve disc and body of 1/8 in. (letter, June 29, 1984).

l As discussed in Section 2.2.2, the temperatures in the containment are not expected to exceed 300*F and in general would average 250*F (see Figure 2.4).

NUREG-1037 3-4 l

p Considering the seal life curves shown in Figure 2.14, these temperature condi-

.tions'should not result in any significant seal degradation. Therefore,

~ although'the purge valves present a potentially large leakage path, this path is not expected to occur for the accident conditions discussed in this report. ;

3.1.2.6 Total Leak Area Estimate Sources.of; pressure-dependent leakage are identified.in Section 3.1.'2.4. As e noted, the reported separation areas would result in equivalent leak areas if it is assumed that the seals have no. resilience. This assumption would result

in a leak area'of 0.1 in.2 at the containment design pressure of 45 psig.

'Because containments have been shown to be leak tight for pressures up to their design pressure, the separation areas discussed in Section 3.1.2.4 are. reduced by 0.1 in.2 (See further discussion in Section 3.1.1.6.) .The' total estimate

cf-pressure-dependent leakage that is attributed .to the equipment-hatch /'

emergency escape-hatch flange and the personnel airlock is shown in Table 3.2. ,

Table 3.2 Pressure-dependent leak area estimate for a-sub-atmospheric containment i - Pressure, Leak area, psig in.2 ;

0 0.004

, 45 0.004 i 81 0.02 119 0. 4 -

E -3.1.3 PWR Ice Condense'r Containment l

3.1.3.1 ? General Description Leak area' estimates for the PWR ice condenser containment are based on design -

data pertaining'to the Sequoyah plant. The containment vess d -is a free-

), standing welded steel structure consisting of a cylindrical wall, a hemispheri-cal dome, and a_ bottom liner plate encased in concrete. The vessel.is provided

with both circumferential and vertical stiffeners on the exterior of the shell. -

L The containment' free volume is 1,200,000 cubic feet. The reactor shield build-g -ing is separate from the containment vessel and encloses a 5-ft-wide annulus.

. The Sequoyah containment normally operates at atmospheric pressure.and tempera-tures between 60*F and 120*F. The containment is designed for a maximum inter-nal' design pressure of 12.0 psig coincident with a temperature of'220*F.

~

Leak area estimates are based on the accident conditions discussed in Section 2.2.3.

j. 3.1.3.2 Capability. Pressure The capacity of the Sequoyah containment was evaluated as described in TEC's .

.IDCOR Technical Report 10.1. By use of large deflection theory for the <

containment shell structure, an inelastic finite element analysis indicated i

NUREG-1037 ~ 3-5

. _ . _ ___ E , _ _ . _ ~ . - . , , _ _ . _ _ _ _ . _ . . - _ _ _ . _ _ . _ _ . _ _ _ _ _

l linear displacement up to an internal static pressure of 48 psig. At a pressure of 50 psig, the pressure-displacement relationship showed the onset of non-linearity as a result of yielding of the cylindrical shell. The report concluded that the containment would be functional at 50 psig and could contain a pressure of 58 psig without gross distortion. Ames Laboratory also performed an analysis of the Sequoyah containment and calculated a mean containment capacity of 60 psig with a standard deviation of 8 psi (NRC, NUREG/CR-1891). As a result of

.these analyses, it was decided to use a pressure of 50 psig as the capability pressure for Sequoyah in this study.

3.1.3.3 Pre-Existing Leak Area The pre-existing leak area for Sequoyah is based on the maximum allowable leak-age rate of 0.25% of the containment volume per day at the calculated peak con-tainment pressure. This is equivalent to an allowable leakage rate of 227 scfh.

Assuming unchoked flow, the corresponding leak area is calculated to be 0.01 in.2 (letter, June 29, 1984).

3.1.3.4 Large Opening Penetration Leak Areas The large opening penetrations in the Sequayah containment that were investigated are (1) the equipment hatch, (2) two identical personnel airlocks, and (3) the fuel transfer tube. The details of the penetrations and the calculations per-formed to estimate potential leak areas are discussed in Section 2.3.6 of Appendix B.

As reported in Appendix B, the two personnel airlocks present the only potential leak source. The two airlocks are predicted to remain elastic up to a pressure of 50 psig and to have a separation area of 0.4 in.2 at that pressure. Both airlock doors have double tongue-and groove seals made of EPDM Compound No. 603.

Because the temperatures in the containment are not expected to exceed 400 F for 1 to 2 hours2.314815e-5 days 5.555556e-4 hours 3.306878e-6 weeks 7.61e-7 months before reaching the containment capability pressure, as indicated in Section 2.2.3, the seals are not expected to become degraded as a result of accident temperature conditions. However, if the airlock door seals are assumed to have no resilience, then the reported separation area would result in an equivalent leak area. This also assumes that only one of the two redundant doors on each airlock is available to resist the containment pressure.

Because the initial studies of the Sequoyah penetrations did not identify any significant leak sources, more detailed analyses were undertaken by Ames Labora-tory to investigate the leakage characteristics of the pressure seating estuip-ment hatch (NUREG/CR-3952). The effects of containment deformation and hatch postbuckling on the distortions at the sealing surface were included in the investigation. A three-dimensional finite element model of the equipment hatch was developed that includes (1) shell elements for the hatch, sleeve, contain-ment plate, and stiffeners, (2) prestressed bar elements for the swing bolts, and (3) friction elements for the seal surface. The results of the finite ele-ment analysis showed that the maximum relative sliding and rotation at the seal surface was 0.9 in, and 1 degrees at a pressure of 82 psig. The buckling load was predicted to occur in the range of 85 to 90 psia, but postbuckling displace-ments did not cause large seal motions. These analyses demonstrated that the Sequoyah equipment hatch should not leak before very large strains develop in the -in. containment shell plate near the springline, which occurs between 50 and 60 psig. It was also concluded that if the hatch should buckle, the post-buckling deformations would not introduce leakage.

NUREG-1037 3-6

3.1.3.5 ' Purge and Vent Valve Leak Areas The Sequoyah containment purge and vent system is equipped with 12-in. and 24-in.

butterfly isolation valves. These valves are manufactured by Henry.Pratt Company and use ethylene propylene for the seat material. The maximum leak area associated with these valves,fassuming total seat failure, would be 78 in.2 based on leakage through eight'24-in. and two 12-in. valves and a maximum metal-to-metal clearance between the valve disc and body of 1/8-in. (letter, June 29, 1984).

As discussed in Soction 3.1.1.5, scoping tests were performed on similar valves cnd it was demonstrated that they would not leak when subjectert to temperatures

'up to 400*F for 11 hours1.273148e-4 days 0.00306 hours 1.818783e-5 weeks 4.1855e-6 months . Thus, it is not expected that the Sequoyah purge and v nt valves will result in any significant leakage if subjected to the accident canditions discussed in Section 2.2.3.

3.1.3.6 Total Leak Area Estimate Sources of pressure-dependent leakage are identified in Section 3.1.3.4. As nsted, the reported separation area would result in an equivalent leak area if it is assumed that the seals have no resilience. This assumption would result in a leak area of 0.1 in.2 at the containment design pressure of 12 psig.

B:cause containments have been shown to be leak tight for pressures up to their d: sign pressure, the separation area discussed in Section 3.1.3.4 is reduced by 0.1 in.2 (See further discussion in Section 3.1.1.6.) The total estimate of pressure-dependent leakage which is attributed to the two personnel airlocks is shown in Table 3.3.

Table 3.3 Pressure-dependent leak area estimate for a PWR ice condenser containment Pressure, Leak area, psig in.2

0. 0.01 12 0.01 50 0.3 3.1.4 BWR Mark I Containment 3.1.4.1 General Description Lcak area estimates for the BWR Mark I containment are based on design data per-taining to the Peach Bottom plant. The containment is a pressure-suppression system which consists of a drywell, a pressure-suppression chamber (torus) that stores a large volume of water, and a connecting vent system between the drywell and the water pool.

The drywell is a lightbulb-shaped steel pressure vessel with a spherical lower Jpsrtion and a cylindrical upper portion. The suppression chamber is a steel pressure vessel in the shape of a torus that is located below and encircles NUREG-1037 3-7 ,

the drywell. Eight circular vent pipes connect the suppression chamber with the drywell. The Peach Bottom drywell and pressure-suppression chamber free volumes are 159,000 and 119,000 cubic feet, respectively. The pressure-suppression chamber pool water volume is 136,000 cubic feet.

During normal operation, the Peach Bottom containment operates at atmospheric pressure and temperatures up to 135 F. The drywell, suppression chamber, and vent system are designed for accident conditions corresponding to a maximum internal pressures of 62 psig coincident with a temperature of 281*F.

Leak area estimates are based on the accident conditions discussed in Sec-tion 2.2.4.

3.1.4.2 Capability Pressure The Peach Bot' tom containment was evaluated as part of the Reactor Safety Study (NRC, NUREG-75/014. In that study it was concluded that the containment would fail at a pressure between 135 and 185 psig. The failure criterion used in this study was a stress. level in the base material halfway between the yield and ultimate' strength of the metal. Furthermore, the mechanical properties utilized in the study were "the expected properties of the material" rather than the minimum specified in the applicable codes'.

Ames Laboratory has performed a recent analysis of the Browns Ferry containment and calculated a mean containment capacity of 117 psig (NRC, NUREG/CR-2442).

The failure criterion in this analysis was defined as the pressure at which strains reach twice the yield value. The material mean yield stress used in this study was assumed to be 10% greater than the minimum specified yield.

Because the current program is concentrating on identifying potential contain-ment leakage sources before reaching large containment deformations, it is felt that the Ames Laboratory containment capacity (117 psig) should be used as the capability pressure for a typical BWR Mark I containment. There may be specific design details for each particular containment to justify the variation in con-tainment capability' pressures as illustrated by the above two studies. However,

.it. is also apparent that the Reactor Safety Study utilized a less conservative failure criterion (material stress halfway between yield and ultimate) that would result in larger containment deformations before reaching the above calcu-lated containment capability pressure.

3.1.4.3 Pre' Existing Leak Area l l

The pre-existing leak area is based on the maximum allowable leakage rate. For l Peach Bottom the allowable leakage rate is 0.5% of the contained volume per day '

at the calculated peak containment pressure. This is equivalent to an allowable leakage rate of 281 scfh. Assuming choked flow, the corresponding leak area is calculated to be 0.004 in.2 (letter, June 29, 1984). l 3.1.4.4 Large Opening Penetration Leak Areas The large opening penetrations in the Peach Bottom containment that were inves-tigated are (1) the drywell head, (2) the two drywell equipment access hatches, (3) the personnel airlock, and (4) the control rod drive (CRD) removal hatch.

The details of these penetrations and the calculations performed to estimate !

potential leak areas are described in Section 2.3.3 of Appendix B.

NUREG-1037' 3-8

No data were obtained for the two 54-in.-diameter suppression pool and the 24-in.-diameter drywell head access hatches. Because temperatures in the sup-

- pression pool are expected to stay well below those of'the drywell, it was j judged that thermal degradation and subsequent leakage of the suppression pool l hatch seals were unlikely. With regard to the drywell head access hatch, it was .

concluded that if. leakage occurred it would be insignificant in comparison with :

that of the drywell head flange seal.

As reported in Appendix B, potential leakage sources are the drywell head, the -

pressure unseating equipment hatch and the personnel airlock. The drywell head flanges are predicted to unseat at a pressure of 27 psig and have a flange sepa- '

ration area of 53 in.2 at a pressure of 117 psig. The equipment hatch flanges are predicted to unseat at a pressure of 82 psig and have a flange separation area of 2 in.2 at a pressure of 117 psig. .The personnel airlock door frame is predicted to yield at 94 psig and is predicted to have a separation area of 1 in.2 at 117 psig.

All of-the above openings have silicon rubber seals. If it'is assumed that these seals have no resilience, then the reported separation areas would result

' in equivalent leak areas. This also assumes that only one of the two redundant personnel airlock doors is available to resist the containment pressure.

3.1.4.5 Purge and Vent Valve Leak Areas

- The Peach Bottom containment purge and vent system has four lines (two in the drywell-and two in the suppression chamber) each having two 18-in. valves in series. There is also a bypass line with a 6-in. valve in series with an 18-in.

valve. These valves are manufactured by Fischer and use ethylene propylene for the seat material. The 18-in. valves have a disc diameter of 17.5 in. and a

- metal-to-metal clearance of up to 1/8 in. Therefore, the maximum leak area associated with one valve, assuming total seat failure, would be 7 in.2 (letter, June 29,-1984).

. As indicated in Figure 2.8, the temperatures in the drywell could reach 500 F for approximately 2 hours2.314815e-5 days 5.555556e-4 hours 3.306878e-6 weeks 7.61e-7 months . Some estimates indicate temperatures in excess of

- 700 F for approximately 1 hour1.157407e-5 days 2.777778e-4 hours 1.653439e-6 weeks 3.805e-7 months . On the basis of the seal life curves shown in Figure 2.14, these temperature conditions could degrade the ethylene propylene purge valve seals associated with the two lines in the drywell. Thus, the dry-well'could experience a maximum leak area associated with the purge valves of 14 in.2, 3.1.4.6 Total Leak Area Estimate Sources of pressure-dependent leakage are identified in Section 3.1.4.4. As noted, the reported separation areas would _ result in equivalent leak areas, if it is assumed that the_ seals have no resilience. This assumption would result in a_ leak area of 21 in'.2 at the containment design pressure of.62 psig. Because containments have been shown to be leaktight for pressures up to their design pressure, the separation area discussed in Section 3.1.4.4 is reduced by 21 in.2 (See further discussion in Section 3.1.1.6.) The total estimate of pressure-dependent. leakage attributed to the drywell, equipment hatch, and per-sonnel airlock is shown in Table 3.4.

4 NUREG-1037 3-9

Table 3.4 Pressure-dependent leak area estimate for a BWR Mark I containment

" Pressure, Leak area, psig in.2 0 0.004 62- 0.004 82 12.0 94 19.0 117 35.0 As noted, the leak area estimates in Table 3.4 are attributed to the potential lack of sufficient seal resilience. Although not taken into account, the seals could also become degraded as a result of the high containment temperatures.

Section 2.2.4 indicates that temperatures in the drywell could exceed 700 F.

Thus, the seal life, even for silicon rubber, could be affected.

As discussed in Section 3.1.4.5, the high drywell temperatures could have their greatest effect on the purge valve seals. Utilizing the procedures discussed in Section 2.3.4 and the seal life curves shown in Figure 2.14, a temperature-dependent leakage model as shown in Table 3.5 also has been developed.

Table 3.5 Temperature-dependent leak area estimate for a BWR Mark I containment Parameter Estimate Temperature 500 F Soak time 40 min.

Ramp time 40 min.

A1 0.004 in.2 A2 14 in.2 3.1.5- BWR Mark II Containment 3.1.5.1' General Description Leak area estimates for the BWR Mark II containment are based on design data pertaining to the Limerick plant. The containment is a pressure-suppression system that consists of a drywell and suppression chamber separated by a hori-zontal diaphragm slab. The containment is in the form of a truncated cone over a cylindrical section. The upper conical section forms the drywell and the lower cylindrical section forms the suppression chamber. The containment is a reinforced concrete structure, lined with welded steel plate and provided with a steel domed closure head at the top of the drywell. The Limerick drywell free volume is 249,000 cubic feet and the suppression chamber free volume can vary between 161,000 and 149,000 cubic feet The pressure-suppression chamber pool water volume can vary between 116,000 and 128,000 cubic feet.

NUREG-1037 3-10

During normal operation the Limerick containment operates at temperatures up to 135"F. .The containment is designed for a maximum internal design pressure of 55 psig and temperatures in the drywell of up to 340*F and temperatures in the

. suppression chamber of up to 220*F.

Leak area estimates are based on the accident conditions discussed in Section 2.2.5.

3.1.5.2 Capability Pressure

- The ultimate pressure capacity of the containment was calculated as part of the Probabilistic Risk Assessment Study for Limerick (TEC, IDCOR Technical Re-port:10.1). Utilizing an idealized model of the cylindrical portion of the suppression chamber, general yielding of the hoco reinforcement was predicted

.st a pressure of 120 psig. Additional analyses indicated that an internal pressure of 170 psig could be considered as an upper bound where actual catas-trophic failure could be expected. Consideration of the complexity of the struc-tural behavior led to the conclusion that 140 psig was an acceptable upper limit

- of the internal pressure. Therefore, this pressure is used as the capability

-pressure for Limerick in this study.

3.1.5.3 Pre-Existing Leak. Area The pre-existing leak area for Limerick is based on the maximum allowable leak-age rate of 0.5% of the containment volume per day at the calculated peak con-tainment pressure. This is equivalent to an allowable leakage rate of 404 scfh.

Assuming choked flow, the corresponding leak area is calculated to be 0.005 in.2 (letter, June 29, 1984).

3.1.5.4 Large Opening Penetration Leak Areas The large opening penetrations in the Limerick containment that were investi-gated are (1) the drywell head, (2) the two drywell equipment hatches, (3) the personnel airlock, (4) the.two suppression chamber access hatches, and (5) the control rod drive removal hatch. The details of these penetrations and the calculations performed to estimate potential leak areas are described in Section 2.3.4 of Appendix B.

As reported in Appendix B, potential leak sources are the drywell head, the two drywell equipment hatches (both are pressure unseating), and the suppression chamber access hatches. The drywell head ' flanges are predicted to unseat at a pressure of 85 psig and have a flange separation area of 33 in.2 at a pressure of 140 psig. The equipment hatch flanges are predicted to unseat at a pressure

. of 75 psig and have a flange separation area of 9 in.2 at a pressure of 140 psig. The suppression chamber access hatches were analyzed with no preload and have a. separation area of 1.5 in.2 This last separation area is probably too high because procedures are being modified on Limerick to implement a specified preload for the suppression chamber hatches.

4 Separation areas were also calculated for-the CRD removal hatch and the drywell heaJ manhole on the basis of no preload in the bolts. However, these areas are small compared with the other areas and have been neglected in further calcula-tions. Furthermore, procedures at Limerick are being modified to implement a preload for both of these openings. >

NUREG-1037- 3-11

All of the above openings have silicon rubber seals. If it is assumed that these seals have no resilience, the reported separation areas would result in equivalent leak areas.

3.1.5.5 Purge and Vent Valve Leak Areas The Limerick containment purge and vent system is equipped with 4 , 6 , 18 , and 24-in tricentric butterfly valves. These valves are manufactured by Clow Cor-poration and use a metal-to-metal seal. The manufacturer reports that this design has been successfully used for sealing applications from cryogenic tem-peratures to 900*F (Clow, Report No. 11-15-83). Thus, these valves are not expected to result in any significant leakage if subjected to the accident con-ditions discussed in Section 2.2.5.

3.1.5.6 Total Leak Area Estimate Sources of pressure-dependent leakage are identified in Section 3.1.5.4. As noted, the separation areas would result in equivalent leak areas if it is assumed that the seals have no resilience. This assumption would result in a leak area of 0.6 in.2 in the suppression chamber at the design pressure of 55 psig. Since containments have been shown to be leaktight for pressures up to their design pressure, the separation area discussed in Section 3.1.5.4 for the suppression chamber is reduced by 0.6 in.2 (See further discussion in Section 3.1.1.6.) The total estimate of pressure-dependent leakage that is attributed to the drywell head, equipment hatches, and sLppression chamber access hatches is shown in Table 3.6.

Table 3.6 Pressure-dependent leak area estimate for a BWR Mark II containment Total leak area, in.2 Pressure, Suppression psig Drywell chamber 0 0.003 0.002 55 0.003 0.002 75 0.003 0.2 85 1. 0 0.3 140 42.0 0.9 As noted, the leak area estimates in Table 3.6 are attributed to the potential lack of sufficient seal resilience. Although not taken into account, the seals could also become degraded as a result of the high containment temperatures.

' Figure 2.11 indicates that temperatures in the drywell could exceed 600*F for 1\ hours. Other studies have predicted temperatures greater than 750*F for approximately 40 minutes and temperatures greater than 650*F for 60 to 80 minutes. Thus, the seal life, even for silicon rubber, could be affected.

NUREG-1037 3-12

r 3.1.6 BWR Mark III Containment 3.1.6.1 General Description Leak area estimates for the BWR Mark III containment are based on design data pertaining to the Grand Gulf plant. The containment is a pressure-suppression l system with the drywell completely enclosed by the containment structure. The lower portion of the containment structure also serves to form the pressure-suppression pool. The containment is a steel-lined, reinforced-concrete struc-ture consisting of a right circular cylinder capped by a hemispherical dome and founded on a flat circular base mat. The drywell is also a reinforced-concrete structure consisting of a cylindrical wall and flat, horizontal, circular roof that contains a circular opening for the drywell head. The drywell head is a short stainless steel cylinder with an elliptical head that is part of the dry-well pressure-retention boundary. The Grand Gulf drywell and containment free volumes are 270,000 and 1,400,000 cubic feet, respectively. The pressure-suppression pool water volume is 136,000' cubic feet.

The containment structure is designed for an internal pressure of 15 psig and a temperature of 185*F. The drywell structure is designed for an internal pressure of 30 psig and a temperature of 330*F.

Leak area estimates are based on the accident conditions discussed in Section 2.2.6.

3.1.6.2 Capability Pressure Bechtel, the architect-engineer for Grand Gulf, has performed an analysis of this containment (TEC, IDCOR Technical Report 10.1). The ultimate' capacity was defined as the pressure at which a general yield state is reached at a critical structural section. This pressuce was found to be 56 psig. However, if actual material strength is used, the mean ultimate capacity is 67 psig, with lower and upper bounds of 62 and 70 psig, respectively. In addition, the lower con-tainment personnel airlock was predicted to have an ultimate capacity of 60 psig.

Brookhaven National Laboratory also analyzed the Grand Gulf containment (NRC, NUREG/CR-1967). The analysis was made on the basis of the ASME Code-specified strengths _and indicated that the hoop reinforcing steel undergoes yielding at a pressure of 52 psig. The failure strength would have been higher if actual material strengths were used. Therefore, as recommended by TEC, a pressure of 60 psig is used as the capability pressure for Grand Gulf in this study.

3.1.6.3 Pre-Existing Leak Area The pre-existing leak area is based on the maximum allowable leakage rate. For

' Grand Gulf the allowable leakage rate is 0.4% of the containment volume per day ,

at the calculated peak containment pressure. This is equivalent to a leak area i of 0.02 in.2 (letter, June 29, 1984).

3.1.6.4 Large Opening Penetration Leak Areas The large opening penetrations in the Grand Gulf containment structure that were investigated are (1) the two personnel airlocks, (2) the equipment hatch, and (3) the fuel transfer tube. The details of these penetrations and the NUREG-1037 3-13

calculations performed to estimate potential leak areas are described in Section 2.3.5 of Appendix B.

As reported in Appendix B, it was determined that containment leakage of these large opening penetrations will not result from pressure loadings up to the capability pressure of 60 psig.

In addition to pressure loading, the personnel airlock door seals were investi-gated for possible thermal degradation as a result of the development of stand-ing flames in the wetwell above the suppression pool. The flames result from the installation of deliberate ignition devices, which are designed to burn hydrogen in a controlled manner for a range of postulated degraded core (not full core meltdown) accidents. The details of this investigation are described in Appendix A and are summarized below.

The accident sequence considered predicted a hydrogen release to the drywell about 100 minutes after the beginning of the accident. _It assumed that the hydrogen passes through the suppression pool and is ignited at the surface of the pool and burns for 50 minutes with a stationary diffusion flame. The heat fluxes from this flame to outside containment and the drywell walls were deter-mined by Sandia using the HECTR computer code. A heat-transfer analysis of the drywell personnel airlock doors was performed using these heat fluxes. The doors are sealed by two inflatable EPDM seals around the door edge between the door and the door frame opening. The seals on the door facing the suppres-sion pool (inner seal) were predicted to reach a temperature of 345 F, whereas the seals on the door facing the drywell (outer seal) were predicted to reach a temperature of 264 F.

On the basis of the seal life curves shown in Fig-ure 2.14, seal degradation would not be expected. On the other hand, it has been indicated that the double inflatable seal design may be inadequate for applications in a temperature environment of 330 F and more (letter, October 7, 1983). However, since the temperatures on the inner door seals are only above 330 F for a brief time, it is expected that leakage will not result from the diffusion flame.

The containment personnel airlocks are the same design as the drywell personnel airlock. Therefore, it is expected that the containment airlocks also will not leak as a result of the above predicted diffusion flames. The seal temperatures for the containment airlock should be less than that predicted above because the outer personnel airlock door is exposed to the environment outside the containment.

3.1.6.5 Purge and Vent Valve Leak Areas The Grand Gulf containment purge and vent system is equipped with 6-in. and 20-in. butterfly isolation valves. These valves are manufactured by Henry Pratt Company and use ethylene propylene for the seat material. The maximum leak area associated with these valves, assuming total seat failure, would be 19 in.2 based upon leakage through two 6-in. and two 20-in. valves and a maximum metal-to-metal clearance between the valve disc and body of 1/8-in.

(letter, June 29, 1984).

As discussed in Section 3.1.1.5, scoping tests were performed on similar valves and it was demonstrated that they would not leak when subjected to temperatures up to 400 F for 11 hours1.273148e-4 days 0.00306 hours 1.818783e-5 weeks 4.1855e-6 months .

Thus, it is not expected that the Grand Gulf purge and vent valves will result in any significant leakage if subjected to the accident conditions discussed in Section 2.2.6.

NUREG-1037 3-14

3.1.6.6 Total Leak Area Estimate On the basis of the results presented in Sections 3.1.6.4 and 3.1.6.5, the Grand Gulf containment is not expected to leak before reaching the capability pressure of 60 psig when subjected to the accident conditions discussed in Section 2.2.6.

The containment pressure and temperature responses used in these studies were based on the assumption that there was no bypass of the suppression pool by -

means of leakage through the drywell wall. As discussed in Appendix F to Appen-dix B, a drywell leak area of approximately 9 in.2 is estimated at a drywell differential pressure of 4 psi. Furthermore, as indicated in Figure 2.13, even without diffusion flames the drywell temperature would rise very rapidly after about 200 minutes into the accident sequence. Some studies have predicted the drywell to reach temperatures in excess of 900 F, remaining at this level until the containment reaches its capability pressure of 60 psig (about 800 minutes into the accident sequence). On the basis of these temperature conditions, a heat transfer analysis was performed for the drywell personnel airlock and it was determined that at about 300 minutes after initiation of the accident sequence, the seal on the drywell side of the airlock will reach temperatures of approximately 870 F, and the seals on the suppression pool side of the air-lock will reach temperatures of approximately 550*F. (See Appendix A for the details of this analysis.) Because these seals will remain at these temperatures until the containment reaches its capability pressure, it is expected that the pressure integrity of the drywell personnel airlock will be lost. As discussed in Appendix F to Appendix B, the upper bound leak area for the personnel air-lock with the seals entirely blown out is approximately 125 in.2 This leak area could occur as early as 5 to 6 hours6.944444e-5 days 0.00167 hours 9.920635e-6 weeks 2.283e-6 months after initiation of the accident.

The effect that the above bypass leakage may have on the containment response is discussed in Section 3.2.6.

3.2 Impact of Leakage on Containment Response The impact of leakage on containment response depends on the magnitude and pres-sure/ temperature dependence of the containment leak area, as well as on the mag-nitude and timing of the containment pressure / temperature loadings. If a detailed analysis of containment penetrations indicates that significant leak paths do not exist beforehand and do not develop before containment failure, then containment loading patterns will have only a marginal effect on the predicted releases from containment. On the other hand, should significant leak areas develop at pres-sures or temperatures less than containment failure, the accident-induced leakage can result in early release frou containment, gradual containment pressure relief, and delay / prevention of containment structural failure by gradual overpressuriza-tion.

Containment response analyses have been performed to determine the magnitude of the leakage through the leak areas estimated for the six containment types iden-tified in Section 3.1. These analyses include the effect of the leakage on the overall containment pressure / temperature response and are described in the sec-tions that follow.

3.2.1 PWR Large Dry Containment Extensive analyses of core meltdown accidents at the Zion facilities have been performed. On the basis of these analyses, a TMLB' accident sequence (using NUREG-1037 3-15

1 1

the notation of the Reactor Safety Study) was selected by the CLWG as a basis for assessing the influence of containment leakage. A TMLB' accident (extended loss of all ac power coupled with failure of the auxiliary feedwater system) results in core meltdown with a high primary system pressure (at the setpoint of the relief valves) and a loss of all containment heat removal systems (CHRSs).

For the Zion containment, the TMLB' accident sequence is predicted to result in containment pressure vessel.

failure many hours after the core debris penetrates the reactor The timing of containment failure was shown in NUREG-0850 (NRC) to be a strong function of the water inventory in the reactor cavity. Decay heat boiling of water in the cavity results in more rapid containment pressurization than if the core debris has to heat up and de gas concrete in order to pressurize the containment.

The quantity of water that would be in the cavity is uncertain; therefore, two reactor cavity configurations have been considered for this leak-age study, namely:

(1) A reactor cavity that is flooded before vessel failure and supplied with sufficient water to maintain a coolable debris bed, thereby eliminating core / concrete interactions (all decay heat is used to boil water).

(2) A reactor cavity that is dry before vessel failure. The only water assumed to reach the cavity is that from the accumulators (after vessel failure and primary system depressurization). The accumulator water is rapidly boiled off during the initial core debris / water interactions resulting in exten-sive core / concrete interactions.

The containment pressure and temperature histories for the no-leakage case are plotted in Figure 3.1. Note that for this flooded-cavity case the core remains quenched and core / concrete interactions do not occur. The containment pressure would not be expected to reach its capability pressure (149 psia) until after 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months . For the case with leakage as described in Section 3.1.1 (plotted as a solid line on Figure 3.1) the leakage rate is sufficient to delay or possibly prevent the containment from reaching its capability pressure. The corresponding containment leakage rates are shown in Figure 3.2.

3.2.2 PWR Subatmospheric Containment The leak area estimates for a PWR with a subatmospheric containment are describ-ed in Section 3.1.2. Details of the containment leakage analysis are given in BNL Technical Report A-3736R. Preliminary leak areas were based on the contain-ment loading calculations for a TMLC' accident (extended loss of all ac power coupled with failure of the auxiliary feedwater system) using MARCH 1.1 (NRC, ,

NUREG/CR-1711) with its core / concrete interactions based on INTER (SNL, SAND 77- '

0370). These pressure and temperature calculations indicated the possibility of significant seal failure and substantial leakage. Note that for Surry the reactor cavity will be essentially dry and the containment loading is very sen-sitive to core / concrete interaction modeling (refer to Section 2.2.2). More recent results using CORCON (SNL, SAND 80-2415) give substantial reductions in both_ temperature and pressure loadings as shown in Figure 2.4. On the basis

~

of the loads given in Figure 2.4 (for Surry) revised leak areas were estimated (refer to Table 3.2). For the Surry containment response to a TMLB', these leak areas are inconsequential. Because the loading estimates in Figure 2.4 never leak areas reach(<0.01 95 psia, the pressure response is not significantly affected by the in.2),

NUREG-1037 3-16

3.2.3 PWR Ice Condenser Containment The estimated leak areas (refer to Section 3.1.3) for the PWR ice condenser plant are extremely small and do not have a significant effect on the contain-ment response given in Figure 2.6.

3.2.4 .BWR Mark I Containment The leak area estimates for a Mark I containment are described in Section 3.1.4.

Details of the containment leakage analysis are discussed in BNL Technical

-Report A-3706. The accident sequence chosen to illustrate the effects of con-tainment leakage is a TQUV-type sequence in which the main steam system is iso-lated and all reactor vessel injection capability is lost at the time of a reac-tor trip from 100% power. Because of mass loss out of the safety / relief valves (SRVs) and the lack of coolant injection, the core eventually becomes uncovered.

In this sequence, automatic depressurization system (ADS) actuation will not occur, and manual actuation is assumed not to occur, so that the reactor coolant system (RCS) remains at high pressure (setpoint of SRVs). The uncovered core becomes molten and the debris falls into the reactor vessel lower plenum where, eventually, the corium attacks the reactor vessel bottom head.

When the reactor vessel bottom head fails, the corium falls onto the dry con-crete floor of the drywell and the corium/ concrete reaction begins. As steam is liberated from the concrete, previously unoxidized zirconium in the corium is oxidized, releasing large amounts of energy. The specific geometry and major assumptions are described in BNL Technical Report A-3706. The specific case analyzed assumes a high-temperature debris (4130 F) with maximum spreading (in order,to maximize gas generation rates). The pressure responses for the no-leakage and leakage cases are shown in Figure 3.3. The pressure rises rapidly because of the generation of noncondensable gases and the capability pressure of the containment (132 psia) is reached within 2 hours2.314815e-5 days 5.555556e-4 hours 3.306878e-6 weeks 7.61e-7 months . hote that the pressure response beyond the capability pressure is also shown. For the case with leakage in the drywell head (refer to Figure 3.3), the pressure remains well below the capability pressure of the containment and after a short peak near 100 psia the pressure drops back to 80 psia where the head would be expected to reseat itself with a very limited leak area (<2 in.2). The corresponding leakage rates are shown in Figure 3.4.

A less likely but potentially more severe leakage path has been identified (Sec-tion 3.1.4) in the containment purge and vent lines. These are double isolation valves that, for the case analyzed, use ethylene propylene for the seat material.

Under the severe thermal loading during core / concrete reactions, the seat may fail completely and open up the metal-to-metal clearance of about 14 in.2 This leakage path is considered to be much less likely than the path through the dry-well head seal because the second seal is well isolated from the containment at-mosphere. The purge and vent valve failure is modeled as a temperature-dependent leakage (see Section 3.1.4) and the leak areas are given in Table 3.5. The pres-sure response is also shown in Figure 3.3. For this case the pressure rises to slightly over 100 psia before the seals are expected to fail. However, once the seals fail, the containment is rapidly depressurized to the atmospheric pressure.

The corresponding leakage rates are also shown in Figure 3.4.

NUREG-1037 3-17

! 3.2.5 BWR Mark II Containment Three. cases involving no leakage (base case), drywell leakage, and wetwell leak-age specified in Table 3.6 were analyzed for the Limerick plant (refer to Fig-ures 3.5 and 3.6). The TQUV sequence was selected for assessing the influence of containment leakage. A TQUV accident (loss of the power conversion system coupled with a failure of all makeup water) results in core meltdown with a high primary system pressure and a loss of all containment heat removal systems.

The MARCH predictions indicate that the reactor vessel lower head fails' at approximately 2 hours2.314815e-5 days 5.555556e-4 hours 3.306878e-6 weeks 7.61e-7 months af ter scram. At this point molten core materials (corium) fall onto the. diaphragm floor and start interacting with concrete.

The corium/ concrete interaction modeled by the MARCH /CORCON analysis predicted a release of a large amount of gases and energy in the first three hours after the contact of corium with the drywell concrete floor when the vessel fails.

This results in a rapid increase of containment pressure and temperature. For the base case (in which no leakage was modeled), the containment pressure be-comes stable 3 hours3.472222e-5 days 8.333333e-4 hours 4.960317e-6 weeks 1.1415e-6 months after vessel failure and the peak pressure is about 130 psia. !

The peak pressure is 25 psia below the containment capability pressure. For the case of drywell leakage, the leakage becomes significant beyond 210 min-utes (approximately one hour after vessel failure) when the containment pressure exceeds 84 psia. A peak pressure of 104 psia occurs at about 275 minutes, at which time a maximum volume leakage rate of 1200 cubic feet per minute is predicted.

The wetwell leakage case exhibits similar behavior to that of the drywell leakage case. Because the wetwell leakage is smaller, a peak pressure of 118 psia is predicted at about 375 minutes. The maximum volume leakage rate

-in the wetwell leakage case is one order of magnitude smaller than that in the drywell leakage case. Comparisons of containment pressure and total volume of gas leaked are shown in Figures 3.5 and 3.6, respectively.

3.2.6 BWR Mark III Containment For the Mark III containment, the CLWG chose to concentrate on the TQUV sequence (loss of the power conversion system plus failure of all makeup water). The large Mark III containment can accommodate high gas generation rates but unlike the Mark I and II containments the Mark III is not inerted. This allows hydrogen burns to occur alcng with the associated high temperatures. However, explos'ive mixtures are avoided by the use of hydrogen control devices so that only local-

-ized burning (diffusion flames) is expected. The anticipated containment res-ponse is shown in Figure 2.13 for the no-leakage case (hydrogen burns have been suppressed).

The penetrations.

and CLWG has emphasized the local effects of the diffusion flames on equipment However, even without diffusion flames the drywell tempera-ture (refer to Section 2.2.6) will approach 600'F to 1000*F during core / concrete interactions. Under these severa conditions, the seals separating the drywell from the wetwell are expected to degrade over a period of about 5 to 6 hours6.944444e-5 days 0.00167 hours 9.920635e-6 weeks 2.283e-6 months after the start of the accident with a potential leak area of approximately 1 ft2 (refer to Appendix A). With this large bypass area, the gas generated by core / concrete interaction will not be sufficiently rapid to clear the downcomer vents. Thus, all gases and fission products generated 5 to 6 hours6.944444e-5 days 0.00167 hours 9.920635e-6 weeks 2.283e-6 months after the start of the accident are expected to pass freely to the wetwell airspace with-out pool scrubbing and condensation.

Under these circumstances the containment pressure rises more rapidly but the time to reach its capability pressure is NUREG-1037 3-18 '

. _ _ __ _ _. __ _ _ . _ _ _ _ ~ _ _ _ - . _

l L 5not greatly affected because'the containment had almost reached its capability pressure even without pool bypass. ,Thus,-the containment failure mode and timing is not' greatly affected by this potential pool bypass path. Bypassing the pool at this point in the accident will result in some enhanced fission product re-lease. This enhanced fission product release will be characteristic of the

'- i fission products released during core / concrete interactions. ;

'3.3 ' Impact ~on Radiological' Consequences During al core meltdown accident,. fission. products,may be released into the con-

~

tainment building and a number of systems are available to help contain or miti-gatei the radionuclide release. If these; systems fail or are compromised, a fraction of the radionuclides may be released to the atmosphere with corres-

,ponding adverse effects on the surrounding environment. These effects or con-a sequences can be measured in terms of health effects on the surrounding popula-

tion and property damage in.the surrounding area (damage indices). A number of these damage. indices (e.g., early. fatalities and injuries) are short term

, ~(within 2 months of the accident).and require the dose to exceed certain thres- i holds. 'For example, a dose of 320 rem to the blood-forming organs is required with supportive medical treatment (175 rem without treatment) to result in an early, fatality,in a small fraction of the population receiving such a dose.

' These damage.-indices.are therefore a strong function of the timing and magnitude of the/ fission. product. release-(for constant population and meteorology assump-tions) and also of the emergency response of the population. Other damage in-

' dices:such as latent fatalities and interdiction of crops or land are long term (50 years lafter'the accident) and measured over large distances (500 miles aroun'd the reactor site). These damage-indices are a function of the type and amount of

- radionuclidesnreleased and are relatively insensitive to the emergency response of the population. The implications of our work on containment performance for both short- and long-term damage indices are discussed for each of the six con-

tainment types in the following paragraphs.

4 4

We'havelnoted above that early damage indices require the release.cf significant quantities of fission products with relatively-little. warning so that' protective

, actions'are limited. If the quantity of fission products released is low, pro-tective action may not be so important or, if containment failure occurs with

-significant warning time, the population can be evacuated or sheltered and .thus theLthreshold doses necessary to cause early health effects can be avoided. In

.Section 2.2.1 we noted that for PWRs with large dry-containments the potential for.failing the containment close to the time of vessel failure was relatively low. .In addition, structural failure is predicted many hours after vessel-fail-

-ure'from gradual pressurization. The time between reactor vessel and contain-ment failure allows for aerosol agglomeration and settling to occur in contain-ment. 'Thus, the airborne fission product mass is much_ lower at'the time of.

+

structural failure. Consequently, the quantity of fission products released tfor. late containment failure times is significantly lower than if conta nment i failure occurs shortly after vessel failure. In addition, the long time between the start of core damage and the release of fission products to the atmosphere allows a substantial warning time necessary for evacuation or sheltering of the population. It was shown in the Zion Probabilistic Safety Study and in testi-l many at Indian Point hearings that ' late containment failure times were very effective in reducing risk of early fatalities.

3-The containment leakage estimates for a PWR with a large dry containment (refer to Section 3.1.1) indicate that containment leakage will be significant-only at NUREG-1037' 3-19 M , ~ . - _ . _ _ , - . _ . _ - - _ . - - - - _ , . _ - _ _ - . _ . _ _ . _ _ _ _ . . , _ _ . , _ _ . _ - _ _ _ _ _ , . _

/

relatively high containment pressures (120 psia). If the reactor cavity is dry and extensive core / concrete interactions occur, containment pressure will not be relatively high.(<100 psia) for.many hours after pressure vessel failure.

Therefore, on the basis of the Zion results, leakage during extensive core /

concrete interactions does not appea.- to be a problem for large dry contain-ments. This is an important conclusion because core / concrete interactions that persist for several hours can result in the release of a significant fraction of the refractory- fission products to containment. Consequently, if containment leakage had been predicted during the core / concrete interaction period;the potential-for the release of these fission products to the environment would have existed. However, as the predicted leakage at these pressures is' rela-tively low the. fission products should be well retained in containment. '

If the reactor cavity is flooded and a coolable debris bed is formed, th'e decay

heat boiling.of. cavity water will produce containment pressures greater than 120 psia (and hence increased leakage) about 8 hours9.259259e-5 days 0.00222 hours 1.322751e-5 weeks 3.044e-6 months after accident initiation.

However,1f~the reactor cavity is flooded and a coolable debris bed is formed, core / concrete interactions will be minimized. Hence,.the refractory fission products will not be released from the. core debris in the reactor cavity. This implies that:only fission products generated during the in-vessel melting phase will be released to containment. Consequently, aerosol agglomeration and

, settling will reduce the fission products suspended in the containment atmosphere before the beginning of significant leakage.

The leakage estimates for the PWR subatmoepheric and ice condenser containments were.found to be relatively small (refer to Sections 3.1.2 and 3.1.3 and would

not be expected to impact containment-failure modes or offsite' consequences.

-The containment leakage estimates for a BWR with a Mark I containment (refer to Section 3.1.4) indicate two potential leakage 4 paths from the drywell. Leakage through the drywell head appears to be a function.of pressure and will result i n ' leak areas sufficient to prevent further pressurization of the drywell late in the accident' sequence. The other leakage path, which was considered to be less. probable than drywell head leakage, results from failure of the purge and

-vent. seals from high temperatures early in the accident sequence. The purge and vent. seal leakage (although less probable) would potentially have the greatest impact on offsite consequences. This is simply because the. leakage areas are ' greater than for drywell head leakage. Leakage in Mark I containment can be-large and has both positive and negative effects. A negative effect is the potential'for significant release of the fission products earlier than would have been predicted based on threshold models. This earlier release when coupled with the potential for greater pool bypass (less fission product scrubbing) could result in increased offsite consequences. However, a positive effect of leakage vs. gross failure is that it could result in significant aerosol agglom-eration ard settling in the reactor building. In addition, the standby gas treatment system could be utilized to scrub the fission products under these circumstances. However, the above uncertainties associated with leakage vs.

- gross failure could be eliminated by the use of current operating procedures that call- for wetwell venting.

The containment leakage estimates for a BWR with a Mark II containment (refer to l Section 3.1.5) also indicate the potential for leakage at high pressures. Dry- '

well leakage in.the~ Mark II containment is as important as drywell leakage in the Mark I containment and has the same positive and negative effects as NUREG-1037 3-20 e - D-- -----q-m------et,-eab- n_ --.a.m--'--m--w- s-mww,eemp+-wcw-,as-s. --.vem-eeem r,ww.---w.%eemr g

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discussed above. However, if the wetwell is vented (modifiad BWR operating procedures), there will again be no driving force to produce significant drywell leakage in the Mark II design.

There are unique design features associated with the BWR Mark III containment that help to mitigate the consequences of severe core-melt accidents. The dry-well (which contains the primary system) is completely enclosed by the contain-ment (wetwell). In the Mark III design, containment failure (loss of containment integrity) would still leave the drywell integrity intact (containment function maintained) and the fission products would have to pass through the pool. This subjects the fission products to significant pool scrubbing, which greatly re-duces offsite consequences. Hence, drywell integrity is importart for minimiz-ing the offsite consequences of core melt accident sequences. Thus, the potential for flames in the containment (wetwell), resulting from in-vessel hydrogen pro-duction, to cause loss of drywell integrity was an important issue. Calculations (refer to Section 3.1.6) indicate that drywell integrity will be maintained even with these hydrogen diffusion flames in the wetwell. However, the drywell

- personnel access hatch has the potential to fail because high temperatures occur during core / concrete interactions. This failure provides a mechanism for pool bypass. However, the timing is expected to be late in the accident after most of the fission products have already been scrubbed by the pool. Hence, the impact on offsite consequences of loss-of-drywell integrity late in the accident sequence is not expected to be significant. The leakage estimate at 4 psid for the reference plant based on observed leakage testing is important because of the potential for suppression pool bypass early in the accident sequence and should therefore be carefully assessed in future studies. However, the leakage paths at this pressure would be tortuous with the potential for significant retention of aerosol fission products.

It should be noted that the above comments are qualitative and are intended to indicate potential trends. The leakage estimates in Section 3.2 must be incor-porated into containment failure mode t.nd fission product release analyses to determine the quantity and characteristics of the radionuclide release. Finally, it should be noted that the point of release of the radionuclides is important (auxiliary building vs. directly to the atmosphere) and will certainly influence the consequences. In summary, the above discussion clearly shows the importance of the present activities of the CPWG on the potential for #ission product re-lease from containment during core meltdown accidents.

4 NUREG-1037 3-21

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NUREG-1037 3-23

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Figure 3.5 BWR Mark II containment response NUREG-1037 3-26

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4 CONCLUSIONS The Containment Performance Working Group (CPWG) has completed studies of six reference plants that are considered representative of six different contain-l ment types (Zion /large dry PWR, Surry/subatmospheric PWR, Sequoyah/ ice conden-ser PWR, Peach Botton/ Mark I BWR, Limerick / Mark II BWR,' Grand Gulf / Mark III

.BWR). The CPWG concentrated its efforts on identifying upper bound leak area estimates that could occur before the containment shell capability pressure is I reached. -This pressure generally corresponds to the point at which the contain-ment reaches an initial general yield state. The CPWG believes that for the severe accident conditions described herein, the leak area estimates in this report will not be exceeded. On.the other hand, smaller leak areas may also be justified. The results presented should be utilized to assess the impact of

. containment leakage on the radiological consequences of an accident. However, until more test data are available, these results should be coupled with the results obtained from threshold models. The following paragraphs summarize the potential effects of the predicted upper bound leak area estimates.

The leakage estimated for the PWR large dry containment could delay or possibly prevent the containment from reaching its capability pressure. However, the leakage will not occur'until relatively late in the accident sequence so it is expected to have a minor influence on offsite consequences.

The leakage areas estir.ated for the PWR subatmospheric and ice condenser con-

.tainments were found to be relatively small, did not affect containment response, and were also expected to have little impact on offsite consequences. .

The leakage estimated for the BWR Mark I and II containments alters containment response, and its effect on offsite consequences should be carefully assessed.

However, it is also noted that changes to the BWR operating procedures involve the use of wetwell venting. If.the wetwell'is vented, there will be no driving force to produce significant drywell leakage (even with high drywell temperatures).

For BWR Mark-III containments, it is concluded that the drywell personnel air-

. locks will maintain their integrity even in the presence of hydrogen diffusion flames in..the wetwell. Because suppression pool bypass early in the accident is important, tests on drywell personnel airlocks should be conducted to confirm this finding. The drywell personnel airlock is predicted to result in signifi ,

cant bypass. leakage from high temperatures in the drywell during core / concrete

-interactions. However, this occurs late in the accident sequence and is.ex-pected to have little impact on offsite consequences.

It should be noted that the above comments are qualitative and are intended to indicate potential trends. The leakage estimates must be incorporated into con-tainment failure mode and fission product release analyses to determine the quan-tity.and characteristics of the radionuclide release. The point of release of the radionuclides is important (auxiliary building vs. directly to the atmos-phere) and will certainly influence offsite consequences. As discussed in Sec-tion 2.3.1, NRC-sponsored research on large-scale steel containment has indicated that containment functionality can be expected to be available for membrane NUREG-1037 4-1

i y

~ strains between 2% and.3% and that the containment capability pressure can be higher than.previously estimated. The capability pressure of a containment and its behavior under a severe accident challenge is very much dependent on plant-specific. design details, and have such important implications for offsite conse-quences that future studies on severe accident risk must incorporate realistic containment behaeior models.

In addition to the above findings, this study has led to the following general conclusions:

The potential for. containment' leakage through penetrations before reaching currently reported containment, capability pressures should be considered g- in severe' accident risk estimates.

The potential. for s'i gnificant leakage'before reaching currently reported containment capability press'ures appears to be greater for BWRs than for PWRs.

~

Leakage before reaching containment capability pressures can also occur p with PWRs, but such leakage is much more plant specific.

- Failures ~ of, nonmetallic seals for containment penetrations (primarily pressure unseating equipment hatches, personnel airlocks, drywell heads, and purge valves) are the most significant sources of containment leakage.

. Altho ~ ugh peneric studies of. containment types are useful in identifying 3

. sources of containment leakage, final. conclusions may need to be plant specific.

Current' efforts rely on analysis and engineering judgment. Additional test data are needed.to better quantify.the leak tightness of containment penetrations when subjected to severe accident conditions.

20n the basis of the results to date, both analytical and experimental studies

' should continue to better characterize containment leakage before reaching con-l tainment capability pressures as defined above.' Furthermore, efforts should P

-be made to better define the confidence levels associated with these capability

- pressures. Future studies should include the following:

l i ,

. Tests to fully assess the behavior of penetration seals under severe accident' pressure and temperature conditions, including the effects of-

[

aging and radiation.

I -

Sensitivity studies to assess the potential variation of containment

-leakage within the family of each containment type.

Sensitivity studies.to determini, the. magnitude and timing of containment L ' leakage that can.have a significant effect on radiological consequences.

. An assessment of the potential for plugging of identified leak paths.

An. assess' m ent of the survivability of' equipment inside containment during important severe accident sequences. <

k.

4

NUREG-1037 4-2 8

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F 5 REFERENCES Argonne National Laboratory (ANL), " Plan for a Penetration Integrity Program,

October 1982.

Battelle Memorial Institute (BMI), BMI-2104, "Radionuclide Release Under Specific LWR' Accident Conditions," J. A. Gieseke et al., Vols. I-VI, 1984.

Brookhaven National Laboratory (BNL), draft, "A Preliminary Review of the Limerick Generating Station Severe Accident Risk Assessment," Vol.1, Core Melt Frequency, M. A. Azarm et al., August 15, 1983.

-- , BNL-NUREG-33835, " Containment Failure Mode and Fission Product Release Analysis for the Limerick Generating Station: Base Case Assessment," H. Ludewig et al., March 1984.

-- ,' Technical Report A-3706, " Containment Loading for Severe Accidents in BWRs with a Mark I Containment," K. R. Perkins, W. T. Pratt, and G. A. Greene, November 1984.

-- , Technical Report A-3736R, "The Influence of Leakage on Containment Perform-ance During a Core Meltdown Accident," R. Juang et al., February 28, 1985.

Clow Corporation, Report No. 11-15-83, " Purge and Vent Valve Operability Quali-fication Analysis," 5. M. Mondahl and J. E. Krueger, November 16, 1983.

Commonwealth Edison Company (Ceco), " Zion Probabilistic Safety Study," Docket 50-295/304, September 1981.

Cybulskis, P., R. Wooton, and J. Kolb, " LWR Core Meltdown Accident Sequence Phenomenology," Trans. Am. Nuc; Soc., Vol. 41, pp. 407-410, 1982.

EG&G Idaho,'Inc., TREE-1979, " CONTEMPT-LT-028 - A Computer Program for Predict-ing Containment Pressure-Temperature Response to a Loss-of-Coolant Accident,"

.D. W. Hargroves, March 1979.

-General Electric Co. (GE), "BWR-6 Standard Plant Proiabilistic Risk Assessment, 238 Nuclear Island, GESSAR-II," Rev. 2, Docket 50-447, February 1982.

l Letter, October 7,1983, R. L. Baer (NRC) to F:. C. Lewis (NRC), C. E. Norelius (NRC) and J. E. Gagliardo (NRC), " Potential Problem on Airlock Seals in Contain-ments and Drywells."

-- , June 6, 1984, B. E. Miller (BNl) '.v J. . .cra (NRC), " Scoping Test on Containment Purge and Vent Valve SeC MM erial," draft report, May 1984.

-- , June 29, 1984, B. E. Miller (BNL) to V. Noonan (NRC), "An Estimation of Pre-Existing Containment Leakage Areas and Purge and Vent Valve Leakage Areas Resulting from Severe Accident Conditions."

NUREG-1037 5-1

l l

Memorandum, February 15, 1984, M. Silberberg (NRC) to CLWG members, "Contain-ment Loads Working Group Meeting at the Brookhaven National Laboratory, Septem-ber 15-16, 1983."

-- , December 4, 1984, J..F. Costello (NRC), to Distribution (H. Denton et al.)

" Quick-Look Report on Large Steel Model Test."

Northeast Utilities, " Millstone Unit 3 Probabilistic Safety Study," Docket 50-423, August 1983. .

NUREG reports, see U.S. Nuclear Regulatory Commission.

Parker Seals Company,- 0-Ring Handbook, OR5700, Lexington, Kentucky,1977.

Philadelphia Electric Company, " Limerick Generating Station - Probabilistic Risk Assessment," Docket 50-352/353, March 1981.

-- , " Limerick Generating Station - Probabilistic Risk Assessment," Docket 50-352/353, April 1983.

Power Authority of the State of New York (PASNY) and Consolidated Edison Company (Con Ed), " Indian Point Probabilistic Safety Study," Docket 50-286, March 1982.

Rashid, Y. R., " Review of WASH-1400 Surry Containment Capability Assessment,"

prepared for EPRI Project 2288-2, February 1983.

Riskalla, S., S. H. Simmons, and J. B. MacGregor, "A Test of a Model of a Thin-Walled Prestressed Concrete Secondary Containment Structure," Transactions of the 5th International Conference on Structural Mechanics in Reactor Technology, North Holland Publishing Co., New York, 1979.

Sandia National Laboratories (SNL), SAND 77-0370, "A Preliminary Model for Core /

Concrete Interactions," W. B._Murfin, August 1977.

-- , SAND 80-2415, "CORCON-M001: An Improved Model for Molten Core-Concrete Interactions," J.~R. Muir et al., 1981.

Stone and Webster Engineering Corp. (SWECO),. Report TP-84-13, " Containment Integrity at Surry Nuclear Power Station,"_ February 1984.

Technology for Energy Corp. (TEC), IDCOR Technical Report 10.1, " Containment Structural Capability of Light Water' Nuclear Power Plants," July 1983.

U.S. Nuclear Regulatory Commission (NRC), NUREG-0850, Vol.1, " Preliminary Assessment of Core Melt Accidents at the Zion and Indian Point Nuclear Power Plants ' and Strategies for Mitigating Their Effects," November 1981.

-- , NUREG-0956, "A Reassessment of Accident Source Term Evaluation Methods and Assumptions" (to be published).

p NUREG-1037 5-2

k 1-- ,LNUREG-0974, " Draft Environmental Statement Related to the Operation of r- Limerick Generating Station, Units 1 and 2," June 1983; " Final Environmental ,

-Statement Related to the Operation of Limerick Generating Station, Units 1 and '

2," April'1984; Docket Nos.~50-352 and 50-353.

-- , NUREG-0979, " Safety Evaluation Report-Related to the Final Design Approval of the GESSAR II, BWR-6 Nuclear Island Design," April 1983; Supp. No. 1, July 1983; Supp. No. 2, November-1984.

-- , NUREG-1079, " Containment Loads Working Group Report" (to be published).

F

m. - , NUREG-75/014 (formerly WASH-1400)," Reactor Safety Study: -An Assessment

of ' Accident' Risk in U.S.- Commercial Nuclear Power Plants," December 1975.

, -- , NUREG/CP-0033, Proceedings of the Workshop'on Containment Integrity, Vol. I of II, pp. 45-62, " Resolution of Containment Structural Response Issues Under Degraded Core Conditions," T. J. Marcinak, E. P. Stroupe, and-

'S. H.~Fistedis, October 1982.

I '

., NUREG/CR-1185, " COMPARE-MOD 1 Code Addendum, R. G. Gido et al., August 1979.'

i

-- , NUREG/CR-1659, "Reactory Safety Study Methodology-Applications Program:

S:quoyah 1 PWR Power Plant," D. Carlson et al. , February 1983.

.' -- , NUREG/CR-1711, "Battelle Columb'us Laboratory MARCH Code Description and Users 1

- Manual," R. O. Wooten and H. I. Avci, October 1980.

. -- , NUREG/CR-1891, " Reliability Analysis of Containment Strength - Sequoyah y and McGuire Ice-Condenser Containments, L. G. Greimann et al., August 1982.

. -- , NUREG/CR-1967, " Failure Evaluation of a Reinforced Concrete Mark III Con-

- toineent Structure Under Uniform Pressure," 5. Sharma et al., September 1982. ,

1- -- , NUREG/CR-2182,'" Station Blackout at Brcwn's Ferry Unit One - Accident' S quence Analysis," D. H.. Cook et al., November 1981. .

-- , NUREG/CR-2442, " Reliability Analysis of Steel Containment Strength,"

L R.'G. Greimann et al., June 1982. ,

- -- , NUREG/CR-2825, "BWR/ Mark I Accident Sequence Assessment," D. D. Yue and j

- T. E. Cole, November 1982.

(' -- , NUREG/CR-2973, " Loss of DHR Sequence at Brown's Ferry Unit One - Accident

! S:quence Analysis," D. H. Cook et al., May 1983.

v

, NUREG/CR-3028, " Review of the Limer

'erating Station Probabilistic

' Risk Assessment,". I. A. Papazoglow et al. dary 1983.

t

. -- ,=NUREG/CR-3179, Appendix B, " MARCH 1.1 Code Improvements for BWR Degraded Core Studies," R. M. Harrington and L. J. Ott, October 1933.

U

-- , NUREG/CR-3278, " Hydrogen Burn Analysis of Ice-Cordenser Containments,"

R; G. Gido and A. Koestel, April 1983.

- NUREG-1037- 5-3 4

, . . , . - ,-e.,,n-- - - , - - - , , --.w,, .-w--a- - w ,,-.-+------,-,,--.,am, - , . - , - ,-+m --- -- - -,w-,- - - . . , .e e -

-- , NUREG/CR-3300, Vols. 1 and 2, Draft, " Review and Evaluation of the Zion Probabilistic Safety Study," July 1983.

-- , NUREG/CR-3952, "Sequoyah Equipment Hatch Seal Leakage," L. G. Greimann et al.,

August 1984. ,

-- , NUREG/CR-4135, " Review of the BWR-6 Standard Plant Probabilistic Risk Assessment: Internal Events, Core Damage Frequency," BNL-NUREG-51852 (to be published).

-- , NUREG/CR-4143, " Review and Evaluation of the Millstone Unit 3 Proba-bilistic Safety Study: Containment Failure Modes, Radiological Source Terms and Offsite Consequences," M. Khatib-Rahbar et al. (to be published).

-- , Testimony for the Indian Point Investigative Hearing, Docket 50-247/286, pending.

WASH-1400, see U.S. Nuclear Regulatory Commission, NUREG-75/014.

Weinstein, M. B., " Primary Containment Leakage Integrity: Available and Review of Failure Experience," Nuclear Safety, Vol. 21, No. 5, September-October 1980.

NUREG-1037 5-4

V -

APPENDIX A

-EFFECT OF'NON-UNIFORM THERMAL ENVIRONMENT ON DEGRADATION OF CONTAINMENT PENETRATIONS DURING SEVERE ACCIDENTS 11.

NUREG-1037

EFFECT OF NON-UNIFORM THERMAL ENVIRONMENT ON DEGRADATION OF CONTAINMENT PENETRATIONS DURING SEVERE ACCIDENTS Introduction In the analyses of postaccident conditions, the survival of containment penetra-tions is evaluated for the average environment which is predicted to exist in the containment at different times after an accident. However, in many cases some phenomena may occur which would produce local releases of thermal energy and, as a result, some of the penetrations may be exposed to considerably higher local heat fluxes than are predicted to exist in the containment atmosphere.

The present study includes the pressurized-water reactors (PWRs) with large dry containments (Zion) and boiling-water reactors (BWRs) with Mark III containments (Grand Gulf).

Sources of Non-Uniformity of Thermal Environment (1) Non-Uniformity Attributable to High Temperature Steam Release Normally, steam released from the primary system will diffuse throughout containment and it is not expected that any localized high temperature regions will exist. However, if a containment penetration happens to be directly exposed to a jet of hot steam it may reach considerably higher temperatures than those existing elsewhere in the containment. During an accident, steam can be released to the containment either through a break in the primary system piping or through safety and relief valves. During most transients only the latter release can occur and, since steam from these valves is released to the pressurizer relief tank, no direct expo-sure of containment penetrations to hot steam jets is possible. In some instances, however, if sufficiently large amounts of steam are released, the temperatures in containment may reach fairly high levels.

(2) Non-Uniformity Attributable to Burn of Combustible Gases During a severe accident, combustible gases can be generated within the containment. These gases are hydrogea and carbon monoxide. The main source of hydrogen is reaction of steam with zirconium fuel cladding and to a smaller extent with the reactor's stainless steel structural compo-nents. Some hydrogen will also generate from radiolytic decomposition of water. The amount of hydrogen generated during a transient in PWR and BWR plants is predicted to be 2000 to 3000 lb (NRC, NUREG/CR-2228 and NUREG/CR-2530). The source of carbon monoxide is a reaction of melted core with concrete. Carbon monoxide will generate after the core melts through the bottom of the reactor vessel and slumps into the containment sump. The amount of carbon monoxide generated during TMLB' transient is estimated at 30,000 lb (NRC, NUREG/CR-2228).

The hydrogen generated in the primary coolant system can be either released to the containment where it nixes with air and the mixture is burned with NUREG-1037 A-1

4 a flame propagating throughout the containment, or it can burn at a point of release. -Hydrogen can be released locally either directly from the break in the primary coolant system or, in the case of BWRs with Mark III containments, from the surface of the suppression pool. In this case the steam-hydrogen mixture released to'the drywell is forced to pass through the suppression pool water where steam condenses leaving pure hydrogen which is then released from the surface of the pool. Hydrogen released from this source burns with a stationary (diffusion) flame.

~During the premixed burning, the hydrogen flame itself does not produce sig-nificant thermal flux, but the resulting hot gases emit large amounts of

-heat by radiation and convection. Since a hydrogen-air mixture can be

-ignited at any location in the containment and since the resulting flame

'will-travel throughout the containment, there is always a possibility that any containment penetration could be exposed to high thermal fluxes. How-ever, the flame moves at a relatively high velocity (on the order of

. 2-10 ft/sec) and the hot gases behind the flame front cool rapidly, so that the exposure of penetrations to high heat fluxes is rather short. Sta-tionary flame, on the other hand, releases its thermal energy locally and hence the objects located in the immediate vicinity of the flame are sub-jected to a severe environment. The presence of carbon monoxide does not change these mechanisms significantly and it can be considered as an addi-tional source of. thermal energy similar to hydrogen, but of lesser magnitude.

From the above discussion it can be concluded that for the PWRs and BWRs, the only significant type of non-uniform thermal environment to which containment penetrations could be exposed is caused by hydrogen and carbon monoxide burn with. flame propagating close to the penetrations.

In:the BNL analysis of containment response in a PWR plant during degraded core accidents (HRC, NUREG/CR-2228), hydrogen burn is predicted for the TMLB' tran-sient, with containment spray during injection and recirculation phases. The maximum temperature reached is approximately 900 F and this temperature excur-sion lasts only a very short time. The' analysis ignores the presence of carbon monoxide. In order to correct for this omission, the length of temperature peak should be' increased. It can, however, be demonstrated that the maximum tempera-ture would not change significantly because of considerably lower heat of com-bustion of carbon monoxide. The BNL analysis (NRC, NUREG/CR-2228) predicts that

the combustible gas starts burning at about 850 minutes after the beginning of the transient. .This is about 550 minutes'after the 38 psi containment pressure

~

rise that results.from a large amount of steam generated by interaction of melted core with sump water.

The analysis of containment response in a BWR plant with Mark III containment predicts hydrogen release to the drywell in about 100 minutes after the begin-ning of'the accident. It is assumed that this hydrogen passes through the sup-pression pool and is' released to the wetwell. It is then ignited at the surface

g. of the pool and burns for 50 minutes with a stationary diffusion flame. The

' heat fluxes from this flame to outside containment and drywell walls were deter 3 mined by Sandia using the HECTR computer code. The temperature in the drywell, after remaining for about 200 minutes at a 130*F to 240 F. level, rises very l rapidly, reaches approximately 960 F, and remains at this level for about

.> 600 minutes. The corresponding containment pressure determined by the analysis varies between atmospheric at the beginning of the accident and about 80 psia toward its end.

NUREG-1037- A-2

- v- -- - - - - ..-m w., -

>,y , , , , , -- nm ,

r Characterization of Containment Penetrations The effect produced by the_non-uniformity of thermal environment on degradation of containment penetrations will depend on their locations, thermal characteris-tics, and materials of construction. In the case of premixed burning, the non-uniformity is caused solely by the energy released from burning gases and since it can occur at practically any location in the containment, nearly all penetra-tions have equal probability to be exposed to the burning gas environment. In the case of diffusion flame, only the penetrations located close to the flame will be exposed to high thermal energy fluxes. Thermal response of penetrations is determined by thermal capacity (proportional to the mass of penetration),

geometry, and the nature of material. This last property is defined by thermal diffusivity k ,

" . _~ cp Materials of construction of penetrations will also determine the degree of degradation caused by a given thermal response. Thermal degradation of seals and gaskets made of organic materials is the main cause for change in leak characteristics.

The following five categories of containment penetrations exist in PWR and BWR plants (Ceco, MP&L Co.).

- electrical penetrations

- piping penetrations

- equipment and personnel hatch

- fuel transfer and sump penetrations

- purge and vent valves Their thermal characteristics are described below:

(1) Electrical Penetrations A typical electrical penetration consists of a metal pipe passing through the containment wall. Inside this pipe there is a stainless steel canister contain-ing electrical cables. The canister is welded on inside and outside ends to the pipe. The welds form an hermetic barrier. The cables are passed through the stainless steel header ~ plates located on both sides of the canister. The electrical conductors passing through these plates are sealed. These seals form critical elements of the penetration. Only one side of the penetration would be exposed to burning gases. The projecting portion of the pipe and the canister could pick up a significant amount of thermal energy by radiation and convection and its temperatures may rise signifcantly. However, even if this temperature exceeds the upper safety limit and the inboard seals fail on the inboard side of the penetration, the outboard seals will protect the penetration from leakage.

As discussed in draft NUREG/CR-3234 (NRC), thermal energy transfer through the canister is very low and outboard seals will never reach high temperature.

(2) Piping Penetrations Piping penetrations for both hot and cold pipes do not contain any material which could be degraded at high temperatures; hence, even if they are exposed to high heat fluxes they would not change their leak characteristics.

NUREG-1037 A-3

(3) Equipment and Personnel Hatch An equipment and personnel hatch is a bulky structure with a very high thermal inertia. .It is not expected, therefore, that in most cases its temperature would increase significantly if exposed to heat fluxes from burning gases.

This was verified by the thermal response analysis presented in the next sec-tion. However, when the part of a hatch containing seals is exposed directly to the flame, gasket material may become overheated and may degrade. Personnel airlocks are also bulky structures with large thermal inertia. Furthermore, each airlock has double doors so in PWRs only he inboard door is exposed to the thermal environment inside the contains a .

In BWRs, however, personnel airlocks in the drywell wall may become exposed on both sides to high energy fluxes and under certain circumstances the seals on drywell and wetwell sides may fail, resulting in the development of a leak path between the drywell and the wetwell.

(4) Fuel Transfer and Sump Penetrations Because of their locations within the containment it is very unlikely that these penetrations will be exposed to the thermal fluxes from burning gases.

(5) Purge and Vent Valves The valves on the purge exhaust and supply system are heavy, hence they have large thermal inertia and can absorb relatively large amounts of thermal energy without a significant change in their temperatures. In addition, each of the penetrations is provided with two valves in series: one on the inside and one on the outside of the containment wall. Only the valves inside containment will be exposed to high heat fluxes from burning gases and the outside valves would not be affected. Leak characteristics of the purge system will be therefore preserved even if the inside containment valves fail. Thermal response analy-sis for these valves is presented in the following section.

Thermal Response of Penetrations Exposed to Non-Uniform Thermal Environment Thermal models of containment penetrations were developed and temperature rises of the most sensitive components of these penetrations were determined. Two types of models were developed: one for the equipment hatches and purge' valves in PWRs and BWRs exposed to a premixed hydrogen burn and the other for the per-sonnel airlock in the drywell wall of BWRs with Mark III containments exposed to a standing hydrogen flame. Thermal environment in the second case is much more severe since both sides of the personnel airlock are exposed to elevated tem-peratures for an extended period of time.

(1) Environment of Premixed Hydrogen Burn (a) Equipment Hatch A thermal model of the equipment hatch was developed from the drawing presented in Figure A.1.

The most thermally sensitive part of the equipment hatch is the cover seal made out of organic material. Temperature rise of this seal was calculated.

NUREG-1037 A-4

V-In setting the model the.following assumptions were made:

: -Equipment hatch is symmetrical and as a first approximation a one-dimensional model'can be used.

Temperatures (f containment walls do not change during hydrogen burn and remain equal to the initial temperature.

Initial temperature of the equipment hatch cover is equal to the inside

' containment temperature just before hydrogen burn.

Material with homogeneous thermal properties is assumed (steel). The effect produced by the presence of the seal is neglected.

With these assumptions the mathematical model, Figure A.2, is developed.

The model in Figure A.2 represents a cross-section through the hatch door and door frame which contain the seal made out of organic material.

In_this model Regions 1 and 2. represent the hatch door and Region 3 the frame to which the door is attached. .

Region 1 Making balance for finite element in Region 1:

-(Thermal Energy Absorbed) = (Incoming Thermal Energy from Outside) - (Outgoing Thermal Energy to Outside) + (Thermal Energy Conducted to Element) - (Thermal Energy. Conducted out of Element)

'2n. r Ar S p c At = 2n r n

^#' *i A0 - 2n rn* 0# * *o A0 n n n tn ) k(tn - tyy)

+2n(rf-Ar)k(t or S

A0 - 2n(rn +0

Ar Simplifying: ..

A'tn * " (Ar)Z ,

k +(t n-1 - 2t n + tn+1) - (t n-1 -t+1) n where:

k' a= cp thermal diffusivity (1) substituting for:

At = t' - t n n

$j =h g (tc -t) n to "N o (t -t) n o NUREG-1037 A-5

7 L

' assuming:

Ar. r , i.e.,

n 20 n .

.and rearranging the_ terms:

ty=B tn-1 + t n+1 +

~ (2 + N) t n

+R (2) where:

B=a hep (3)

N _ (h9 k o

+ h(ar)2

)S (4)

(hgtc + h f_ td) (ar)2 k 5 (5)

Region 2 (with' adiabatic surface)

~In Region'2 -

tn-1 = tn =tt and ti='B t2+tt - (1 + N) +R (6)

Region 3 (with heat sink)

In Region 3 heat is transferred to the heat sink which remains at the temperature t,

. . I t n+1 = to andtg=Bj tn-1 + tgi

+t n H - (2 + N)

+R ] (7)

Evaluation of B, N, and R parameters in equations 2, 6, and 7 (i) B (dimensionless) as B=a p. 2 For carbon steel:

k'= 26 Btu /hr ft2 op c = 0.12 Btu /lb *F p = 487 lb/ft8 a = 0.445 ft2 /hr NUREG-1037 A-6

Ff _17 - - .

ItIis-assumed that M = 1 in.

~B =.(64.08)(as) (8)

(ii)'N (dimensionless)

-y, i bk o) , (ar)2

.S where:

'hj .='hRi + hci>

1 '

(t c+460)4- (t + 460)4

.hRi.= oF 1/c)+1/c -

g (tc- t) assuming:

e = 0.35 (h'ot; gas) g "c = 0.8 (steel)

F = 1.0-o = (0.-1713)(10 s) Btu /hr ft2 (oR)*

h ci=0.19(tc -t)II3

. (t'+460)4-c (t + 460)4 hRi = (5.513)(10 M) t-t C ,,

hRi, hci, hRo, and h co are evaluated using one sis.31 e value of solid surface temperature in the middle of the area considered.

. h, = hRo + h co 7 *'

(t+460)4-(t,+460)4 hRo = oF (1/c + 1/c - 1) t-t,

' (t + 460)4'- (t+460)4 '

hRo^= (1.142)(10 9)

(g_g) h,=0.19(t-t,)1/3 o ,

Thickness of hatch cover is S = 1.5 in. and N corresponding to hatch cover is N=_(2.13675)(103)(h j + h,

-NUREG-1037 A-7

t Thickness of hatch frame is S = 3.25 in. and N corresponding to hatch frame is N=(2.13675)(103)(h9+h)(1.5/3.25)=(9.8620)(104)(h9+h) o g T(iii) R(*F)

R = (hg t+h c g to ) gp)2 k S R is calculated'similarly to N:

For. hatch cover:

R=(2.13675)(103)(h9t c + h, t, }

For hatch frame:

R=(9.8620)(104)(h.t+h Results 9 c o t) o Equations 2, 6, and 7 were evaluated for the conditions shown in Figure A.3.

tc.(max) = 900 F tc (min) = 240 F O =.120sec(durationoftemperaturespikE) sp t, = 240 F (initial temperature)

Final temperature of the seal is 242.3*F, temperature increase is 2.3F*.

.If more than one hydrogen burn is assumed, the corresponding temperature increase can be determined by multiplying 2.3F by the number of burns.

(b) Purge Valves

. The most thermally sensitive part of a purge valve is the gasket attached to the valve disc. Thermal modeling (Figure A.4) was made using the following assumptions.

Purge valves are located far enough from the containment walls so that thermal effect of the walls can be neglected.

Purge valve is attached to the ducts which shield the valve disc and the gasket from the directed thermal effects of burning gases.

Purge valve heats uniformly. This is a very conservative assumption because in reality the outside of the valve will reach higher temperatures !

than the disc which is located inside the body of the valve. ;

- i Initial. temperature of the purge valve is equal to the inside containment temperature just before hydrogen burn.

4 NUREG-1037 A-8

J i

1+ Material with homogeneous thermal properties is assumed.

Using.these assumptions,'the energy balance for the entire valve is:

l V'c ph=hg A(tc - t) (9) ].

Integrating over a small time increment A0 and assuming constant h g .

t 'dt Ah g d3 f't t e

-t -f0+A0 6

Vcp t' = t-exp[-Ka0] + tc(1 - exp[-KA0]) (10)

.where: _

h K=p A

cg p hy=hRi + hei

- The values of h Ri and h ci are the same as in the analysis of equipment hatch.

Evaluation of K Volume of valve material:

V = 10.38 ft 3-Areas exposed to energy transfer from environment:

28.45 ft2 K=

2( =(4.691)(102)(hRi + bci)

.Results Equation 10 was evaluated for the same conditions as the equipment hatch, using a computer program.

Final temperature of the gasket is 241.7*F.

Temperature increase is 1.7F*.

Similarly to the equipment hatch analysis, for multiple hydrogen burn tempera-ture increase'is determined by multiplying 1.7F* by the number of burns.

(2) Environment of Standina Hydrogen Flame Personnel airlock in the drywell is exposed on one side to a standing hydrogen flame-and on the other to high temperature steam in the drywell. Since both NUREG-1037 A-9

l l

these thermal energy sources contribute to heating up of personnel airlock, both of them participate in raising the seal temperature. Temperature time histories l in the wetwell and drywell are shown in Figure 2.13 in the main body of this report.

In addition to temperature rises, there is also a corresponding in-crease of pressure in the containment (see Figure 2.13). The information presented in these curves is used as an input to the calculational procedure used in determining thermal response of different elements of the personnel airlock.

Thermal model for the personnel airlock was developed from the corresponding technical drawings. A schematic sketch of the personnel airlock is shown in Figure f 5.

Therm the most sensitive parts of the personnel airlock are rubber seals in the L m both wetwell and drywell sides. The temperatures reached by these seals during a severe accident were calculated.

In developing the model the following assumptions were made:

Wetwell side wall in the personnel airlock is circular and the model developed for the equipment hatch applies (equation 1).

Drywell side wall of the personnel airlock is assumed to be uniformly heated.

Temperature of the drywell wall changes during the accident, but it remains uniform within the wall.

There is a perfect mixing of air inside the personnel airlock.

Inflatable rubber seals fail at 600 F.

When the seal fails, the ingress of steam to the personnel airlock is con-trolled only by pressure difference. No resistance to the flow of steam throughout the damaged seal is considered.

In thermal analyses the following three distinct intervals are considered:

standing hydrogen flame in wetwell - 50 minutes low temperature in drywell - 30 minutes high temperature in drywell - 600 minutes Thermal responses of personnel airlock seals in each individual interval are determined.

(a) Drywell Wall Temperatures of the drywell wall were determined by solving analytically the differential equations obtained from energy balance:

i W p c dty =A($wD+#DD)do l Substituting for $wD and $0D the above equation can be solved for different intervals:

NUREG-1037 A-10

.r General format equation is:

=t y = G[1 exp(-H 0)] + I exp(-H 0) - J -0 (11)

The values for parameters G, H, I, and J are as follows:

Interval G H I J a 306 2.874 t yg 1.1 y <t 219 2.906 t 0 Dsat y b

(tty2tDsat 193 0.0718- ty 0 t

y <t 0 sat 298 2.906 t y 0 (ty1tDsat 807 0.169 ty 0 l

-(b) Personnel Airlock's Wall Facing Wetwell-The. mathematical model used for calculating seal temperatures of this portion of a personnel airlock is identical to the model used in ' calculating temperatures

.in.theequipmenthatch,exceptthatinthiscasefhasafinitevalue n

'ar =n rn Substitutingforfandforheatfluxesinequation1andrearrangingthe n

terms,theexpressionsfor(tsw)frdifferentregionsbecomeasfollows:

Region 1

.(tsw n = Bf1 - f[tsw)n-1+(1+ t (12) sw +1 + -(2+N)(tsw)n + Rf

, where:

.B. =a Asp

-For interval a:

0.2$wD+h(t p g -t)(ar)2y (13) k(tg - t y}S >

(tg- 0.8 ty ) $wD + hp tp(tg -t)(ar)2 y (14) k(tg - t y)S The values for $wD and t gare determined by Sandia's HECTR code.

i NUREG-1037 A-11 t

For intervals b and c:

{ N= " ( ")

k (15)

+

4 R _ (h, J t G) + (h, t p) (ap)2

. k S (16)

, Region 2 (with adiabatic surface)

In Region.2 n = 1 and-

-(t',{=Bf2t2+t2 - (2 + N) ,

+ Rf (17) 1 Region 3 (with heat sink)

~

In Region 3 heat is transferred to or from the drywell wall which can become either a heat sink or.a heat source

{tsw)n+1=ty t =8 sw)n 1 - f)(tsw)n-1+(1)t, +

(2+N)(tsw)n

+

Rf (18)

. Total number of nodes (n) in the personnel airlock wall is 57.

The: seal is located at node 10.

.(c) Personnel Airlock's Wall Facina Drywell (Figure A.6)

This wall is composed of two parts: a circular end plate containing a door and a cylindrical portion. In developing the mathematical model it is assumed that the in thecircular end plate cylindrical is uniformly heated and temperature gradients exist only part. Similarly to the model of the wall facing wetwell, the model for the wall facing drywell includes three separate regions for which different mathematical: expressions were developed.

The general form of.these expressions is similar to those for the equipment hatch described earlier in this report (equation 1).

, Region 1

~

(tsDn=B(tsDn-1+(tsD)n+1+ - (2 + N) (tsD)n+Rf (19)

I where:

B=a h (20)

NUREG-1037 A-12 o

a

N= 0+hf p (21) cyl.

(hD ~ t D) + (hp . ' t )

2 R= (22) cyl Region ~2 (end plate)

In Region _2)

(ts0n-1=(ts0n=(t sD 1 (tsD 1 = B t sD2'+..-(1+N')(tsD1+R'f (23)'

where:

cyl

A*

B'=B(2Sep.r S a- (24)

(25)

N'=Nh).

(26)

R'=R(h)

Region'3 (next to drywell wall) lIn Region 3

.(tsD)n+1=ty

+

(t'D)n=Bf(t')n-1+t sD y - (2 + N) (tsDn+Rf J27)

Total number of nodes.(n) in the personnel airlock wall exposed to drywell is

11. The seal is located in the first end plate node.

(d) Air in Personnel Airlock When all the seals retain their integrity, the air inside the personne1' airlock

~

, is heated only by the heat transferred through the personnel airlock walls. How-ever,'when one of the seals fails, steam and hot' gases from the containment enter the personnel airlock and bring with them additional thermal energy. It is assumed that when a seal fails, gases are admitted to the personnel airlock Linstantaneously until pressures equalize.

Scyl = 0.5 in.. (thickness of cylindrical portion); S,p = 1 in. (thickness of end plate portion).

NUREG-1037 A-13 4

k Temperature in the' personnel airlock is determined by makirg energy balance:

(Thermal Energy Content of Personnel Airlock Gas) = (Initial Thermal Energy Content of Personnel Airlocks Gases) + (Energy Transferred Through the Walls)

+ (Thermal Energy of Gases Coming in Through Damaged Seal) ps("a c, + ms cs ) * "a c, tp+ E ($

t a) 40 + m3 c t (28) s 0 The value of ms is determined from the gas laws:

~

V P P s D pf

,s = D (29)

(460+tps) (460+tpf}

and the value of heat fluxes is given,by the following expression:

$=h(t p s -t,p (30)

Temperature of gases in the personnel airlock can be determined at any time in the accident by equations 28-30.

The value of PD is obtained " rom Figure 2.13 in the main body of the report.

Evaluation of Parameters in Equations 11 through 30 (1) Geometrical Parameters x = 10 ft W = 3.5 ft '

r = 5 ft Ax = 1 ft ar = 1 in.

S,p = 1 in.

S l cyl = 0.5 in.

V' = 1178 ft3 (2) Temperatures and Pressures Initial Pressure of Personnel Airlock Gas: 14.7 psia Initial Temperatures 'in the Containment: 160*F Wetwell Gas Ten,perature during Hydrogen Burn (calculated by HECTR Code):

tg = 307 + 87.130 *F' Seal Failure Temperature: 600 F NUREG-1037 A-14

J

'(3) Heat Fluxes and Heat Transfer Coefficients WETWELL Heat Flux from Hydrogen Flame (calculated by HECTR Code):

$g = 1046 - 1920 Btu /hr ft2 h, = h +h Btu /hr ft2 op ew Rw Convection Heat Transfer:

he ,= 0.19 (t sw -t g )1/3 ~Btu /hr ft2 op

- Radiation Heat Transfer:

hgg = (0.1713)(10 s) e (t sw +460)4-(tg+460)4 BW/h R2 op 8'

1/c + /c - 1 c = 0.8 DRYWELL Interval a, Condensation Heat Transfer:

hD = 175 Btu /hr ft2 op Intervals b and c:

For tsD < tD sat Condensation Heat Transfer:

h '= 175 Btu /hr ft2 op D t For tsD > tD sat hD=bcD + hRD Btu /hr ft

F Convection Heat Transfer:

hcD = 0.19 (ts0 -t)/D Btu /hr ft2 op Radiation Heat Transfer:

(ts0+460)4 (t0+460)4 hRD = (0.1713)(10 8) e (tsD -t) D NUREG-1037 A-15

PERSONNEL AIRLOCK For total heat transfer to personnel airlock without steam h

p

=0.19(t -t)1/3+[(0.1713)(108)]c(t+460)4.(t s p s p

+460 For total heat transfer to personnel airlock with steam

h p =0.19(t s -t ps)1/3+[(0.1713)(108)3e '( ts +460)4-(t s +460)4' (4) Personnel Airlock Temperature Substituting to equations 28 and 29 c, = 0.24 Btu /lb 'F cs= 0.5 Btu /lb "F m, = 70.68 lb tp = 17.03 t, + Q + 0.5 tD * *s - and 17.03 + 0.5 m s

P P D pf

s = 1976 (460+tps) (460 + tpf)

Results ~

Temperatures tures. reach highest of seals are determined for the locations at which these tempera-values.

For the seals in the door facing wetwell it is 10 in~. from the center of theendplate:{tsw)to.

For uniform the temperature door facing the t drywell it is assumed that the whole seal remains at sD*

.These a specially seal developed temperaturescomputerwere determined by solving equations 11 through 30 u code.

The temperatures were determined for the wetwell and drywell conditions specified in Figure 2.13 in the main body of the

. text and. assuming standing flame in the wetwell lasting 50. minutes.

The determination was made for two cases:

(a) assuming no seal failure (b) assuming drywell seal failure at 600*F The'results are shown in Figures A.7 and A.8 for cases (a) and (b), respectively. '

NUREG-1037 A-16 9

The highest temperatures reached by the wetwell and drywell seals are tabulated below:'

Highest seal temperature, *F Case Interval Wetwell Drywell '

Seals survive Hydrogen burn 345 264-High drywell 531 908 temperature I Seals fail at 600*F Hydrogen burn- 345 264 High drywell 557 873 l 4

temperature Conclusions

.The review of thermal characteristics of containment penetrations in PWR plants

.has indicated that the only non-uniformity of thermal environment to which these penetrations would be exposed can be caused by burning of flammable gases (hydro-

' gen and' carbon monoxide) generated in the plant during the transient. These burns may cause local hot spots as a result of high energy release. These localized energy releases will'be of short duration and only the components with relatively small thermal inertia will reach high temperatures. Thermal responses were analytically determined for two components most'likely to fail when exposed to elevated temperatures: the seal in the equipment hatch and the gasket in the purge valve. 'In both these cases temperature rise did not exceed 3F*. 'It was concluded,-therefore, that the effect of thermal ~non-uniformities on containment penetrations in ~PWR plants is negligible.and would not cause their failure.

In BWR plants the inflatable seals in personnel airlocks may be exposed during

. severe accidents to thermal energy fluxes coming from a standing hydrogen flame ~

, on the surface of,the suppression pool. Although these fluxes may significantly raise' seal temperature, it still wi_11 remain below the value at which the seal may fail. -Later in the accident, however, sudden increase of drywell tempera-ture may cause the seals to exceed the failure limit and a leak path may develop

between drywells and wetwells.

References Commonwealth ~ Edison Co., " Final Safety Analysis Report, Zion Station, Units 1 and 2," Section 5.

- Mississippi Power & Light'Co., " Final Safety Analysis Report, Grand Gulf Nuclear

. Station, Units 1 and 2," Section'6.

i U.S. Nuclear Regulatory'Commis'sion, NUREG/CR-2228, " Containment Response During

' Degraded Core Accidents Initiated by Transients and Small Break LOCA in the Zion / Indian Point Reactor Plants," July 1981.

NUREG-1037 A-17 i -. . . . ,' , - . -,,,, ..-.-. -. - . - . . . . - _ _ . _ . . - _ . - . _ - . _

=

~ ~

e - r

.v. ,

l

(' . -

j

~

- NUREG/CR-2530', " Review of the Grand Guli Hydrogen Igniter System, March

^ 1^

~ ?g -

---; draft NUREG/CR-3234,-"The Peter.tial for Containment Leak: Paths Through Electrical Penetration Assemblies Under Severe Accident Conditions," April 1983. j i

s*. -e ,

, J

/

e

\'..

4 r' -.

t J

[ . .1 5 O p i :

y .Y'-

~

m R

a 4

.g e1 a

t r d w.

4 b

,y.

y e

j t 1

4

% P 3

{

t n 1 i ~. e s

NUREG-1037.'~ A-18

y 1

Nomenclature A -

. area of. purge valve exposed to ' thermal energy 2

transfer from environ-ment,'or area of drywell wall element, ft a - surface area of'an element of per_sonnel airlock wall, ft 2 B

parameter defined by equations 3 and 20

- B' -

paramster defined by equation 24 -

c .- specific heat, Btu /lb F

- D-

' gas constant, psia, ft3/lb mol R F

view factor (in radiation equation). -

-G

. parameter defined by equation 11 H' .

. parameter defined by equation 11 h - heat transfer coefficient, Btu /hr ft2 op I -

parameter defined by equation 11 J -

parameter defined by equation 11 K.

parameter defined.by_ equation 10 k - thermal conductivity, Btu /hr ft F ,

m, .

-- mass of air in personnel airlock, lb M,.

- molecular weight of steam (water), 1b m - mass of' steam ingressed to personnel airlock through the damaged seal, lb s

N -

parameter defined by equations 4, 13, 15, and 21 N' .

parameter defined by equation 25

n. -

number of nodes P

pressure of gas, psia

(

Q

- -heat transferred to personnel airlock gas, Btu R

parameter defined by equations 5, 14, 16, and 22 parameter defined by equation 26 R' -

r - radius of circular surfaces in equipment hatch or personnel airlock, ft Ar - increment of radius of circular surface, ft

'NUREG-1037 A-19

a p 15 thickness =of equipment hatch cover, thickness of frame, and thickness of personnel _ airlock walls, ft .

t -

-temperature, *F

'At

~

temperature increment, F t '. temperature after elapse of time interval AG, *F

-V -

volume of personnel airloc'k, fta W --

thickne s of drywell wall, ft '

x. --

length of personnel airlock, ft Ax -

increment of length of personnel airlock, ft a -

thermal diffusivity, ft2/hr c -

~ thermal emissivity of surface c -

thermal emissivity of gas g

'O -

time, hr AO -

time increment, hr p -

density of component material,1b/ft3 o' -

Stefan-Boltzman constant, Btu /ft 2 hr ( R)4 heat fluxes, Btu /hr ft2

$ DD heat flux from drywell to drywell wall, Btu /hr ft2 '

heat flux from hydrogen flame, Btu /hr ft 2

$wD heat flux from wetwell to drywell wall, Btu /hr ft 2 t

NUREG-1037 A-20

. Subscripts a -

air c -

PWR containment

~cD convection to personnel airlock wall exposed to drywell environment ci -

incoming heat by convection co. -

outgoing heat by convection cw -

convection to personnel airlock wall exposed to wetwell environment cyl -

cylindrical portion of the personnel airlock D -

drywell conditions DD-' -

heat flow between drywell and drywell wall D -

Saturated at drywell condition sat ep -

end plate portion'of the personnel airlock G -

wetuell gas-g -

gas

-i -

incoming n -

value at n mode o -

outgoing; initial temperature; outside temperature p -

personnel airlock condition pf -

personnel airlock at seal failure ps -

personnel airlock air / steam mixture

! RD -

radiation to personnel airlock wall exposed to drywell environment l

l Ri -

incoming heat by radiation i

Ro -

outgoing heat by radiation Rw -

radiation to personnel airlock wall exposed to wetwell environment s -

steam j s0 -

surface of personnel airlock exposed to drywell environment I

sp -

temperature spike NUREG-1037 A-21

I sw -

. surface of personnel airlock exposed to wetwell environment v -

drywell wall vo -

initial drywell wall: temperature v1 -

drywell wall temperature at the end'of first interval v2 -

drywell wall temperature at the end of second interval w -

wetwell conditions wD -

heat flow between wetwell and drywell wall 1 -

value at node 1 2 -

value'at' node 2 NUREG-1037 A-22

1 2

Area Considered t

l $ } l I

t

/

Q 3 :

Figure A.1 Equipment hatch i

., NUREG-1037 A-23

Y Finite Difference Solution

[If f Adiabatic Surface f . .

AX t h

n t

2 6 Region 2 1

Ir Outside Inside Containment Containment C-r*--l U

t,

out p *in t, a Region 1 U

J t n+1

___^, , ,

4 p

n-1 P

n :

s 2

P i n+1 x- -r -= c y

t,,,

Region 3 v

i n

I l////////U'f//////b 'o j

Heat Sink Figure A.2 Model of equipment hatch t

NUREG-1037 A-24

-- - - - - . . --n.. -n, , - , - -,. , -- , , , . - - _ , , . - , - . --.--.----.--....--.- . _ . .--.-.- _ ,. ,,., ., . - -,- ,. -

900.0 800.0' -

E L 700.0 -

lii 600.0 -

I h 500.0 -

400.0 -

1

$ 300.0 -

N 200.0 -

100'0

' - I I I I I I I I I O 100.0 200.0 300.0 400.0 500.0 600.0 700.0 800.0 900.0 1000.0 1100.0 Time (Minute)

Figure A.3 Compartment temperature during hydrogen burn in PWR e

NUREG-1037 A-25

l l

Gasket -

Duct Disc

' T M a

a

/

52" 41" 42

p -..".- t ir 1 I

--.- e s" gj' V

y

--* 12" +---

Figure A.4 Model of a purge valve I

NUREG-1037 A-26

. . . ~ _ . - _ - . _ _ _ _ _ _ _ _ _ _ _ ._ _ . . . _ . . _ _ .__ _ .

Drywell Wetwell Side Walt Drywell Side t,, t:0 h

h" N t, k

2 .wC

wD ,

DD

x :

wWe Figure A.5 Personnel airlock NUREG-1037 A-27

t, C X

-o-t n tn.3 tn+i tn tn-1 f t3 t2 Drywell" Wall - ti Personnel Airlock

+ t n tn_i ((

t n +1 tn tn-1 f{ 13 t2 tw ,

W Region 3 Region 1 Region 2 Drywell Side l

l Figure A.6 Model of personnel airlock i

i I

l NUREG-1037 A-28

900 -

Drywell Wall 800 -

tw j) to

\ Personnel Airlock Wetwell Drywell Side Side 600 -

It y 500 - /, - -- -- tw k !

t /

/

<m -

/

I I

h j

/\

300 -

/\ /

]

(

HFlarnel*

0 1 2 Time (Minuto)

Figure A.7 Thermal response of personnel airlock seals with drywell side seal survived NUREG-1037 A-29

900 -

'Drywell Side Seal Failed Drywell Wall 800 -

tw / tD

+

\ Personnel Airlock

~

Wetwell /

/ Drywell Side Side Seal Failed 600 -

[ , - - - - - - -

$ /

s 500 -

/

5 I E

l

, I I

400 -

1 I

A

/\ I 300 -

/ \ /

200 -

H Fla m e} e 0 1 2 3 Timo (Minute)

Figure A.8 Thermal response of personnel airlock seals with drywell side seal failed NUREG-1037 A-30

ll e*

er APPENDIX B LEAK AREA ESTIMATES FOR POWER REACTOR CONTAINMENTS DURING SEVERE ACCIDENT CONDITIONS NUREG-1037

i 9

J LEAK AREA ESTIMATES FOR POWER REACTOR CONTAINMENTS

~DURING SEVERE ACCIDENT CONDITIONS Final Report B. L. Barnes T. L. Bridges '

W. T. Dooley T. E. Rahl M. J. Russell Applied Mechanics Branch EG&G Idaho, Inc.

November 1984 i

l

(

NUREG-1037 APPENDIX B :

r CONTENTS

1. INTRODUCTION ..................................................... I

2. EVALUATION OF CONTAINMENT LARGE OPENING PENETRATIONS . . . . . . . . . . . . . 9 2.1 Types of Containment Large Opening Penetrations ............ 9 2.1.1 Equipment Hatches .................................. 9 2.1.2 Drywell Heads ...................................... 9 2.1.3 Personnel Airlocks ................................. 13 2.1.4 Fuel Transfer Tubes ................................ 13 2.1.5 Miscellaneous Pressure Unseating Penetration Closures ............................... 18 2.2 Methods of Analysis ........................................ 18 2.2.1 Analysis Methods for Personnel Airlocks ............ 18 2.2.2 Analysis Methods for Pressure Seating Closures ..... 22 2.2.3 Analysis Methods for Pressure Unseating Closures ................................. 22 2.3 Containment Large Opening Penetration Details and Leak Area Estimates ........................................ 24 2.3.1 Zion ............................................... 24 2.3.2 Surry .............................................. 26 P.3.3 Peach Bottom ....................................... 27 2.3.4 Limerick ........................................... 29 2.3.5 Grand Gulf ......................................... 32 2.3.6 Sequoyah ................ .......................... 33

3. EVALUATION OF CONTAINMENT PIPING PENETRATIONS . . . . . . . . . . . . . . . . . . . . 35

4. RESULTS .......................................................... 56

5. CONCLUSIONS ...................................................... 58

6. REFERENCES ....................................................... 59 APPENDIX A*-CONTAINMENT PENETRATION ELASTOMER SEAL TEST . . . . . . . . . . . . . . . A-1 APPENDIX B--ZION CONTAINMENT LEAK AREA ESTIMATES ...................... B-1 APPENDIX C--SURRY LEAK AREA ESTIMATES ................................. C-1 APPENDIX 0--PEACH BOTTOM LEAK AREA ESTIMATES .......................... D-1 APPENDIX E--LIMERICK LEAK AREA ESTIMATES .............................. E-1

Work supported by the U.S. Nuclear Regulatory Commission. Office of NucIcar Regulatory Research, under Inoragency Agreement DOE 40-550-75 with the U.S.

, Department of Energy.

NUREG-1037 U APPENDIX B

i i

APPENDIX F--GRAND GULF LEAK AREA ESTIMATES ............................ F-1 APPENDIX G--SEQUOYAH LEAK AREA ESTIMATES .............................. G-1

FIGURES PWR large dry and subatmospheric containment . . . . . . . . . . . . . . . . . . . . . 3 1.

2. BWR Mark I containment ........................................... 4

3. BWR Mark II containment .......................................... 5

4. BWR Mark III containment ......................................... 6

5. PWR ice condenser containment .................................... 7

6. Pressure seating equipment hatch ................................. 10

7. Pressure unseating equipment hatch with integral personnel. airlock ................................................ 11

8. Drywell head flange details ...................................... 12

9. Personnel ai rlock with rotating doors . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14

10. Personnel airlock with flat bulkheads ............................ 15

11. Equipment hatch with emergency escape airlock .................... 16

12. Fuel transfer tube ............................................... 17

13. Control rod drive removal hatch .................................. 19

14. Personnel airlock isometric plot ................................. 21

15. Model for determination of flange separation for pressure unseating closures ............................................... 23

16. Zion personnel airlock leak area estimate . . . . . . . . . . . . . . . . . . . . . . . . 25

17. Peach Bottom personnel airlock leak area estimate . . . . . . . . . . . . . . . . 28

18. Embedded piping penetration ...................................... 36

19. Flanged piping penetration ....................................... 37

20. Flued head piping penetration .................................... 38

21. Single bellows piping penetration ................................ 39

22. Double flued head piping penetration ............................. 40

23. Flued head / bellows piping penetration ............................ 41

24. Double bellows piping penetration ................................ 42

25. Typical piping penetration with bellows at the precess pipe ...... 43 TABLES

1. Surry leak area estimates ........................................ 27

2. Peach Bottom leak area estimates ................................. 29

3. Limerick leak area estimates ..................................... 31

4. Zion piping penetration summary by type and size ................. 45

5. Surry piping penetration summary by type and size ................ 46

6. Peach Bottom piping penetration summary by type and size ......... 47

7. Limerick piping penetration summary by type and size ............. 48

8. Grand Gulf piping penetration summary by type and size ........... 49

9. Sequoyah piping penetration summary by type and size ............. 50

10. Parameters used in the analysis of piping penetrations ........... 52 NUREG-103/ jjj APPENDIX B

LEAK AREA ESTIMATES FOR POWER REACTOR CONTAINMENTS DURING SEVERE ACCIDENT CONDITIONS

1. INTRODUCTION Nuclear power plant risk assessment studies are performed to predict

.offsite releases of radioactive fission products for' severe accident conditions. In the past, virtually all of the risk assessments were performed characterizing the plant containments using a threshold model

, which defines a threshold pressure, with some associated uncertainty, at which the containment will suffer a loss of holding capability resulting in a significant release of fission product inventory. If the containment pressure loading is calculated to be below the threshold pressure, the containment is considered to be intact and the offsite consequences are therefore quite low. Some recent analyses have pointed out that significant containment leakage may occur prior to reaching ultimate pressure capability. Currently, it is felt by many, that characterizing containment behavior using leak before failure models will provide more realistic results than the threshold models. NRC established a Containment Performance Working Group (CPWG) consisting of NRC staff members, Brookhaven (BNL) and Sandia National Laboratories (SNL) and EG&G Idaho, Inc, personnel to assess the " Leakage.Befare Failure" concept of predicting containment responses during severe accidents. EG&G Idaho's responsibility includes determining containment leak area estimates for containment piping and large opening penetrations as a function of containment pressure and temperature for severe accident conditions. These leak area estimates are to be combined with leak area estimates for containment electrical penetrations and purge and vent valves for performing containment response analyses for severe accidents. The leak area estimates for purge and vent valves and electrical penetrations are being determined by BNL and SNL respectively. The containment response analyses are also being performed by BNL.

Containment behavior is being evaluated for each containment type.

CPWG has selected the following six reference plants for evaluation:

I APPEN0!X B NUREG-1037

r ,

-1. PWR, Large Dry--Zion

2. PWR, Subatmospheric--Surry

3. BWR, Mark I--Peach Bottom

4. BWR, Mark II--Limerick

,5. BWR, Mark III--Grand Gulf

6. PWR,' Ice Condenser--Sequoyah Figure 1 illustrates the PWR large dry and subatmospheric types of containments. Figures 2-5 tilustrate each of the remaining four basic types of nuclear power plant containments.

Containment leak area estimates were determined for each of the six reference plants for severe accident pressure and temperature conditions.

Originallyitwasintendedthatthree(high, medium,Iow)leakarea estimates as a function of containment pressure would be determined for each' plant.' The high leak area was to be one which has a small probability of being exceeded. The medium area was to be a best estimate value. The low leak area was to be one which has a high probability of being exceeded. For penetrations with elastomer type seals, the leak areas were to be determined by first determining the structural response of the seal joints due to pressure loading and then using seal behavior assumptions based on engineering judgement or test results to obtain the three desired leak areas. Well into the program a small scale series of seal behavior tests were conducted to try to determine the seals behavior characteristics under these conditions. A double o-ring configuration with neoprene seals and a double tongue and groove configuration with silicon rubber seals were tested. The results of these tests, contained in Appendix A, suggest that seal leakage for flange separations below 0.030 in, will be nearly insignificant. Due to the limited amount of testing performed it was concluded that additional seal testing must be performed before attempting to characterize the behavior of seal materials during severe accident NUREG-1037 2 APPENDIX 8

E -

n. ester Sesel centeinment liner .

Troueys 1 2

, e m J . 2 %,

',)

Steem genereters Pressurteer ,.

l.

.i

O

/\

n_ \ \

week wene

.\.-

z:,." .

(typteell m i

" ,e *

,'.=

'. e, m ara - '.',.

, yr-- J1 I . .,.

g

. . . . . a bn.w h som

g. '

,'.'4*h h*,

6

^

7/.' - nefueling , *

.: .V '< cmity i :/ . A V t1 l ="b"g f

?. *' f,*. ',t'

'4 i f?,: [L. - ;.-

~

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Y 'e ',.'.

j ),.;* . useusuu ;. . . , f.'., ,' E I(-

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. .s* . :e. <t ... -

s s,. . ..

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l w. .. .. .

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[ r ,"; w [s'l- , ..' & -- -

.,h.

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ett

.;,?;;.,. t....e).,..;: 1 ; e,'v,',..

.~, ' .: 9.r.5.*6. :n ,su,,.

.e

. . .. :.y; :.;si[:

.*.,.'.x.

.%: .e .e ,e ..), .lo,' . . .. ;.,',

. ."t?:,

,,; .g': . .'. fr. '?/.,1,;i e .*. .

Messeersev6ty Figure 1. PWR largo dry and subatmospheric containment.

NUREG 1037 3 APPEN0!X 8

Reestor $uilding L -.- L ^i containment) e j_Orywell a -r .

n i n

r' H ' -

, =

\ ... I

Meector "

shield well . -

'of0~*

-- , Reenter 0'e

_ eel

, s' '

, Drywell

, Vent pipe PodestelN .

70 ft N P."""" #

~

l %

- internel

. ,,ing heede,

/

. ,Downeemere

R pTorut suppreselen M . '5- f*o N

/

A chamber

-Prenwee

~

'o,'S **8='*a pool Figure 2. BWR Mark I containment.

I NUREG-1037 4 APPENDIX 0

I s a i m ,

M \ f l

j n cto,

['

vessel N i

h- F I e /

i i

l ' Primary l

coritainment % ' -

Wet well %

%g _ ~

Dry weil % .

da t , Pedestal % 4 a I.

I Olaphragm% -

l i

Floot N -

1 x- 14 _A = >

Oowncomers - '-

',s 1 .,

fecojay,

[ ;. ,

%QD, ressbon .. . . .

p.. ,

I . I 1

l figure 3. BWR Mark !! containment.

l l

l i

i NURI.G-10.1/

5 APPINDix 0

)

i >

r-l Primary containment O.25-in steel liner Polar crane

. . e.

CCD/,

t

?. t ' ?.-

l

. ,e , -

.. '?

2.5 ft eenerete tg (Detail drowing not to scale)

Upper containment poet

&

Note: Upper containment v* \ ..

t ',

' M?',

,- ~ . -

dome end so.or t

.h g s.4 wetwell enmmunieste l' ')

? t ,

1 '

. . 'f' with eadi other l.? \ 7 l l q.

e i 1i

'] l g.. . ,.. .c e.,. - l,). I i ic. . .,,..,. ,

.. - Temr'orary 4 %', e. j, p K fuel storage

[I {l 't Drywell l l "' # .,

. ,1

( ' y m- . i.'. .

. = ~ .'.

i t , ,

, , . .., ; ? ; . .,,-

m . -

I. -

N , '.5 '

s[

- ,l.

.c .. . :_.,,ese,o, : .. . .

,t ' '4, '. vessef '.. '.';

/ /

i '.1, p orywell wel f'.

' .t

/ <t

.: Vent / , Shield I',lWetwel ess !

,'*p t,T:Wb.,s. ,, annulve Weir 4., ~, ,,

/:-

.,' '..' l.

well .,'<

P.

' . -n. - .:t 7 ', ',,

- ) m al ', ' , __ .

't,,,,,,

i. '

namenna g

'. Suppression . . I; 1 ' '

Hor rontal A.,

/ .

.: poet = W "" ,

rede tes

\'l;'l.; r/ '

:. , ' = ',.'.

f,* ig*. 1 ,;,e ;.,,.,,l =

m s.* . .. .

f. .. - .h,,

a

,,.,s.,

a.,',. .. i .. C.

' ,., ., p.

. . . .p. , ,. t r .* . . 'g'.

, , . . e,. .. ,.f. . ' . , g, ..' g.. , .e contsnment

.,. . :. .j, L,t,. . *.,.; , : .j^ t. .:. o . :. n t

rigure 4. DWR Herk !!! contatnment.

i NURCG 1037 6 APPLNDIX 0

I I

eenesinment x

%.einment son t k, ~ .

..g

' 1 .

I A A -

,en n , _ 7, -.

- L u- .em - , .

1 v

.j l

1 f- ,h

\

g. l

. t . .

5 l l e ;

p - . ..-n, .

.e 1 . .

1 vent evenine .. s u s t

'g y,l = g ;.;;;n - '

t ;

. . cI t vi- o h. .,, a y

w. + - . -

,e

, ,,i ,, ...

,n_gr

~ 4, , .

w

, w..a, yme, ....n.. g. . . . . . ..:::

1

i g.,%j;0'.;.?.:.1. ::

i- - - - 4. . . A. , -

e , ,, , . ,, ,, .

w

. W..ifi N

.n. - _- .e ~w. n. . .

n

- ~ .

p...s*l'.

menue:m .

a

'***.../*.*j:,. t s..........*.,,,.

... ,,.s ,...e s ...o ....

Deew, woes inesses eosimseinen name veawn.

T l

figure 5. PWR ice condenser containment. ,

l l

i 7 APPCN0!X 8 !

WRCG 1037 I

conditions. SNL is currently undertaking such a program for the Office of Nuclear Regulatory Research of the NRC. Therefore, rather than determine three containment leak area estimates as originally planned, it was decided g that only one leak area estimate would be determined for each containment type based on the penetrations' calculated structural response (flange separation) only, realizing that this is an upper bound value. In addition to the assumption that the seals have no resilience, it was also assumed that the personnel airlocks would have only one door available to resist the containment pressure and thermal effects of the pressure unseating closures was neglected. Both of these assumptions are consistent with the goal of determining upper bound leak area estimates.

Sections 2 and 3 describe the analyses performed to determine the leak area estimates for large opening penetrations and piping penetrations respectively.

NUNEG-103/ 8 APPENDIX 0

2. EVALUATION OF CONTAINMENT LARGE OPENING PENETRATIONS 2.1 Types of Containment large Opening Penetrations Leak area estimates were determined for containment large opening penetrations for each of the reference plants. Combinations of the following types of large opening penetrations were evaluated for the six reference plants depending on the specific containment (a) equipment hatches, (b) drywell heads, (c) personnel airlocks, (d) fuel transfer tubes, (e) CRD removal hatches, (f) manhole hatches, and (g) suppression chamber access hatches.

2.1.1 Equipment Hatches The containments for the six reference plants all have one or two equipment hatches. The equipment hatches are either pressure seating with mating flanges on the inside of the containment as shown by Figure 6 or pressure unseating with mating flanges on the outside of the containment as shown by Figure 7. These hatches are typically spherically shaped heads ,,

with containment opening diameters varying from 12 ft to 20 ft. Some of these hatches have an integral personnel airlock such as that shown by Figure 7 or an emergency escape airlock. The equipment hatch seals are either double o-rings or double tongue and groove or double gumdrop shaped seals. The seal materials are generally silicon rubber with some being either neoprene or ethylene propylene.

2.1.2 Orywell Heads BWR Mark I and !! containments have drywell heads with inside diameters of 32 ft to 37 ft-7-1/2 in. These heads are semi-elliptical carbon steel heads with pressure unseating flanges as shown by Figure 8.

The seal designs for these heads are either of double tongue and groove design or a double gumdrop shaped design. The seat materfat for the two drywell heads evaluated is silicon rubber.

9 APP [NDIX 0 NURLG 1031

~

1

~

e f,%

k

F .. 'RJ :

ffj?.T'i

. .? Containment Liner A

,m I

e

, ~

l I

i N

f s s

! N

.' W.

s

? e t

h

. . ...,.'. . f, -

1 ,-

l * . '. ' t . . . .t .,, .,%

,8,*,5'.**,.',...r,.'

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+.

i

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I Detail A MJAlt03 4 Figure 6. _ Pressure seating equipment hatch.

HUAEG 1037 10, APPEN0!X B

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m oe

c e.e . tm., -

y ,.s .g. -

w s , W .. 'L y, s.I mr g4,yj pg.

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,,,,,,, , smcon.nu,bber a..n. .

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"- \' '

/

[J,t /~

/

Sd lA h ,

n......

2 o

m m

z o.e x

co Figure 7. Pressure unseating equipment hatch with intergral personnel airlock.

1 I

l l

m .

l}

s ll 1

hI ANNNN .

N

'-}.

k

~

\

k Silicon-Rubber \\% $5

,=

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kasxxws.ss_xw

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h.Il il EEEEE

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8 Figure 8. Drywell head flange details.

12

- NUREG-1037 .

APPENDIX B~

2.1.3 - Personnel Airlocks-Each of.the containments evaluated ha've one or two personnel airlocks. These _ personnel airlodks can be classified by four different ,

types. The first type shown by Figure 9, has a pressure seating door on the inside and a pressure unseating door on the outside. The rotating doors have radial lugs which eng' age with the barrel section. The doors have double o-ring' seals. The second type of personnel airlock has flat exterior and interior bulkheads with both doors being pressure seating as shown by Figures 7 and 10. The door seals for this type of personnel airlock generally are a double tongue and groove design, however, one ,

personnel airlock evaluated had only single tongue and groove seals. The third type of personnel airlock is the same as the second only it has two inflatable seals at the outer periphery of the door between the door and

~

door frame. The door pressure loading is transmitted to the bulkhead through vertical door stops located at the door sides,-thus the inflatable seals are not.affected by the door pressure loading. Relative deformations ,

of the door and door frame do not affect these seals' ability to maintain a seal. Thermal degradation of these seals would, however, result in very significant leak areas. The fourth type of personnel-airlock evaluated is the emergency escape airlocks as shown by Figure 11. These airlocks are much smaller with circular 30 in diameter doors which are both pressure seating. The door seals for the one emergency escape airlock evaluated were single seals with two raised sealing surfaces (double dog eared design). >

2.1.4 Fuel Transfer Tubes i

Containment fuel transfer tubes consist of a 20-24 in. diameter pipe i

i extending from the reactor cavity through_the reactor cavity wall, containment wall, and fuel pool wall. As shown by Figure 12, the Fuel Transfer Tube (FTT) is sealed on the containment side by a blind flange and on the fuel, pool side by a closure valve. The FTT is supported by a seal I plate which attaches the FTT to the containment liner and by a vertical i support off of the fuel pool. For a leak to occur between the FTT and its I' !

I 13 APPENDIX B

-NUREG-1037

_-._._;.____ _ __ __.._.7,,__ .,c__., ._ , ,,

~ . . ,

f f

- l 2

C j .

Containment '

Liner pm w- ... .

o A

w . .

N . *

] . ,

. 2.*.*. .

r

,_s Q

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'} ._

................................................................................. ......c

n. ,. . . . .. . .

m

,Ilj > T i k-w i

t .. .

. . - Double Gasketed

~ Rotate to Lock" Type Closure.

Both Ends

%v 2

m o

m z

o X

cm Figure 9. Personnel a'irlock with rotating doors.

z C

=

m o

e p.a O

W N

I inside

- Containment

- EPDW

'-Wm . !>

E

' /

/

, . WM , $

,, ...... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .., . //////////

m j

l

{

=a = DETAll A I.l n 5 T..l. .,5 D

H l

o v

$ Figure 10. Personnel airlock with flat bulkheads.

~

X CD

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e

.m .=m, i

l Containment l

Liner

x h '

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l l

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I

.:=mm , .: ?.

.* ?

Ifll : ; .. . . . . . *. 1 f; . ! .

..- -/

! ,' M. !

H li t

. 1 ii . .

-Figure 11. Equipment hatch with emergency escape airlock.

16 NUREG-1037 APPENDIX B

~___

. x C ^

.= Containment -

a L.mer s

w --

\ ,, ,, ,,, , ,, , ,__ T

~

,. e. .

Fuel Building

. ~

Rellows

. ~

..5 .. . ;,

(typical '-

3 places) e. ' ' . . . 0. '..

. -. . ~e, . / .. ~.,.

,.i e... . ,' '

"$ Closure Valve

t. - .. ,.

_. y 4' ',' "~.:

q. .. . . ' '

JUI IU1 r fugt fugl i r fuU1 I4 "

" .l Double Gasketed

Blind Flange M j . .... . .. ,.

1 IMlf-"UUU l L dC p

.. . , . . . - Support

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8' c,. . : .,

. .. . . .. e

. ~

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TLB784-1

>o m

o Figure 12. Fuel transfer tube.

z t C3 g e X

CD

. , l containment penetration sleeve, the leak must penetrate a bellows on the containment side, the seal plate, and a bellows on the outside of the containment.

2.1.5 Miscellaneous Pressure Unseating Penetration Closures The remainder of the containment large opening penetrations evaluated

'are all pressure unseating closures such as the CR0 Removal Hatch as shown

. by Figure 13. Other similar containment penetration closures are the drywell head manhole hatches and the suppression chamber access hatches.

These closures are basically blind flanges with diameters of 24 in, to 52 in. attached externally to the containment. The seals for these closures are either double o-rings, double gumdrop shape seals, or double tongue and groove. The seal materials are generally either silicon rubber,.

neoprene, ~or ethylene propylene.

2.2 Methods of Analysis 2.2.1 Analysis Methods for Personnel Airlocks Personnel airlocks and emergency escape hatches with flat bulkheads were analyzed to determine the amount of leak area possible between the doors and bulkheads. These containment penetrations have two pressure seating doors to provide resistance to the containment pressure when they are both properly closed. However, Licensing Event Reports (LERs) document several instances where one of the doors was ajar for various reasons.

With that in mind, the leak area estimates for personnel airlocks were performed assuming only one door is available to resist the containment pressure.

The analyses of the personnel airlock bulkheads and doors were performed for the various bulkhead designs using linear elastic three-dimensional finite element models of the bulkheads. The computer code SAP-IV (Reference 1) was used to perform these analyses.

NUREG-1037 18 APPENDIX B

y .

t- ' g, y .I If2 i u

/

' /

l A

[- .

T.

e i ,

e // /M ,

s ..

... - :-- - Detail A=

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s .... ..

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. .. . .... Silicon-(

t .

, ' ' . . .j

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.x sN w'

i t

j t ,

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' ' . . .f. N p .?

.l. ..

.d, m

I Figure 13. Control rod drive removal hatch.

NUREG-1037 19 APPENDIX B n, -.r.,-,,:, - . - , , . , , , - - - , . ,,-.,.,,-,-n n~,--,,,,---,---n-,---, , - , - . - , - -

Advantage was taken of symmetry about the bulkhead's horizontal and vertical center 1.ine axes, allowing the use of one quadrant models of the bulkheads. The bulkheads were either modeled using plate elements or a combination of plate and beam elements, depending upon the bulkhead design. Figure 14 is an isometric plot of the computer model representation of the Limerick personnel airlock external bulkhead. This bulkhead was modeled using a combination of beam and plate elements.

Short, stiff truss elements were placed between the door and door frame.

If any of these elements went into tension, indicating separation, they were removed from the model. For this type of bulkhead, separation generally occurred at the top and bottom of the doors only.

In addition to determining the separations, the-bulkhead structural members were evaluated to determine stress levels. For cases where yielding of the support beams was initiated before the ultimate pressure capacity of the containment was reached, the rate of increase of leak area to increase in pressure was assumed to be 10 times the rate before yielding. This assumption is based on the common use of a plastic modulus of one-tenth the modulus of elasticity for austenitic stainless steels.

Similar analyses were performed for the bulkheads with radial stiffening gussets using models with plate elements only. The maximum separation determined from these analyses was found to occur at door corners. All element stresses were determined to be below the material yield strength for these bulkheads.

The analyses of the personnel airlocks did not include thermal loading '

due to mean wall temperatures or thermal gradient loading. For personnel airlocks it is difficult to predict whether the coincidental thermal loadings would add to or negate the leak areas due to pressure loading.

Thermal degradation of the seal materials was evaluated where it appeared this could result in significant containment leakage.

1 l

NUREG-1037 20 APPENDlX B

k N

N

% g N

\

k BEAM ELEMENTS VERTICAL CENTERLINE Figure 14. Personnel airlock isometric plot.

21 NUREG-1037 APPENDIX B

2.2.2 Analysis Methods for Pressure Seating Closures The evaluation of the pressure seating hatches consisted of checking these heads for buckling stability to assure their collapse pressure is beyond the ultimate pressure capacity of the containments. Hand calculations were generally performed to make this check, however, an extensive nonlinear analysis of Sequoyah's equipment hatch was performed by Ames Laboratory to evaluate all leakage aspects of that pressure seating hatch. The results of that analysis (Reference 2) suggest leakage is not likely for this type of closare. The model used to perform this analysis incorporated a maximum size hatch imperfection allowed by the ASME Boiler and Pressure Vessel Code. Buckling of the hatch was determined not to occur below the ultimate pressure capacity of the containment. Mismatch of the penetration flange and hatch flange due to ovalization of the penetration flange was not sufficient to allow leakage. Flange control was maintained around the entire hatch perimeter. None of the pressure seating hatches evaluated contribute to containment leakage.

2.2.3 Analysis Methods for pressure Unseating Closures Hand calculations were performed to determine the flange separation, as a function of containment pressure, of pressure unseating closures.

These calculations were performed using the two spring models depicted in Figure 15. The solution for the flange load is as follows:

-F F2 =K1(P 2)

(4) where K1 = bolt stiffness K2 = flange-shell stiffness F2 = flange load NUREG-1037 22 APPENDIX B

F = preload P = applied load.

The initiation of flange separation occurs when F2 becomes zero. Setting F2 = 0 in the above equation ~and solving for the required applied load to cause flange-separation (P 3

) results in the following equation:

p S ,(K1 K2

+ K2) p

The flange separation (A) for applied loadings greater than P3 is given by the following equation:

P-P 3 A=

K1 P

D K1 K2 I

t P

Figure 15. Model for determination of flange separation for pressure unseating closures.

NUREG-1037 23 APPENDIX B

^

1 i

i The analyses performed using this method neglected any structural properties of the seal materials, since the seal materials are very soft relative to.the other stiffness values. For situations where the seal ;

design was such as not to allow the mating flanges to make metal to metal contact, the flange separation was determined assuming there was no

, preload. As can be scen from the equations used to perform these analyses, i calculation of the flange separation due to pressure loading is somewhat sensitive to the stiffness values; however, determining the preload is of i

greater concern. This is particularly true when the preload must be [

cetermined from a specified bolt torque value, since the coefficient of :

friction can take on such a wide range of values depending on the I lubrication conditions used. Thermal effects were not included in the analyses of these pressure unseating closures. The thermal conditions

would tend to reduce leakage since the bolts on the outside of the containment would be somewhat cooler than the flanges. Detailed thermal i

calculations would be required to evaluate this condition in detail.

l Neglecting this' condition is consistent with the objective of this effort

{ which is to obtain upper bound leak area estimates.

2.3 Containment Large Opening Penetration Details '

and Leak Area Estimates 2.3.1 Zion I

! The Zion containment large opening penetrations consist of an equipment hatch with an integral personnel airlock, an emergency escape hatch, and a fuel transfer tube. Detailed information beyond that contained in the FSAR was not available for the fuel transfer tube, i j emergency escape hatch, and the equipment hatch. Based on evaluations of

{ similar penetrations for other containments a leak area of zero is

justified for these penetrations. Zion's fuel transfer tube is a similar design as that shown by Figure 12. The equipment hatch is a pressure ;

seating hatch with double gaskets. The emergency escape hatch appears to k j be of a similar size and design to that for Surry's containment

(Figure 11). That hatch does not contribute to containment leakage during !

severe accident conditions. !

i ,

24 NUREG-1037 APPEN0!X B

. Details of the personnel airlock was obtained from Reference 4.

. Zion's personnel airlock is essentially the same as the one shown by Figure 10. The personnel airlock barrel has a 122 in, diameter and is approximately 10 ft-18 in, long. The bulkheads and doors plate material is 1 in. thick stiffened as shown by Figure 14. The analysis of the exterior

' bulkhead of this personnel airlock is contained in Appendix B. The resultant leak area as a function of containment pressure is as shown by the following figure.

6.36 - - - - - - ' - - -

i i

i Ec i*

O !

m i 8 i m

x e

a)

-I 3,sg .. . . . - -

, l i :

! l i

i !

0.00 0.0 107.0 134.0 Pressure (psi) n.....

Figure 16. Zion personnel airlock leak area estimate.

NUREG-1037 25 APPENDIX B

2.3.2 Surry Surry containment's large opening penetrations consist of an equipment hatch with an integral emergency escape hatch, a personnel airlock, a fuel transfer tube, and a construction access hatch at the apex of the containment. The equipment hatch is a pressure seating type closure with double neoprene o-ring seals. Its mating penetration flange has an inside diameter of 14 ft-6 in. The attachment of the emergency escape hatch to l

the equipment hatch is by means of a pressure unseating flange connection.

The seals at this flange are double silicon rubber o-rings. The connection bolts (32 1 in. diameter) have a specified torque of 160 ft-lb. The emergency escape airlock barrel has a 5 ft-9 in. diameter and is <

12 ft-8-1/2 in. long. Pressure seating 30 in. diameter doors are located at both ends of the airlock. The configuration of the equipment hatch and l integral emergency escape hatch is as shown by Figure 11.

I l Surry's personnel airlock is as shown by Figure 9. The doors attach to the barrel section by a rotating lock arrangement. Twenty-four tabs at the periphery of the door wedge behind retaining tabs attached to the barrel end flanges. The rotating door seals are double neoprene o-rings.

l-l Surry's fuel transfer tube is typical of that shown by Figure 12.

A construction access hatch is provided at the apex of the l containment. It is a pressure seating, I in, thick, 27 in diameter blind l flange. It is seal welded to the containment liner,

The analysis performed to determine the leak area estimates of these penetrations are contained in Appendix C.

The only large opening penetrations which will contribute to containment leakage during severe accident conditions are the personnel '

airlock and the equipment hatch-emergency airlock flange seal. The calculated flange separations and corresponding leak area estimates for these penetrations as a function of containment pressure is tabulated below, i

26 NUREG-1037 APPENDIX B

TABLE 1. SURRY LEAK AREA ESTIMATES Containment Pressure Flange Separation Leak Area Penetration (psig) (in.) (in.2) l' 81 0.0 0.00 100 0.0006 0.13 119 0.0011 0.26 b

2 30 0.0001777 0.05 44 0.0002402 0.07 100 0.0004900 0.14 119 0.0005753 0.17

a. Equipment hatch-emergency airlock flange,

b. Personnel airlock.

2.3.3 Peach Bottom Peach Bottom containment's large opening penetrations are the drywell head, two equipment hatches, one personnel airlock, and one control rod drive removal hatch.

Peach Bottom's drywell head is as shown by Figure 8. It has a 32 ft inside diameter. The head is bolted with 68 2-1/2 in, diameter bolts which have a specified preload torque of 850 ft-lb. The drywell head seal is a double tongue and groove design with silicon rubber seals.

One of Peach Bottom's equipment hatches is a 12 ft inside diameter pressure seating spherical head as shown by Figure 6. It has a double tongue and groove seal design with silicon rubber seals.

Peach Bottom's other equipment hatch is a pressure unseating hatch with an integral personnel airlock as shown by Figure 7. This hatch penetration also has a 12 ft inside diameter. The hatch is secured in place by 24 1-3/4 in, diameter swing bolts which have a specified preload torque of 1900 ft-lb. The seal design for this hatch is also a double t:ngue and groove design with silicon rubber seals.

27 NUREG-1037 APPENDIX B L- - - - - - - - - - - - - - - - - - - - - - _ - - _ - - - - - - - - - - - - - - - - - - - - _ - - - - - - - - - - - - >

The personnel airlock is integral with the above described unseating equipment hatch. It has flat bulkheads with pressure seating doors. The inner bulkhead is stiffened with radial gusset plates. The exterior bulkhead is stiffened with horizontal beams. The door seals are a single tongue and groove design with silicon rubber gaskets.

1 Peach Bottom's CR0 removal hatch is a pressure unseating 36 in blind flange similar to the one shown by Figure 13. It is attached to its mating penetration flange with 36 1-3/4 in, diameter bolts. The flange seals are double silicon rubber o-rings.

The analyses performed to determine the leak area estimates are contained in Appendix D. It was determined that leak areas would occur during severe accident conditions for the personnel airlock and the pressure unseating equipment hatch and drywell head. Figure 17 shows the l

personnel airlock leak area estimate as a function of containment pressure.

I 2.10 - ~ -

i C i od i O ,

8 X y c 8 -

a ,

0,30 ' - -

93.1 150.0 Pressure (psi)

Figure 17. Peach Bottom personnel airlock leak area estimate.

NUREG-1037 28 APPENDIX B

p The leak area estimates for the pressure unseating closures is summarized in the following table.

TABLE 2. PEACH BOTTOM LEAK AREA ESTIMATES Pressure Required Leak Area Estimate To Unseat Flanges at 160 psig Closure < (psig) (in.2)

Drywell head 26.6 78.13 Equipment Hatch 82 4.15 2.3.4 Limerick Limerick's large opening penetrations are the drywell head, two Gquipment hatches, a personnel airlock, two suppression chamber access hatches, and a control rod drive removal hatch.

Limerick's containment drywell head is a 2 to 1 semi-elliptical head with an inside diameter of 37 ft 7-1/2 in. The head is constructed of 1-1/2 in, thick plate of SA-516, Grade 70 carbon steel. The drywell head is secured to the mating drywell wall flange by 80 2-3/4 in, diameter, 35 in, long bolts. The bolt material is SA-320, grade L43. Specified bolt preload is 157 Kips per bolt. The original seal design was to be a double tongue and groove design with silicon rubber seals. Due to a mismatch of the tongues and grooves, this design has been modified. The tongues were removed from the drywell head, and gumdrop shaped seals will be used. The drywell head has a 24 in. inside diameter manhole. This manhole has a 1-1/2 in, thick blind flange cover. The cover is secured to the manhole penetration flange by 16 equally spaced S/8 in, diameter, SA-193 grade B7 bolts. The manhole cover seal design consists of a double tungue and groove with 1/2 in, square cross section silicon rubbar gaskets. The drywell head assembly was designed for a pressure of 62 psig, 7 psi above the containment design pressure of 55 psig.

29 NUREG-1037 APPENDIX B L

Limerick's' containment has two 12 ft inside diameter equipment hatches. Both hatches provide access to the drywell at elevation 258 ft-0 in. One of the equipment hatches is a 1 in, thick dished head with a 9 ft-7 in, spherical radius. This head is sealed'to the outside of the containment, so that containment internal pressure tends to separate

the flanges. The other equipment hatch contains an integral personnel l ai rlock.~ This hatch also seals to the outside of the containment. Both hatches are secured to the containment by 24 1-3/4 in. diameter swing bolts with a preload of 50 Kips per bolt. Both batches have double silicon rubber seals. The seal joint for the hatch with the personnel airlock !

I consists of double gumdrop shape seals in the containment flange with a flat hatch flange. The other hatch seal joint is a double tongue and groove design. Both hatches were designed for a pressure of 62 psig, j t

l Limerick's personnel airlock is ar. integral part of one of the 12 ft ditmeter equipment hatches. It consists of a 12 ft-7-7/8 in, long horizontal barrel section with an inside diameter of 8 ft-6 in. Each end has a reinforced bulkhead with a 6 ft-0 in, by 2-ft-6 in, rectangular !

door. The-inner 5 ft length of the barrel section is constructed of 2 in, thick plate to resist external pressure loading, since most of this length l t

is located inside the containment pressure boundary. The bulkhead at the l inner end of the barrel is reinforced with redial gusset plates between the '

bulkhead and the 2 in, thick barrel. The outer barrel section is constructed of 3/8 in, thick plate. The outer bulkhead is stiffened by horizontal plate stiffeners across the top and bottom of the bulkhead door ,

l opening and between the barrel and door frame sides. Both bulkhead doors open towards the inside of the containment so that containment pressure forces the doors closed. The door's seals are a single silicon rubber i j gasket with two raised contact surfaces. A void space between those two l contact surfaces is provided to allow pressure testing of the seal, j!

i Limerick's two suppression chamber access hatches consist of 2-1/4 in. l

-thick blind flanges attached to the outside of 52 in. Inside diameter 1

, sleeve penetrations through the suppression chamber wall. The blind flange

! seals consist of a double tongue and groove joint with rectangular cross section silicon rubber gaskets. The blind flange hatches are secured to l

30 l NUREG-1037 APPENDIX B

their mating penetration flanges by 36 5/8 in. diameter, SA-193 grade B7 bolts. These hatches were designed for a pressure of 62 psig.

Limerick's control rod drive removal hatch is located at elevation 254 ft-6 in. This hatch is also a blind flange on the outside of th containment. The penetration sleeve has a 3 ft-0 in. inside diameter. The hatch seals consist of a double tongue and groove design with silicon rubber seals. The hatch is secured to its sleeve penetration mating flange by 36 7/8 in. diameter, SA-193 grade B7 bolts. This penetration is designed for a pressure of 62 psig.

The analyses performed to determine Limerick's large opening penetration le;k area estimates are contained in. Appendix E. Contribution to leakage during severe accident conditions is attributed to the pressure unseating closures only. The following table summarizes the leak area estimates for those openings.

TABLE 3. LIMERICK LEAK AREA ESTIM/TES Pressure Flange Required to Separation Leak Area Initiate Flange Due t at 140 psig Separation 140 psig 2 Penetration (psig) (in.) (in. )

Drywell Head 84.5 0.023 32.8 Equipment Hatches 74.5 0.0094 8.5 Suppression Chamber

0 0.0046 1.50 Access Hatches CRD Removal

0 0.001 0.11 Hatch a 0.08 Drywell Head 0 0.0011

a. These leak area estimates are based on no preload since the provided closure procedures for these penetrations did not specify a specific preload.

Subsequent discussions with Philadelphia Electric Co personnel indicate that these procedures are being modified and a specified preload will be implemented.

31 NUREG-1037 APPENDIX B

2.3.5 Grand Gulf Grand Gulf's (G/G) large opening penetrations consist of two personnel airlocks, an equipment hatch, and a fuel transfer tube.

The G/G's containment equipment hatch is a 19 ft diameter dished head.

It is located at elevation 172 ft-3 in. It is an internal head, therefore, containment internal pressure will enhance the seal between the hatch flange and the hatch penetration flange. The hatch seals are two concentric 3/4 in, diameter o-rings separated by 1-1/2 in. The o-rings arc EPDM, compound No. E603 with a durometer of 55-65.

The G/G containments have two personnel airlocks, one at elevation 124 ft-8 in, and the other at elevation 212 ft-8 in. Both personnel locks are of very similar design. Each personnel lock consists of a 10 ft-4 in, outer diameter barrel section 15 ft-0 in. long with bulkheads containing personnel doors at each end. The inner bulkhead and door is nearly flush with the containment inside radius. The outer bulkhead and door is 10 ft beyond the containment outer radius. The personnel doors are rectangular 3 ft-6 in, wide by 6 ft-8 in, tall. The doors are pressure seating against door stops located at the doot sides only. The doors are sealed by two inflatable seals around the door edge between the door and the door frame opening. The seals are attached to the door.

The inflatable seal material is EPDM.

The G/G fuel transfer tube is located at elevation 185 ft-4 in. It runs from the spent fuel pool building to the reactor containment building pool, spanning approximately 30 ft. It is supported and contained within a 48 in, bellows to absorb thermal and seismic loadings. Two remotely operated assembly closure doors seal off the 36 in, tube with a double o-ring seal and clamp design. The containment side of the tube is a compression seal which increases its effectiveness with rising containment pressure. The door and mating flange gap is kept within 0.015 in. of parallel. Both ends of the tube are submerged in water.

NUREG-1037 32 APPENDIX B

The evaluation analyses of G/G's large opening penetrations are c;ntained in Appendix F. It was determined from these analyses that c:ntainment leakage of these large opening penetrations will not result from severe accident pressure loading.

Leak area estimates were determined for suppression pool by pass. The discussion of the determination of these leak area estimates is also c ntained in Appendix F. The by pass leak area was determined to be 9.0 in.2 for time 0-5 hours of the severe accident and 135.0 in.2 for beyond 5 hours5.787037e-5 days 0.00139 hours 8.267196e-6 weeks 1.9025e-6 months .

2.3.6 Sequoyah Sequoyah's large opening penetrations are an equipment hatch, two personnel airlocks, and a fuel transfer tube.

The two identical personnel airlocks (located at elevations 696 ft and 737 f t) consist of inboard and outboard, 8 f t-7 in, diameter barrels, 3/8 in, thick, which are welded directly to the 2 in, thick containment p:netration sleeve. The outboard tube penetrates straight through into the shield building, but its connection there is a two ply bellows to absorb radial displacements of the vessel. These bellows do not constitute a containment boundary since they are beyond the point where the airlock barrels are welded to the containment. The barrel ends are fitted with 1/2 in, thick bulkhead plates which house the personnel airlock doors.

Both rectangular doors (3 x 6 ft) open to the inboard side, have double t:ngue and groove seals made of EPDM compound No. 603, and are made from plate.

The equipment hatch is a pressure seating closure. It is located at elevation 742 ft and has a 20 ft diameter opening with a 20 ft spherical radius head made of 3/4 in, thick steel. It is held in place with 20 1-1/4 in, diameter swing bolts and has a double gumdrop seal design which is increasingly effective with rising containment pressures. The shel! edge ring on the hatch has a cross section of 2 x 2 in.

33 APPENDIX B NUREG-1037

The fuel transfer tube design is a 24 in, diameter, 0.375 in thick i

barrel located at elevation 688 ft. It has a double o ring seal design on the inboard blind flange closure. It is attached to the containment wall with a sleeve / flange arrangement where the sleeve is 24 in. 0.D. by 1/4 in, thick and the flange is 1-1/4 in, thick. Bellows are included in the design, but do not form a part of the containment pressure boundary.

The evaluation analyses of these large opening penetrations are contained in Appendix G. It was concluded from these analyses that only j the personnel airlocks would contribute to containment leakage during severe accident conditions. The combined leak area for the two personnel airlocks is:

l l

i Containment Leak l Pressure AT'"

l (psig) ( i n . ,* ) '

0 0 25 0.19 50 0.38 l

1 i

i I

i i

l NUREG-1037 34 APPEN0!X B

3. EVALVATION OF CONTAINMENT PIPING PENETRATIONS Data for all piping penetrations from the six reference plants analyzed have been gathered and classified. Eight different types were defined. The eimplest type, the embedded pipe, consists of a section of pipe with the same diameter and the same or thicker schedule as the process pipe. Gusset plates are welded to the pipe for improved load transfer to the concrete wall. The pipe is connected to the containment liner through a plate of greater thickness than the liner. This isolates the liner from penetration nozzle loads (see Figure 18. All the remaining types of penetrations have a sleeve included in the design. The sleeve is similar to the process pipe used in the embedded pipe type, but its diameter is larger than that of the associated process piping. The various remaining types are defined by the way in which the process pipe is attached to the sleeve. For the flanged type penetration, the attachment is a flange welded to the sleeve with the process pipe (s) passing through holes in the flange face (see Figure 19). In the flued head type of penetration, the flange is replaced with a flued head. This is essentially a forged piece with integral welding attachment points for sleeve and process piping (see Figure 20). This figure also shows a typical sleeve design for a steel containment, which is a forged piece with integral attachment points for the contain:nent wall and flange or flued head. The single bellows type is similar to the flued head type, except that an expansion bellows has been installed in the sleeve between the containment wall and the flued head (see Figure 21). The remaining types of penetration differ from the above in that all have a double boundary. The double flued head type is similar to the flued head type, except that a flued head has been installed on both ends of the sleeve (see Figure 22). The flued head / bellows type is a double flued head design with a bellows / flange assembly replacing one of the flued heads (Figure 23). The double bellows type goes one step further, with a bellows installed between the containment wall and both fluedheads(Figure 24). The final type of penetration defined is the process pipe bellows type. This is similar to the double flued head type, except that a bellows has been installed between the flued head and the process piping (Figure 25).

35 APPENDIX B NUREG-1037

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1 NUREG-1037 36 pp ,

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MJR684 5 37 l

l t NUMG-1037 APPENDIX B ;

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' APPENDIX 8 NUREG-1037

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42 NUREG-1037 APPENDIX B

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cx Figure 25. Typical piping penetration with bello ~s at the process pipe.

A table has been prepared for each plant that lists the number of piping penetrations of each type by diameter of sleeve (diameter of embedded pipe for that type). See Tables 4 to 9.

The pressure boundary for_ a piping penetration includes the penetration itself and the connecting piping through the first isolation

' valve (s). Data has been gathered on the connecting piping from the six reference plants with varying degrees of completeness. Data for the majority of penetration piping was obtained for the sub-atmospheric, ice-condenser and MK-III BWR plants. Data for roughly half of the piping was obtained for the large dry PWR plant. Limited representative data was obtained for the MK-I and II BWRs.

Two types of loading were considered.

The first was the effect of differential pressure across the penetration. This is the same type of loading considered for the large opening penetrations. The second type is unique to piping penetrations.

Deflections of the containment wall will occur during an accident because of the internal pressure, and because the containment will experience thermal expansion due to accident temperatures.

Piping which passes through the penetrations is typically attached to structures both interior and exterior to the containment as well as to the containment itself. The resulting differential support displacement will induce loads in the piping and the penetration. This has been evaluated.

The effect of the differential pressure was evaluated using simple hand calculations. Evaluation of the displacement loading was more complex. The first step was to screen piping data for the plants which had substantial piping data available. The screening was based on the idea that the capability of a piping system to accommodate displacement controlled loading is dependant on the amount of piping in the system routed perpendicular to the direction of displacement. This comes from the significantly greater flexibility of piping for a bending load vs. an axial load.

It is also supported by the nature of the plastic response of piping J systems.

Piping components such as elbows and tees typically experience higher stresses than the associated piping. This is reflected in the stress intensity factors for tees and elbows, which are typically greater NUREG-1037 44 APPENDIX 8

~ .

t i

u,

. TABLE 4. ZION PIPING ~ PENETRATION'SLM ERY BY TYPE AND SIZE 1.

. ' Type

. e, Double flued Bellows at Diameter. Double Flued head /-

t-

-g (in.) fbellows -bellows process pipe head

~

--- -- -- 2 6

-- 2

-- 5 8 '

2 1 5 10

-- 9

-- 5 12-14 4 8 2 6' 16 6 7 2 18- -- ,

3

-- 2 20' --

s 24 1

-- -- 1 28- 1 32- 4 48' -- --< -- 2 >

~

150- 4

-This summary does not include; P-2, -3, -17, -18, -20, -21, -26, -35, -36, -47,

-49, -50, -61, -62, -64, -67, -73, -83, -84, -85, -87, -90, -100, -101, -103,

-105,-106,'-107,.-109,-123,-124(32 penetrations).

APPENDIX B NUREG-1037

i FL

+

TABLE 5.

SURRY PIPING PENETRATION

SUMMARY

BY TYPE AND SIZE-Type

!- ' Diameter Embedded

, (in.) pipe Flued ,

, head 0.745 1 I '

4 2

14 --

2.5 1 --

3 13 --

4 13 --

o .

.6 19 8 11 3 --

10-2 --

12 4 18- 3 4

'20 1

-24' 8 --

26' --

36 3 2 --

46 --

3 v

I i

L 4 NUREG-1037 40' APPENDIX B~

e s

TA8LE'6. PEACH BOTTOM PIPING PENETRATION

SUMMARY

BY. TYPE AND SIZE Type Diameter Single Flued Flanged Embedded

.(in.) Bellows Head

-- 378

'l -- --

-- -- 8 1.5~

-- 2

-- 2 3 ;

-- 1 4 --

6 -- 4

-- 2

-- 2 8

~12

-- 6 21 14

--- 2

-- -- -- 2 18 22 2 --

26 1 2 '

28 2 30 -- 1

-- 1 34 ,

36 1

! 42 4 44 4 NUREG-1037 APPENDIX B

, + ,

4

.t s

v .

' TABLE 7. >

LIMERICK PIPING PENETRATION

SUMMARY

BY TYPE AND SIZE' ,

Type .

Diameter. Flued ' '

Embedded (in.) Head Pipe Flinced-1 --

1 378 1.5 --

4 2 --

.7 --

8 --

3 --

6 --

4 2 10 --

, 6 9

' 3 --

' . i .:

8~ --

9

'10' 74.

2 --

12 11 2 ---

'14 ,

1 --

16 3 6 --

18 2 2- --

, 20- 24 2 --

'24 4 6 --

'30 1 --

36 1 --

42 6 -- .:

T F

d d

A 4

f J

, NUREG-1037 48 APPENDIX B 4

F

---s -%.- , rm-., ,- ,,.-.,e,- , -- ,-eii,--.. ..--...-+c , .--.--.-,,,,-,-e- -c, . , . -, ,.-c-

TABLE 8. . GRAND GULF. PIPING PENETRATION

SUMMARY

BY TYPE AND SIZE Type Diameter Embedded' (in.) Pipe F1anced 3 5 8 4 2 5 5 1 6 4 8 8 2 7 10 4 2 12 3 15 14 2 l' 18 1 7 20 2 24 5 3 26 -- 6

.36. -- 8 42 2 48 -- 4 I

i NUREG-1037 APPENDIX B

TABLE 9. SEQUOYAH PIPING PENETRATION

SUMMARY

BY TYPE ANO SIZE Type .;

l Diameter Single Flued

.(in.) Bellows Head Flanced 3 --

4 2 4 --

1 --

6 --

10 12

. 8 --

2 --

10 --

1

, 1 12 --

2 --

14 1 -- --

16 8 2 --

18 2 -- --

20 4 'l 1 22 1 ~~ --

24 4 1 --

28 1 -- --

30 7 -- --

52 4 -- --

1 NUREG-1037 50 APPENDIX B L

than one. Therefore the plastic response of a piping system would tend to concentrate in tees and elbows. Further, the significant loading in a piping system is typically bending moment and not axial force, so that the nature of the plastic response is predominately rotational and not translational. The result is that plasticity converts a piping system into a mechanism, with the straight runs of pipe acting as links and the elbows and tees acting as hinges. In the response of such a mechanism to an externally applied displacement, the relationship between the magnitude of displacement applied and the magnitude of the resulting hinge rotations is a function of the length of the links. In other words, for a given applied displacement, piping systems with more piping routed perpendicular to the direction of displacement will suffer less plastic deformation at the elbows and tees. Other factors considered in the screening methodology are size of piping (larger piping is less flexible) and material properties.

Application of the screening process to the piping data of plants with substantial data resulted in a list of penetrations ranked by increasing susceptibility to damage from the displacement loading. Piping systems predicted to be most susceptible were chosen for detailed analysis. For the plants with limited data available, all the penetrations were chosen for analysis. There were two exceptions to this, piping data for the Mark III BWR and large dry PWR plants were obtained too late in the schedule to allow detailed analysis. This data was reviewed to ensure that the piping systems which were analyzed are representative of the piping systems which could not be analyzed.

Detailed analysis consisted of linear and very limited plastic analysis of the piping systems using the NUPIpE-II computer code (Reference 3). Piping external to the containment was subjected to combined thermal and pressure induced displacement. The range of combined displacements used in analysis is presented in Table 10, along with the pressures and temperatures used in their derivation. The range of pressures and temperatures used were chosen to accommodate containment behavior analyses being simultaneously performed at Brookhaven National Laboratory (Reference 6). Thermal displacements were calculated on the assumption of a uniform thermal NUREG-1037 APPENDIX B

TAet.E 10 PAAa8ETERS USED IN T90E AAt44.YSIS Of PIP 81gG PEIIETIh4TIONS l

I s.:

w l C D Centsineont Itange of Deflections ieinfeer1 Temperature Pressure l Plant bene Ptont TC f*F1-Wertical stedle t fasial fin.I fin.1 Sescry Sashotmospheric 450 134 1.19/2.52 3.16/3.36 PW, Segeseysts Ice-condenser 425 65 0.34/0.59 2.69/2.76 - "

Mdit w tottee (pdIt/psK- 8 600 147 1.79/2.86 1.78/1.95 Limerick Stat /peK-I l 550 155 1.49/6.77 0.81/3.2s um N

D N

'O FWt EE

=

expansion from 70*F to the indicated temperature for all containment types. Thermal expansion of concrete containments was assumed to be l

controlled by expansion of the steel rebar. The calculations were based on an expression for thermal expansion offered in the eighth edition of the AISC Steel Manual. Pressure induced displacements were calculated differently for the various plants. The Surry pressure induced displacements were based on the maximum measured radial deflection at 45 psig presented in Table 15.5-6 of the Surry FSAR. This value was linearly extrapolated to 134 psia and increased 50% to account for nonlinear effects. The result was used for the pressure induced radial deflection of all Surry penetrations. Vertical deflections were obtained using the radial deflection and a thin walled cylinder assumption to obtain a vertical strain. This was converted to displacements for the penetrations based on the elevation of each above the base mat. The Sequoyah pressure induced displacements were based on a thin-walled cylinder assumption. Variations in the containment wall thickness were considered, as were variations in elevation of the penetrations above the base mat. Pressure induced displacements for Peach Bottom were based on the expression for the displacement of an internally pressurized sphere offered in the fifth edition of Roark. Interaction between the steel and concrete shells was neglected. The sphere was assumed to rest on its skirt, with relative motion between sphere and skirt occuring as the sphere expanded. Pressure induced displacements for Limerick were based on the containment study found in Reference 5. Due to the internal structure of the Limerick containment, the internal supports were assumed to displace vertically as a linear function of radial distance, from zero at the shield wall to the Reference 5 value at the containment wall. Supports outside the containment were assumed to be fixed. Piping inside the containment was subjected to the pressure induced displacement only, because the interior piping is surrounded by the same thermal environment as the containment wall. On the other hand, snubbers outside containment were assumed to function normally; those inside were assumed to be locked-up because the interior environment is, by definition, beyond the design envelope to which snubbers are qualified. In all cases, the first step was NUREG-1037 APPENDIX 0

to obtain a purely elastic solution for the conditions described above. If the most highly stressed point in the system indicated tnat a plastic hinge would form, and that point was in a straight run of pipe, then a very limited plastic analysis was performed. This consisted of installing a hinge in the model at the appropriate point, and applying the correct plastic moment to the ends of the pipe elements connected to the hinge.

This was an iterative process, because the orientation of the hinge predicted by the linear analysis was different than the orientation of the hinge in the converged plastic solution. Adjustment of the hinge for the succeeding iteration was based on the direction of maximum transverr.e moment in the preceding iteration. Convergence was obtained when the L transverse moment generated at the hinge became acceptably small. This procedure was attempted with two hinges in the model, but convergence did not occur. Although the elastic and limited plastic solutions fall short of an adequate plastic solution, they did provide enough information to ,

identify locations in the piping system most likely to develop a plastic hinge. The geometry of the model, the direction and magnitude of the applied displacement, and the location of probable hinge points were then used to predict maximum plastic rotations at the hinge points. The prediction was based on the idea that the formation of plastic hinges turns the piping system into a mechanism, as discussed earlier. Plastic rotations for hinge points that were explicitly modeled were taken directly from the limited plastic solutions rather than applying the procedure described above. '

Analysis of the single and double bellows type of penetration has not been done. These penetrations will accommodate relative displacements between piping supports primarily via bellows deflection. The state of the art analysis for bellows consists of elastic calculations that are correlated to proprietary empirical data not readily available. Since development of such data would be prohibitively expensive, the only ,

reasonable way to obtain an evaluation of bellows capability is to contract with the bellows vendors to perform the evaluations. This has been done by the owner of the Sequoyah Plant, with a resulting prediction of zero bellows leakage for Sequoyah.

NUREG-1037 54 APPENDIX B

Results of the analyses indicate that at the defined point of containment failure, piping penetrations will be in a state ranging from fully elastic in the best case to having plastic hinges formed in the pressure boundary which undergo plastic rotations in the range from 0 to 1 degree. Plastic hinges are predicted to occur exclusively in the piping, primarily near the junction to the penetration structure, but also at elbows and tees in the pressure boundary. The separate treatment of differential pressure loading and displacement loading was found to be justifiable because the piping was~ invariably schedule 40 or stronger wall piping. Stresses in this piping resulting from containment accident pressure were found to be negligible.

Based on the above results and the assumption that the pressures and temperatures used were not substantially different from those associated l tith containment instability, a failure in the piping penetrations before the containment becomes unstable is not likely. Although piping penetrations could fail after containment instability is reached, a zero leak area has been assigned for piping penetrations because the point at which containments become unstable is past the currently defined point of gross containment failure, (1% strain),

i t

4 I

NUREG-1037 APPEN0!X B L

4. RESULTS It was determined from detailed evaluations of containment piping and large opening penetrations that containment leakage during severe accident conditions is possible for two types of large opening penetrations. These types are personnel airlocks and pressure unseating closures. The containment leak area estimates for each reference containment as a function of containment pressure is:

Containment Pressure Leak Area (psig) (in. )

Zion 0 0.0 50 0.75 107 1.52 120 3.55

! 134 5.63 Surry 0 0.0 44 0.07 75 0.18 100 0.27 119 0.43 Peach Bottom 0 0.0 26 0.08 50 13.86 82 32.71 94 40.41 120 57.86 160 84.70 NUREG-1037 56 APPENDIX B

.a 4

Containment Pressure Leak Area (esto) (in. 2 )

Limerick 1

0 0 40 0.5 74.5 0.9 84.5 2.3 ,

120 28.3 140 43.0 ;

i Grand Gulf f i

No Leak Area Seouoyah 0 0 '

25 0.19 50 0.38 l

i I

e WREG-1037 APPENDIX B' f

l

5. CONCLUSIONS l

l !s is concluded from the analyses performed that the large opening j containment penetrations most likely to leak during severe accident l conditions are personnel atriocks and pressure unseating penetration closures. The evaluations performed for pressure seating closures other than personnel airlock doors suggest containment leakage is not likely for them. Evaluation analyses performed of these closures have shown that l metal to metal contact is maintained, they are not susceptible to buckling l instab111ty for pressure loading below the containments ultimate pressure

! capacity, the strain concentrations at the penetration-containment shell interface is not likely to cause leakage, and ovalization of the i penetration flanges is not sufficient to cause leakage. The analyses l

performed for personnel airlocks and pressure unseating penetrations closures have shown for some cases that severe accident pressures below the containments ultimate pressure capacity will cause metal to metal or flange separation. The leak area estimates contained in Section 4 reflect this l metal to metal separation condition. The reporting of these leak area

! estimates should not be interpreted to mean that leakage of these l containments will definitely occur during severe accident conditions, the 1

objective of this task is to determire upper bounds of leak area estimates. That being the goal, the following conditions were imposed:

(1) the leak area estimates assume the seals have no resilience, (2)thermaleffectswereneglected,and(3)thepersonnelairlockanalyses took credit for only one door, assuming the other door to be open. To obtain best estimate leak area estimates, the inclusion of thermal effects and characterization of seal behavior is required. The inclusion of thermal effects requires performing detailed heat transfer anlayses for specific accident conditions. The characterization of seal behavior currently is nearly impossible due to the lack of definitive data.

It is concluded from the piping penetrations evaluations performed that piping penetrations and associated piping are not likely to contribute to containment leakage during severe accident conditions prior to the l

commonly defined point of gross containment failure (1% strain).

NUREG-1037 58 APPENDIX B

6. REFERENCES

1. Computer Program Manual, " SAP-IV--Structural Analysis Program for Static and Dynamic Response of Linear Systems," University of California, Berkeley, California, June 1973. t

2. L. Greimann, F. Fanous, D. Bluhm, "Sequoyah Equipment Hatch Seal Leakage," Ames Laboratory, July 1984.

3. NUPIPE-!!, Quadrex Corporation, Version V4P!FIL released December 15, ,

1980, by Applied Mechanics Branch, EG&G Idaho, Inc.

4. " Primary Containment Ultimate Capacity of Zion Nuclear Power Plant for Internal Pressure Loss," Chicago Bridge & Iron Company, SL-3817, December 1980.

5. Technology for Energy Corporation, IDCOR Technical Report 10.1, Containment Structural Capabilty of Light Water Nuclear Power Plants, July, 1983.

6. C. Hofmayer, Personal communication Brookhaven National Laboratory, t

NUREG-1017 APPENDIX B

7 APPENDIX A*

CONTAINMENT PENETRATION ELASTOMER SEAL TEST o Work supported by the U.S. Nuclear Regulatory Commission, Office of Nuclear Regulatory Research, under Ineragency Agreement DOE 40-550-75 with the U.S.

Department of Energy.

^~I NUREG-1037 MMB

CONTAINMENT PENETRATION ELASTOMER SEAL TEST INTRODUCTION Under the predicted extremes of nuclear reactor containment pressure associated with severe accidents, the mating metal surfaces of some

" pressure unseating" penetrations will separate, thereby partially removing the static precompression on the elastomer seals between these surfaces.

Two seal designs similar to those employed by containment penetrations on the Surry and Peach Bottom nuclear power plants were leak tested as functions of flange separation and at containment pressures and temperatures approximating severe accident conditions. Both seal designs -

were found to be very leak resistant for the short time periods tested at all but the most extreme conditions of flange separation and operational temperatures. Flange separations required to produce incipient leakage at ambient temperatures greatly exceeded worst case severe accident predictions.

PROBLEM DESCRIPTION For nuclear reactor containment pressures less than the design pressure, separation of the " pressure unseating" bolted flanges is impossible because the combined preload of the flange bolts is greater than the product of the design pressure and the area of the penetration over which this pressure acts. However, given that predicted severe accident pressures exceed design pressures by a factor of about 2.5, some " pressure unseating" penetrations employing bolted flanges have been shown to separate on the order of 0.010 to 0.030 in.

The average amount of flange separation is very easy to calculate given the principle dimensions of the penetration, number and size of flange bolts, and a peak severe accider,t pressure. With considerably more effort, one can even accurately calculate the separation of the flanges at any point around the perimeter of the sealing surfaces. For symmetrical penetrations, the separation being in general less near the flange bolts

^~2 NUREG-1037 APPENDIX B

and greatest equidistant between the flange bolts. If the penetration is assumed to be without an elastomer seal, it is even possible to make a reasonably accurate prediction of gas leakage by analytically modeling the flange separations as one or more convergent / divergent nozzles in series releasing gas to the atmosphere.

The great difficulty appears when an attempt is made to calculate gas leakage with the presence of an elastomer seal between the flanges which fully or partially blocks the flow path. To even calculate with certainty whether or not a seal leaks at all is very difficult. Because such calculations are very difficult, and the results would most certainly be controversial, the problem appears to best lend itself to experimental methods which are not without controversy, but which can if done correctly, provide credible results.

Because the cost of testing full sized penetrations was prohibitive, it was decided to test two seal designs with full size cross-sectional dimensions, but of reduced circurr.ference or length. The idea being that if the cross-sectional dimension of the test model were full size and the circumference reduced, the leakage per inch of circumference would be essentially the s me as on a larger diameter full size penetration. Flange separation was incrementally increased using shims between the mating flanges. Equivalence of results between full size and reduced circumference incrementally separated models is compromised in the areas shown below:

1. Unlike the full sized penetration, the reduced circumference model employs flat relatively rigid metal surfaces which make the flange separations very nearly uniform around the sealed perimeters. The full size penetration, on the other hand, is relatively flexible so separation near the flange bolts is less than the separation at positions equidistant between bolts. This translates to constant leak areas versus variable leak areas for model and full size penetrations, respectively. Predicted rotation of full size flanges for " pressure seating" designs is similarly compromised by models not duplicating rotation.

l

^~

NUREG-1037 APPENDIX B

APPENDIX A*

l CONTAINMENT PENETRATION ELASTOMER SEAL TEST l

l l

1 1

0 Work supported by the U.S. Nuclear Regulatory Commission, Office of Nuclear Regulatory Research, under Ineragency Agreement 00E 40-550-75 with the U.S.

Department of Energy.

l l

^~I NUREG-1037 APPENDIX B 1

L

l l

CONTAINMENT PENETRATION ELASTOMER SEAL TEST INTRODUCTION l

Under the predicted extremes of nuclear reactor containment pressure associated with severe accidents, the mating metal surfaces of some

" pressure unseating" penetrations will separate, thereby partially removing the static precomoression on the elastomer seals between these surfaces.

Two seal designs similar to those employed by containment penetrations on the Surry and Peach Bottom nuclear power plants were leak tested as functions of flange separation and at containment pressures and temperatures approximating severe accident conditions. Both seal designs were found to be very leak resistant for the short time periods tested at all but the most extreme conditions of flange separation and operational temperatures. Flange separations required to produce incipient leakage at ambient temperatures greatly exceeded worst case severe accident predictions.

1

< PROBLEM DESCRIPTION i For nuclear reactor containment pressures less than the design pressure, separation of the " pressure unsenting" bolted flanges is impossible because the combined preload of the flange bolts is greater than the product of the design pressure and the area of the penetration over which this pressure acts. However, given that predicted severe accident pressures exceed design pressures by a factor of about 2.5, some " pressure ,

unseating" penetrations etr. loying bolted flanges have been shown to separate on the order of 0.010 to 0.030 in.

The average amount of flange separation is very easy to calculate given the principle dimensions of the penetration, number and size of flange bolts, and a peak severe accident pressure. With considerably more effort, one can even accurately calculate the separation of the flanges at ;

any point around the perimeter of the sealing surfaces. For symmetrical penetrations, the separation being in general less near the flange bolts NUREG-1037

^~2 APPENDIX B

and greatest equidistant between the flange bolts. If the penetration is assumed to be without an elastomer seal, it is even possible to make a reasonably accurate predtetion of gas leakage by analytically modeling the flange separations as one or more convergent / divergent nozzles in series releasing gas to the atmosphere.

The great difficulty appears when an attempt is made to calculate gas leakage with the presence of an elastomer seal between the flanges which fully or partially blocks the flow path. To even calculate with certainty whether or not a seal leaks at all is very difficult. Because such calculations are very difficult, and the results would most certainly be controversial, the problem appears to best lend itself to experimental methods which are not without controversy, but which can if done correctly, provide credible results.

Because the cost of testing full sized penetrations was prohibitive, it was decided to test two seal designs with full size cross-sectional dimensions, but of reduced circumference or length. The idea being that if the cross-sectional dimension of the test model were full size and the circumference reduced, the leakage per inch of circumference would be essentially the same as on a larger diameter full size penetration. Flange separation was incrementally increased using shims between the mating flanges. Equivalence of results between full size and reduced circumference incrementally separated models is compromised in the areas shown below:

1. Unlike the full sized penetration, the reduced circumference model employs flat relatively rigid metal surfaces which make the flange separations very nearly uniform around the sealed perimeters. The full size penetration, on the other hand, is relatively flexible so separation near the flange bolts is less than the separation at positions equidistant between bolts. This translates to constant leak areas versus variable leak areas for model and full size penetrations, respectively. Predicted rotation of full size flanges for " pressure seating" designs is similarly compromised by models not duplicating rotation. ,

^~ '

NUREG 1037 APPENDIX B

r -

2.

Given model and full size double concentric elastomer seal designs, the ratto of inner seal circumference to outer seal circumference in both cases approaches unity, but they are not equal. The smaller the models cir:umference, the greater the departure from unity.

3.

The incremental separation of the flanges used in these tests differs from the actual separation on full size penetrations which is a linear function of pressure above the limiting pressure where flange bolt preload is exceeded by pressure times penetration area forces.

Considering the above compromises, it was decided to perform a scoping test using a pair of 18 in.-900 lb ANSI reducing flanges (32 in. 0.0. x 4 in, thick with 20 bolts 1-7/8 in, diameter) to test two different seal designs each employing a different elastomer seal material. The plan was to first machine the flanges to retain double concentric neoprene, type W 0-Rings of 19.74 in. I.O. x 0.50 in, diameter and 22.74 in. I.D. x .50 in, diameter in trapezoidal 0-Ring grooves as schematically shown in Figure A-1. This configuration is used for the Surry PWR personnel air lock hatch, fuel transfer tubes, CR0 hatches, and equipment hatches.

After completing the first test series on the 0 Rings, the flanges were i

remachined, this time to duplicate the tongue-in groove configuration typical of the Peach Bottom equipment hatch and drywell head seals. Figure A-2 schematically depicts the double rectangular tongue-in groove Garlock Silicone compound No. 8364 seals.

These seals were retained in 19.5 in. I.D. and 22.5 in. I.D. grooves and were, before compressed installation, nominally 0.765 in, wide x 0.523 in, deep. All flange and seal dimensions were made to the approximate middle of the dimensional tolerance range specified on the engineering drawings of the full size commercial containment penetrations.

This was intentionally done because the results obtained here were known to be sensitive to dimensional variations involving the seal size and the mating metal surfaces.

NUREG 1037 A'4 APP M B

~

TEST APPARATUS Figure A-3 schematically shows the apparatus used to perform these tests.

The fully instrumented 18 in.-900 lb flanges (Item 1-Figure A-3) were heated electrically and pressurized by nitrogen from a pressure controlling valve (Item 11) which'in turn was supplied by a large nitrogen bottle trailer (Item 16). Three sizes of turbine flowmeters (Item 13) operating in parallel and individually throttled by manual valves (Item 12) allowed for measuring a broad range of steady gas flows which would leak from between flanges whose separation was maintained and controlled by six circumferentially placed 5 matched shim stacks. Thermal insulation placed around the flanges sped heatup time-and minimized thermal gradients within the approximately 1800 lb of flange material. All essential test data was simultaneously recorded on four dual channel 10 in, wide strip chart recorders. A visual flow indicator, not shown in Figure A-3, consisted of 1 in. wide x .002 in thick plastic strips hung from a round hoop around the outside of the insulated flange pair.

Though not providing a quantitative measurement of gas leakage from between the flanges, the plastic strips were capable of detecting flange leakage far below the ranges measurable with the smallest of the three turbine flowmeters.

TEST PROCEDURE

SUMMARY

.The test plan used is briefly summarized below. In general, the plan

-followed NRC approved model data sheets, each consisting of graphical plots (a family of curves) of estimated leakage versus vessel pressure as a function of flange separations of 0.000, 0.010, 0.020, and 0.030 in. For the neoprene

[

l. type W seals, ther were four such graphical plots to guide the experiment l progress for operational temperatures of 61, 248, 370, and 420 F. For the rectangular silicone tongue-in groove seals, four similar graphical plots were l used which called for flanged seal temperatures of 70, 475, 550, and 606 F.

Additionally, one test using the tongue-ingroove configuration was run as a base line for leak data with no elastomer seals in place. Because the seals were found to be more resistant to leakage than anticipated, much more data was taken for larger values of flange separation than originally planned.

A-5 APPENDIX B NUREG-1037

Figure A-1. 0-Ring configuration prepressurization assembly, o c ag c ri ,.r.si er r.ss ,4s.ti a. ir e,.o. ra o.

psig pstg

. O 8efore Asseanly. Seal tastattee os , '4*o.

, pstg

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- riange lecersoon

siero I

af ter Assemely $ eels Proc 3moressed ly o.QM f aca '

NUREG-1037 A-6 APPENDIX B

Figure A-2. Tongue-in groove prepressurization assembly.

ev o.

pstg

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e sig Flange leseration

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- .049 to .071 (Theoretical Average)

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F1 sage senaretton 19 30f0 After Allgely, $eall 8eeCompresses i

1 NUREG-1037 A-7 APPENDIX B 4

\

l Figure A-3. Schamatic diagram of experimental seal test apparatus.

1 l

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e 2 4 10

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t 18'a6* 900 lb AN$1 sold neck reducing flange pair alth f aces eachtred to retc.n 0.50 inch 0-Rtag seals la double trapesoldal grocees. These flanges were reedchtned to accept rectanjuler sillCone tw gue-la-groove seals after the 0 Ring seal tests mere complete. Ab built diernstons for both seal configurattens are shoen elsewhere in the test.

2 One of three strcumferenttally located dial tnascator gaugas measuring flange separatton to ulthin s .0005 tach. Flange separation =as controlled with sntes placed between the flanges; the dial todicathr gauges mere used to check the proper place.

eent and selection of the shles, integral Jack screus (31 were used to separate the flan 9es for shte insettlon and seal inspection.

3 1/4" bleed valve used to depressurtre the flanges before adjusting flange separatten 4 Electrical resistance heaters ehtch clamp to the flange circueferences. 46J V4 3/ power from the schematically deptCted temperature controller above and to the lef t of the flangen was used to control flange temserature sta a thermucouple feecbaca lint.

5 Dual pressure taps monitortng vessel pressure and the pressure between the seals lannulus) via bourdon tube gauges and electric pressure transducers tied to two t.f four dual channel 10* side strip chart recorders. The calibrated bourden gauges mere used only as a quick visual check of systes pressure and not as a source of test data.

6 Qual trernacouple taps monitoring vesselcentral body tesq>erature and annulus temperature. Calibrated thermocouples wrd elects scally tied to two of four 20* utde dual channel strip chart recorders.

7 8eeuntant bour&>n tbbe gauge placed close to pressure controller (11) for operator convenience.

8 1 1/2" ball bloc,k valve-4000 pst rating 9 l/4* manually operated bleed valve to atmosphere 10 pressure rettef valve to prevent overpressartnetton of flanges Il I 1/2* 2500 lb ANSI valve with attached pressure controller and air driven positioner.

12 One of three manually operated cc,eotnation turbine throttling and block valves used to select and control each of three turbine flow meters.

13 One of three turbine fiw meters of 1/2*. l'. and I 1/2= staes.

14 pressure andteaperature transdacer taps and transducers for monitoring turbine flow meter talet pressures and temperatures.

Transducers were connected to two of four dual channel 10* elde strip chart recorders for permanent recording of data.

15 2* block valve-integral to nitrogen trailer.

' 16 titrogen bottle seal trwca trailer-2000 pstg morning pressure utth apprestaately 40 tadt 4Juelly valved bottles each about 20 feet long.

l l

A-8 NUREG-1037 APPENDIX B

Flange separations for the neoprene, type W 0-Rings of up to 0.218 in were ultimately tested at ambient temperature. Similarly for the silicone seal tongue-in groove configuration, flange separations of up to 0.301 in. were tested.

Referring to Figure A-3, with valve (Item 15) open and one or more valves

'(Item 12) open, the heated flanges were pressurized with nitrogen in at least six stepped approximately equal increments of pressure between zero and at least 160 psig. At each pressure step, the visual flow indicators and the s2allest turbine flowmeter were checked for leakage. If leakage was observed, up to two turbine meters were monitored and recorded at once. By varying the nitrogen supply pressure from the trailer (Item 16) as measured at (Item 14),

rangeability of the turbine flowmeters could be adjusted to accurately measure both small and large leak rates. At the conclusion of each test, the electric power to the flange heaters was turned off, and the internal (vessel) pressure of the flanges was reduced to atmospheric pressure by opening bleed valves (Item 3) and (Item 9) and closing valves (Item 8) and (Item 12). Because the flanges were very rigid and the test pressures very low, only six of the twenty flange bolts were used.

Flange separation was varied by loosening the six flange bolts working through holes in the thermal insulation. Three circumferentially positioned Jack screws located radially outside the seal area were then used to separate the flanges so each of the six shim stacks could be changed to vary flange separation. Once the shims were placed, the jack screws were backed off until the upper flange rested atop the six shim stacks. The six flange bolts were then retorqued to a minimum value of 255 ft-lb to ensure that the bolt preload forces exceeded vessel pressure times area forces thereby precluding further flange separation in response to vessel pressure. Observation while testing of the dial indicators (Item 2) verified that flange separation did not change because of vessel pressures exceeding 200 psig. Shims were changed at flange temperatures up to 606 F using gloves.

l

^~9 APPENDIX B NUREG-1037

EXPECTED RESULTS 0-Rings Because the 0.502 in. diameter 0-Rings in 0.416 in. deep trapezoidal grooves protruded 0.C36 in. above the surface of the lower flange, leakage was anticipated for flange separations in the neighborhood of 0.086 in., the

" theoretical separation point." This geometry can be best understood by examining Figures 1 and 4. Leakage was thought to be certain for flange separations significantly exceeding 0.086 in.

Tongue-in-Groove Seals Refer to Figures 2 and 4. With the silicone seals initially below the surface of the lower flange by 0.069 to 0.071 in. and the tongues protruding below the flat surface of the upper flange by 0.320 to 0.321 in. , a

" theoretical flange separation point" of about 0.249 to 0.252 in. would in all probability define the approximate flange separation where incipient leakage would occur.

ACTUAL RESULTS 0-Ring Seals For ambient temperatures, flange separations were increased in 30 increments

-of flange separation between 0.000 and 0.188 in, before the first leakage was observed at a separation of 0.188 in. This was the point of incipient leakage; Figure A-5 shows maximum leakage was often observed at lower pressures. At higher pressures leakage often dropped off or stopped completely. For incipient leakage to occur at 0.188 in. separation is surprising; this'is far in excess of. 0.086 in. , the " theoretical separation point". Flange separation was next incrementally increased to a value of 0.218 in.; the seals again reseated at high pressure for each of the three separation increments. Leak flow for the 0.218 in. separation was not plotted

)

NUREG-1037 APPENDIX B

Figure A-4. Post pressurization.

Ps *0 te 100 pstg [ pg g, peng

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0.atag Seal post Pressurtsstion Configwetten

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f P, o to tactiy , pg. o anig esig s.s

!! C l [Flaage Seoaratton fein-Groove Post Pressurisattes Conftguretten I

i

^~ APPENDIX B NUREG-1037

I i

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Figure A-5. Double 0-ring seals in trapezoidal grooves.

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0 50 10 0 15 0 7

200 Vessel pressure (psig) o i

Figure A-6. Double rectangular silicone seals in tongue-in-groove configuration.

NUREG-1037 A-12 APPENDIX B f

15 .,

O - 0,020 " flage essereuen 0 - &Oa0 " Fisuse senseenen 4 - 8.040 " Flege sesseeHan O - &000 " Ff enge emnereMen A

v Al

'ltandard cutic feet er seat aisaster perr.0.

or ts.n day based usen 60*F and 14.7 to esta.

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10 3 **The flow rates shown here are estiested to be only accurate O within s 505 (see test).

O I

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0 Test Temperature = 420'F a- -s W 0 -

/ -

&s -

t 200 10 0 15 0 0 50 Vessel pressure (psig)

Figure A-7. Approximate ** seepage for double neoprene, type-W 0-ring seals per inch of inner seal circumference.

10 .

O - c,0 coa rienes meeseenen 0 - 0.020* Fleen essereMan A - 0.0 30 " Flage essereMan m

0 'Stancare cutic feet per say based upon 2o

~

50'F and 14.7 psia.

es V **The riov rates shewn here are estimated to De only e accurate attriin : 505 (see test), -

a 6 -

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I O 200 12 0 14 0 16 0 180 I 10 0 Vessel pressure (psig) :

I Figure A-8. Approximate ** seepage for double silicone tongue-in-groove seals per inch of inner seal circumference.

I A-13 APPENDIX B NUREG-1037

~ _

'because the data was erratic and flow was both time and pressure path I dependent.

These extremes of flange separation were far beyond the range of f practical interest relating to severe accidents. 1 i

' At all stated test temperatures below 420 F, no 0-Ring seal leakage was

' observed for flange separations of 0.000, 0.020, 0.030, 0.040, or 0.060.in.;

because there was no leakage, no plots were made for those temperatures.

However, . at 420*F, the flanges were observed to leak slowly. Leakage was so slow that measurement with a 1/2 in. turbine flowmeter was impossible and detection with the plastic visual flow indicators was difficult at all but the highest pressures near 180 psig.

t Though_ these tests were not designed to measure very low levels of leakage (seepage), a rough estimate of the magnitude of these leaks was made by repeated use of the universal gas law as applied to a pressure-time decay curve. Figure A-7 was generated. based upon calculating the mass m of gas y

in the known volume V of the system with the known gas temperature T and 1

pressure P1 by the universal gas law, P1 Y * *1RT3 . At a later known time when the. pressure had decayed to a level P2 with associated temperature-

'T 2, the mass m2 was again calculated using P2 V = m RT . Next, the 2 2 mass loss of gas during time period Attwas taken as(m) - m )2 = A'm and the universal gas law was used a third time as shown below to calculate the standard volume V3 that Am would occupy at 60 F (520 R) and 14.7 psia = P3 . With V and s At now known, PV 33 = AmRT 3 a rough estimate of the average or integrated leak rate occurring during the pressure decay between P 3 and P2 can be obtained by simply. dividing V s

by At to get the standard cubic feet per unit time lost due to leakage at the average pressure (P1 + P2 )/2.

The principle weakness of this calculation is with the loss of significant figures of accuracy when one subtracts values of m , and m which are I 1 2 i

l

. NUREG-1037 A-14 APPENDIX +B

nearly equal . A quick sensitivity analysis applied _to this data shows that these-numbers are in.some cases only-accurate.to within 50% because of the aforementioned loss of-significant figures.

These calculated values of seal seepage were detectable only at 420'F for the neoprene type W seals; hence they are only of significance for this

'particular material if PWR severe accident temperatures equal or exceed 420 F for' sufficient time to bake, crack, and deformation set the seal material las observed here. Though no planned aging was included in this test series, the 420 F test temperature did permanently deform the' 0-Rings to a cross-sectional shape resembling the capital letter D.

Tongue-in-Groove Seals Unlike the 0-Ring seals which did not-leak until flange separations greatly exceeded the " theoretical separation point", the tongue-in groove seals experienced incipient leakage at low pressures (~35 psig) at a flange separation of 0.213 in, which was less th'an the 0.249 to 0.252 in.

" Theoretical separation point." Figure A-6'for the tongue-in groove seals is

.similar to the 0-Ring seals ambient ~ temperature leakage as shown in Figure A-5 in that maximum seal leakage most often occurred at lower pressures and in many cases slowed or stopped at higher pressures.

' For the silicone seals no leakage was observed for flange separations of

'0.000, 0.020, and 0.030 and 0.060 in at test temperatures of 475 F and 550'F. For this reason no plots of leakage were made for these test temperatures.

At 606*F the silicone seals were observed to leak slowly (seepage)in much the same way the neoprene seals did at 420 F. The seepage was calculated using the same method described earlier. It is significant to note, however, that no seepage was observed below 164 psig as shown in Figure A-8. The silicone seals did resist leakage very well considering that the two seals

. melted together and were later removed as one. These pressure tests were of i

A-15 APPENDIX B NUREG-1037

A short duration; for long periods of time, a temperature of 606 F is likely to result- in gross leakage of the magnitude described by Figure A-9 for the case of tongue-in groove design without seals in place. Even the smallest of

(

temperature increases above 606 F for silicone seals o'f this type are likely '

to result in gross leakage. Again, the significance of this leakage is strongly dependent upo'n how closely BWR severe accident temperatures approach 606*F and remainlat these temperatures long enough to heat the seal material itself to these temperatures. Thermal lag time and steady state = thermal gradients are expected to provide both significant and dependable assurance that penetrations employing these designs and materials survive all but the most extreme of severe accidents' extending over.the longest time periods.

CONCLUSION For the seal designs and materials tested here, flange separation resulting from extremes of severe accident related pressures is not likely in itself to be the source of significant containment leakage. Extremes of temperature lasting over a sufficiently long period of time are however, a significant factor that may contribute to gross containment leakage. These conclusions may not be applicable to seal materials and designs differing from those tested here.

O NUREG-1037 ^'

APPENDIX B

l 2000 , , ,

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**The accuracy of' theos estats cettested to be witete till, O

6 1000 - -

a h eeO 500 a/ .%

= #

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Oc " "'~~ ' '

O 50 10 0 150 200 Vessel pressure (psig)

Figure A-9 Leakage for double tongue-in-groove design without seals.

(

A-17 APPENDIX B NUREG-1037 l

l 1 l

FIGURES Figure 1. 0-Ring configuration prepressurization assembly.

I Figure 2. Tongue-in groove prepressurization assembly.

Figure 3. , Schematic diagram of experimental seal test apparatus.

Figure 4. Post pressurization.

Figure 5. Double 0-ring seals in trapezoidal grooves.

Figure 6.

Double rectangular stitcone seals in tongue-in groove configuration.

Figure 7.

Approximate ** seepage for double neoprene, type-W 0-Ring seals per inch of inner seal circumference.

Figure 8.

Approximate ** seepage for double silicone tongue-in groove seals per inch of inner seal circumference.

Figure 9 Leakage for double tongue-in groove design without seals.

NUREG-1037 A-18 APPENDIX B

i APPENDIX B ZION CONTAINMENT LEAK AREA ESTIMATES B-1 APPENDIX B NUREG-1037

ZION PERSONNEL AIRLOCK ANALYSIS The purpose of this analysis is to determine the leak area estimates for Zion containment personnel airlock. An analysis of the exterior bulk- I head was performed using the computer code SAP IV. The bulkhead vertical and horizontal plane of symmetry were taken advantage of, allowing the use of a quarter model of the bulkhead. Figure 1 shows the computer model of the barrel and bulkhead. Figure 2 shows the bulkhead door elements.

The barrel plate elements are .50 inch thick and the bulkhead and door plate is 1.0 inch thick. The barrel has a 61 inch radius. The door opening 4 ft wide by 6 ft 6 in. tall. The bulkhead door and bulkhead stiffness properties were obtained from Reference 1. Figure 3 is an isometric computer plot of the composite door and bulkhead model. Stiff truss elements were placed between the door and door frome nodes. Should these elements go in tension indicating separation, the elements were removed.

For this model the truss elements between nodes 51-69 and 50-67 were removed The beam member input properties are summarized below.

Beam Axial Major Moment Minor Moment Elements Area (in.2) of Inertia (in.4) of Inertia (in.4) 1-2, 16-17 35.0 1109 62.9 3-7 26.5 751 17.8 8-11 2.6 50.7 .17 12-15 5.2 101.5 .34 18-27 17.5 277.4 30.4 28-33 1.2 13.0 .03 34-37 8.0 194.4 .67 38-41 4.0 97.2 .33 The input pressure loading for the computer analysis is 50 psi.

The displacements at the top of the door are:

.NUREG-1037 B-2 APPENDIX B

--mr a d

TLB2ed.1 40 39 38 37 36 35 34 33 32 31

@l@ @i@ @ @ @ @ @ !~ Z ==6.0"*

30i 29 28i 27[ 26 2 55 24 23 22 215 ~

20

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19 18 i 17 16' -15 14

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13 12

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5 4 3 2 10 9 8 1 Barrel Plate Elements ,, 2, 23 .

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/j (17) . (16) (2) (t)

I '

,6,f . . .. . .. 4 5. .. . .,0) .. . . . . . . 5 3. . . . . . . . . .'. . . . . . . j 5.

/i kg) D. i;14 b .59

5

~

27 '/ v . 4 6 63 Elements

[13) . (14) ;

[15]

/(12) k ' .h h ! 39 [3]

% v -

/

$i '47 2 55 60

]64 [] - Beam Elenients

@ ',@' @ j @ (6)

'a,,,,2,56,,,,,,,,,,,63,,,,,,,,,,,,,,,,,, ,3 2 9 [-- . - --- - . -. I -

9

'. . v 9 ,

Ei @- ![7]

49 5 62 66 30 (6) (9) [10) (11)

Bulk Head Beam & Plate Elements l

Figure 81 B-3 NUREG-1037 APPENDIX B

=

-- _- =. . . . _ . .. - _ . . .-,-- . . . . . . . . _ _ _ . . . .

+

71 70 E9 68 67 (21 [20}' [19) (1Bl

[22] , (k .

[28]

72 8' 80- 79 78

{23]' i )'

~

[29) 73 82 89 88

[24] .

i 74 (30]

83 9C '-- 95 94

'[34) (35: (36) [37]

(2S]

[31}

75 84 91 S6 99

[26: -

[32]

76 85 92 97

, 100 (27)

{b' '

[33]

i 77 (38]- 86 l [39] 93 [40] 98 [4 t] '101 Door Elements TL8884-2 l

' Figure B2

{

l i

B-4 NUREG-1037' APPENDIX B

Node 50 AZ = .079803 in.

-AZ = .068149 in,

-Node 67 separation = .011654 in.

Node 51 ' AZ = .075049 in.

Node 69 AZ =^.066032 in.

2 ,

s

-separation = '.009017 in.

The leak area for a 50 psi loading is

(.009017+0) 12 , ('.009017+ .011654) 12' = .71 in.2

=4 2 2 .,

The maximum stress occurs at top center of the door frame, beam element #1..

6 M2 = 2.15 x 10 in.-lbs ,

The corresponding bending' stress is given by y , M_C _ 2.15x106 (7.7)~ - 14,923 psi

I (1109.3)

Using the material minimum yield strength of 32,000 psi, yielding will be

) 50 = 107 psi. The leak area estimate at 107 psi

- initiated at I '

is (.71) = 1.52 in.2 increasing the rate of leak area to pressure by a factor of 10 above 107 psi pressure loading'results in the following leak

~

area at 134 psi L.A. = 10 ) (134-107) + 1.52 = 5.36 in.2 The leak area as.a function of pressure is as shown by the following figure.

NUREG-1037 B-5 APPENDIX B r

s x

N NQ N %

s t

~

x N

BEAM ELEMENTS VERTICAL CENTERLINE Figure 83. Computer model/ isometric plot.

I NUREG-1037 B-6 APPENDIX B l

, .-. ._ .. ~ .-._ __ .-.

t i

i-l l

i l

L l

i f

t s

I t

E - - - -

l 5.36 - ~ ~ - - - - - - - - - - - - - -

k A

N.

C

! c .

.m -

G u

i m .

t J .1.52 - - - - - - - - - - - - - - - - - - - - - -

l' 0.00 l 107.0 134.0 O.0 I Pressure (psi) no.. . ,

1 l

Figure B4. Leak area as a function of pressure.

l B-7 NUREG-1037 APPENDIX B

N ' I

- Page B-8 contains Zion personnel. airlock bulkhead analysis microfiche. It has been omitted.from this. appendix but is.'.available from

LMr. Goutam Bagchi Office of Nuclear. Reactor Regulation ..

! U.S.1 Nuclear Regulatory Commission' ,

, Washington, D.C. 20555- '

1.

4 i

s.

4 ,e f; 4

~

4 i

a

+

s' :.

~l i h.

I i

(v 1

i

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4 4

1

.I NUREG-1037 - B-8

. APPENDIX B J

f f

,.-i . w$..-- ,- - , - . ..,,,.e. ~ . _ , . - - _ . . - , , - . ~ _ . . . . - - - ..-,.,.,,-_.--_..,4 -

~,...,m----m._----._,,., .-.w...,,-m.b. .m . . ~..

0-7 APPENDIX c SURRY LEAX AREA ESTIMATES C NUREG-1037 APPENDIX 8

I ANALYSIS OF SURRY EMERGENCY ESCAPE AIRLOCK BULKHEAD The purpose of this analysis is to determine the leak area as a

[ function of pressure for Surry's Emergency Escape Airlock. The emergency escape airlock is flange bolted to the containment equipment hatch. The I equipment hatch flange is in compression for containment internal l pressure. This ensures metal-to-metal flange contact and precludes i

l significant leakage between flanges. By inspection and engineering judgement, this compression type seal (double 0-ring) will not leak with increasing containment pressure until thermal degradation of the neoprene 0-rings occurs. The flange connection between the escape airlock and the equipment hatch is a tension-type flange. The leak area of this flange is

! calculated in Appendix B of this report. Buckling of the equipment hatch is not a problem because of the area reinforcement requirements for the airlock opening. The airlock consists of a barrel section with a bulkhead at each end. The inner and outer bulkheads are identical designs with 30 in. diameter doors. These doors are both internal, thus closed by containment internal pressure. The airlock barrel has a 1-1/2 in. thick barrel at the internal end and 3/8 in. thick barrel at the external end.

The bulkheads are constructed of 7/8 in. thick plate with two vertical support beams and a circular . door frame.

Since the bulkheads are flat, they are the most likely to fail or leak first in the airlock design. An analysis of the bulkheads was performed l using a SAP IV computer model of 1/4 of the bulkheads taking advantage of j symmetry about both the Model X and Z Axes. The model node point locations are shown in the following figure. The encircled numbers are plate element numbers. The bulkhead stiffener beam goes through Nodes 1, 15, 27, 39, 51, 74. The beam elements representing the door frame go through Nodes 1 through 10. This analysis was performed using the interior barrel thickness of 1-1/2 in.

L l

C-2 NUREG-1037 APPENDIX B

-J b .

t &

,,' s k ^

, ,% ' e

@. ~'

e, 5:

to 4

+

g

'8J

- ,.+,

33 Jr tob. '* #o 32 to so ,

o d

q**-

,0 , , o, <, .

+,

@ 5 s ,,e

^ s

- x @ 's *,

x.. .iae

- cess

@5 3, L , ,,.

3

,. 3

LO_bj.O N. .

6 ,O

-. ,.c.

Buik Head 7 i i i 1 1 1 sss

y ; .g g 9 ? ?- 2 9 *

1 3

,e e: e. se se e0 er s. is se v.e'0

.,T ,

e.. ' ..

,0 ,,

. n . n. ,... ,, ,. j Q, 1.. 6.. .. ... .... .... .. ,,

i ,,@..

l I . . . ,

18 8 700 i Y =-20.0

'O' 102 10 3 10 . 'OS 'O. to, 108 10 9 110 Barre!

Q- Plate Dement No.s t

Figure C-1. Computer model of the exterior bulkhead. ,

-NUREG-1037 C-3 APPENDIX B

4

'0 4

e The door frame beam properties are:

.a

-7" r 9

,7/8 A=6.13 ~

n a 7/8 * - --

1 4- ,

A=3.5 a

U *-

.b A =7.5 2.5" n

3"

~

Y

6.13(4.44) + 3.5(2.0) 6.13 - 7.5(1.25)

+ 3.5 + 7.5 = 1.45 in.

( 0.875)(4)3 -

+

(2.5)3(3) 2 2 2 Ix == 12 -12 + 6.13(2.99) +'3.5(0.55) + 7.5(2.70) = 119.11 in.4

'At = 11.0 Ay = 6.'45 Ai = 7.5 in.2 Iy" (0.875)3(4)

12. + 12(3)3(2.5)

= 5.85 in.4 3

I =7 bd ,(0.875)3(4) (2.5)3(3) = 16.5 in.4 p 3 3 3 i

C .

NUREG 1037' APPENDIX B

The. bulkhead support beam properties are:

!: ' 14" = ,

u 7/8-[l A=12.25 il

~

7/8 -

5.94" N

a b b o y 7/8 . A=5.25 o

6" :

l Y , (12.25) 6.38 + 5.2(2.97) - 5.25(0.44) = 4.02 in*

(12.25 + 5.2 + 5.25)

x , (0.875)3(6) 12

, (5.94)3(0.875) 12

+ (12.25)(2.36)2

+ (5.20)(1.05)2 + 5.25(4.46)2 = 194.01 in.4

.AT = 10.44 Ax = 5.25 A = 6.73 in.

y

y , (0.875)35.94 12 , (6)30.875 = 16.08 in.4 12 g ,' (0.875)35.94 , (0.875)3(6) = 2.'67 in.4 -

3 3 C-5 NUREG-1037 APPENDIX B t_

4 A pr:ssure. loading of 50 psi external to the barrel and bulkhead tas imposed'on the bulkhead model. A concentrated force of -442 lb in the

~

, y direction was placed at~ nodes 1 and 11, and -884 lb at nodes 2-10 to represent pressure loading transmitted by the door (also 50 psi).

Evaluation of SAP IV Results-In the door frame beam elements 1 through 10, the maximum bending

. moment occurs at the J-end of beam 10.

4 M(2) = 6.89 x 10 in.-lb .

The associated b'ending stress is 4

i- 0 (3.95) = 2285.10 psi a=f=6.89 .

At-130 psi, I3 o = g0 (2285.10) = 5.941 osi; no^ problem.

For the bulkhead stiffener, beam elements 11 through 15, the maximum bending moment occurs at element 11, J-end 5

M(2) = 4.86 x 10 in.-lb 5

MC o = T , 4.86 x 194.0 10 (4.895) = 12,262.7 psi .

At'130 psi, o= 30 (12,262.7) = 31.883 osi; no problem.

9 4

i C-6 NUREG-1037 APPENDIX B

'P For the bulkh;ad plate elements 1 through _60, the maximum' b nding occurs at element 5 4

M xx

= -0.183 x 10 jID '

4

e. 6(0.l83 x 10 ) = 14,341 psi .

t (0.875)

For a pressure loading of 130 psi, o= (14.34i)=37.387osi .

~

This is slightly above yield, but still poses no real problem, since it is a_ secondary stress and the plastic hinge moment has not been exceeded.

Based on these stress results, it is apparent the ultimate capacity

pressure of the emergency airlock exceeds the containment ultimate capacity pressure of 134 psig.

EVALUATION OF 000R FRAME DEFLECTIONS

~

For a 50 psi pressure loading and a uniform loading around the door frame:to acco'unt for the pressure loading transferred by the door, the door frame deflections are as follows:

A i

C-7 NUREG-1037 APPENDIX B.

'ay Node (in.)

1 0.044 2 0.044 3 0.044 4 0.045 5 0.046 l 6 0.048 7 0.049 8 0.051 9 0.052 10- 0.053 11 0.053

.For this loading, the maximum leak width is 0.053 - 0.044 = 0.009 in.,

assuming a perfectly rigid door. This implies the door pressure loading is

!- not uniform and must be greater nearer the stiffener beams. Another SAP IV load case was run with the door pressure loading on the first 3 nodes only, as follows:

y force Node (lb) 1 4420 2 2946 3 1473 The door frame deflected shape then became Node ay 1

1 0.045

! 2 0.045 l 3 0.045 -

l 4 0.045 0.045 A maximum difference

-5

. 6 0.045 of only 0.003 in.

7 0.046 8 0.047 9 0.047 10 0.048 11 0.048 ,

C-8 NUREG-1037 APPENDIX B

' Bas:d on these results no diffcrcntial displacement bttwetn tha door and j

' oor frame deflected shapes.is predicted. -'This could be proved by d ;

' overlaying a model of the door on the door frame of the existing model with

- c'ompression only elements between the door.and door frame. The current :

schedule does not permit such an extensive analysis at this time.nor is it j felt justified.

i i

l l

r C-9 NUREG-1037 .

APPENDIX B

,.r a y,R 4m I

I E 4

)~

4,.

Page C-10 contains Surry emergency escape' airlock analysis microfiche. It has

- been omitted.from this appendix but is avai.lable from:

Mr. Goutam Bagchi

'. Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission Washington, D.C. 20555 i

s

)

1.

4-p ljuREG-1037 C-10

, APPENDIX B

}

l g a l,l . h. .M d'M s

3 h

Page C-10 contains .Surry emergency escape airlock analysis microfiche. It has f been omitted from this appendix but is available from:

Mr. Goutam Bagchi

[: Office of Nuclear Reactor. Regulation b

U.S. Nuclear Regulatory Commission Washington, D.C. 20555-

)

i 1

f I

NUREG-1037 C-10 APPENDIX B

EQUIPMENT HATCH EMERGENCY AIRLOCK FLANGE SEAL ~.

'All sealsLon the hatch, except one, are compressed with increasing -

containment pressure. The exception is the seal between the equipment

~

. -hatch insert neck and the' escap'e -lock door _ tube. This connection is put in tension with increasing containment internal pressure.

The seal design includes a double 0-ring made from silicon-rubber and 32 1-in. swing bolts, which are torqued to 160.ft-lb when the containment

-is closed up. .The pressure'at which.the flange surfaces of the connection <

just begin to separate will be examined first:

Total bolt area = A b

= 32 x "I I = 25.13 in.2 bolt length = 5.0 in.

bolt stiffness = K j=f=25.1 5 E6) = 135.7E6 bolt'preload = h = } = 9600 lbs/ bolt 2 )

. . Total preload = 32 x 9600 = 3.07E5 lbs Flange Area = id(w) = w(72)(1.5) = 340 in.2.

Flange thickness =-5.0 in..

( } = 183.6E6 Flange' stiffness K2* 5 l1 The applied-load ~ required,to produce flange separation is p ,(1+ 2) p , (135. + 1836) 3.07E5 = 3.30E5 lbs l

C-ll NUREG-1037 APPENDIX B t

The seal diameter 11s'72 in.- The pressure area is- - '

~

^

.[(d)2=0.785(72)2=.4071in.2 , ,

y ~The pressure. required for flange separation is P

p?.' S. = 3.30E5 lbs = 81.06 psig .

-A 4071 in.

For-pressures abov'e 81.06 psi the flange separation is-given'by the

additional bolt = stretch.'

s

, , [(P(4071) ~3.30ES)]

K.3 The corresponding leak area is given by

.LA = A (72) .

Fdr. pressures of 100 psig ~and 119 psig the flarige separations and: leak -

careas are:

~

p , Leak-Area

, '(psig) (in.) (in.2)

'100 0.0006 3 -

0.13 119 4

0.0011 0.26 a

V C-12 NUREG-1037 APPENDIX B

, , - .- ,e -..-n--.y.--

ww,,..--, ,,_p. -,,,.,s.m,.,es,_ y, ,,w,-. .-y,yiv.ra% ,,,,,-wm-,.a,. 2-,-r ---t- -ws---t- - -

PERSONNEL AIRLOCK SEAL Personnel Airlock The personnel lock inner door ~is sealed by containment pressure. For determining the upper bound leak area' the inner door is assumed open or failed and the full containment. pressure is placed on the outer personnel lock door. Seals on the outer door are assumed failed. This loading tends to separate the door and flange causing possible leak paths.

Twenty-four tabs on the door and 24 tabs on the door flange interlock when the hatch is pressed against the flange and rotated. This results in full compression of the seal, but very little, if any, straining of the flange. In other words, the door and flange are just touching.

Three series springs are used to approximate the flange separation distance as a function of containment pressure. The first spring approximates ~the compressive stiffness of the 24 door. tabs. The second spring models the stiffness of the 24 cantilevered flange tabs; the third represents the axial stiffness of the cylindrical portion of the flange.

t l

C-13 -

4 NUREG-1037 APPENDIX B

, .-- ,-r n -.-.-,,-,-,,-,-,en, ,.,,ar-- -- e-+-- -

g

. r i

.r e

n

,i L

s n . s i

o l s i n a e t m e i d d D 1 I

b a.

. 6" a s e

T

. 8 l n ..l: a k c c i

ae i p o dmgs y g - l r

"4 eprrin T i a

/o- e r

3 e0 u .

. l e

,N s -

n 2 s -

n e

r .

o P ~

s r

e p

y

\ rs Q 8 I

~

1

= 4 _,

,^-

s. S r

r u

2 C

- , e r

3 u 7 g 1 i

- F hr tco e ao gb HD no laT f

o zC:S.NowN

pr oomzop cm

l ' ~ (f =^ ~

e 4

J A essur.

l <

I 1.

- g Co.c t.b compr.ssi.n

-]-

.~

gJna <c. ..r) .. sid . -

.l c

hg c nneiem ei e s.ce: ... :

e'If.'!".1'il'.TC.*?.'*.**."o'.';'c',*4?.*,".'

l v. ri.a,.-%, n s ,.. m>

tabs = 24 x '[1.75(5.75)] = 2.41.5 in'.2 A

.A hatch. = x(44) = 6082 in.2 F pressure " b atch = 6082 P ,

0.125 - a F = EA e = E A seal s ss 0./5

'A seal =. 2 x [2s(44)(0.718)] = 397 in.2

' ~

E3 ,,j = 900 psi NUREG-1037 APPENDIX B

1 .

AL i

seal " 3 76

0.125 cj = g = 0.167 i

Fg = (0.167)(900)(397) = 59550 lb .

Force of seal + force of pressure = force built up in flange-tab deflections. Find the force in the tabs as a function of pressure:

spring kg f = 241.5(27.E6) = 3.26025E9 lb/in.

l f- spring k(3) f = 3.25 (46)] = 6.3405E9 lb/in.

i spring k(2) cantilever, uniform load Ymax " - I= = 3.83 in.4 W = 1 lb L = 1.75 in.

l -E = 27.E6 y = 0.6E-8 in./lb max k

j=Ymax

1.544E8 for one tab ktotal = 24kj = 3.704E9 lb/in.

C-16 NUREG-1037 APPENDIX B 1

( . . . . _-.- - - - - - - - - - - - - - - - - - ~ ~ - ~ ~ ~ ~ - ~ ~ ~ ' ~ ~ ' ~

1 F o. k,qa-j - = 1.361E9 lb/in.

k,q = j ,

3.260E9 6.3405E9 3.704E9 F = ka F = 1.361E9a

-i F

p

+F s =F tabs. flange 6082 P + (59550 - 476400a) = 1.361E9a

' " 60821.3615E9 P + 59550 A

y leak p

(psig) (in.) (in.2) 30 0.0001777 0.0514 44 0.0002402 0.069 100 0.0004900 0.142 119 0.0005753 0.166 f

i l^

i 1

C-17 APPENDIX B

NUREG-1037

F APPENDIX D PEACH BOTTOM LEAK AREA ESTIMATES 1

NUREG-1037 APPENDIX B D-1

~

PERSONNEL AIR LOCK ANALYSIS

.The purpose of this analysis is to determine the leak area of the

. Peach Bottom personnel lock as a function of containment pressure. This personnel air lock has a flanged barrel, 8 f t-8 in. in diameter by 14 ft long with bulkheads at both ends. The flange connecting the air -lock to

.the containment. flange is 'such that the bolts are subjected to increased tension loading with increased containment pressure. The leak area associated with this flange connection is calculated separately. Each of the bulkheads have a personnel door which is on the containnent side of the bulkheads, thus containment pressure tends to close them. For this air lock the bulkheads are considerably different. The inner bulkhead has ,

24 radial stiffeners between the barrel and the' door frame. These stiffeners are 1 in, thick plates tapered from 24 in. deep at the barrel to 6-3/8 in. at the door frame. The barrel at the inner bulkhead is 2-1/4 in, thick. Due to complexity of the inner bulkhead, the current schedule does not provide sufficient time to develop a model of it. This bulkhead appears to be much stronger than the exterior bulkhead, due to both the number of radial stiffeners and the thickness of the barrel. A detailed analysis-of this bulkhead should be performed at a later date.

.The bulkhead is a 7/8 in, thick plate stiffened by a rectangular 30 in. by 72:in door frame which extends to the air lock barrel at the top and. bottom of the door frame. In addition the bulkhead has 3 horizontal stiffener plates between the sides of the door frame and the barrel. The barrel thickness at the exterior bulkhead is 3/8 in thick.

The exterior bulkhead was analyzed using a 1/4 model with a pressure Lloading of 50 psi. The following figure shows node point locations and element numbers of the model. The node number coordinates can be obtained from the included computer run.

0-2 NUREG-1037 APPENDIX B

~ s Tietta.

43 39 38 37 36 35 34 33 32 3

~'"*'"'

28 27 26 25 24 23 22 2r - Z=0.0" 30 29e 20 19 18' 17 16 15 14 13 12 11

- Z=-12.0"

@ @ @ @ @ @ @ @2@ t

- =-24.0**

10 9 8 7 6 5 4 3 Barrel Plate Elements . ,, 2, 23 . :

l43 . . . . . . 4;42- - . .. . .. d 41 24 l ! ! !

!44 *52 ! 51

$0 25 [1]

(17] { (16) (2)

2. ..."...,.45............53...........'...... . ..

i@!l54 @ l

!S9

@ t4)

Plate 46

[14] (IS}

8 -C Dements (12} i. (13) i .

@ '@i @ ! @ c5i

= . . . .. . . . . p.. ..i n .. . . . . . . . . j .6.e .4

. .g. . _o,,,,a"

@ j@ @ j @ t.i 2, . ..... . .. .. t4....ta..........ie................. .5 t'i

@ l@! . :

62 149 57 66 30 (8) (9) (10] (11]

Bulk Head Beam & Plate Elements Figure Dl. SAP IV model of the bulkhead.

NUREG-1037 APPENDIX B

.t

~

h Beam Properties Door Frame (Beam elements 1-7) .

a- -

sr - r- _y y I !Y Y

o-2.S' h

h r i~ u ....

~

y . ~-(15)(1.25) + (12)(6) + (7)(5.94) = 2.79 in.

(15 + 12 + /)

Ix =y+ 6(2f)3 3

+ (15)(4.04)2 + (12)(3.21)2 + 7(3.15)2 = 589.7 in.4 3

I = 2.5(6)3 +

j 46.0 in.

J= = 6(2.5)3 ,12 1)3 35.25 in.4

'A T = 27 in.2 x = 15 in.2 D-4 NUREG-1037 APPENDIN 8

q

1

~

- tA

+Z

= 12 in.2" Bulkhead Stiffener Beams f J

54W ,r s** ,

y m ,

o v. ,

L l

i.-

)I y , 5.25(3) - (14ll(44) = 0.50 in*

(5.25 + 14J Ix =j f + 5.25(2.5)2 , (j4)(g,94)2 = 60.9 in.4

. A, = 0 Ay t = 5.25 'in.2

.A T = 5.25 in.2 I

y

= 6 (0.875)3 12

= 0.33 in.4

.i D NUREG-1037 APPENDIX B

f a . 6(0j 875)I= 1.34 in.#

Beam elements 12-15 used these values. l Beam elements 8-11 used 1/2 of these values because of symetry.

Door Frame Extension Properties

.(Beam elements 16 and 17) a r .- =;

t tS-n l

S. S" 7

! he!

p ..- =l T u ..

Y ,15(6.75) + 5.5(2.75) - 14(0.44) = 3.19 in*

14 + 15 + 5.5 I = (2. (6) , (5.

3(1) + 15(3.56)2 + 5.5(0.44)2 + 14(3.63)2 = .

y y

, (6) (2.5) + 5.5(1)3 = 45.46 in.4 12 12 3

g , (2.5? (6) , (1)35.5 = 33.08 in.4 3 3 0-6 NUREG-1037 APPENDIX B

AT-= 20.5 in.2 Ax = 15.0 in.2 ;

+A y = 5.5 in.2 EVALUATION OF RESULTS The SAP IV Model was run with door pressure loads distributed uniformly around the door frame and the results showed a separation between the door and door f rame. The model was then rerun with the loading redistributed to account for the separation.

i The maximum bending moment occurred at the top and bottom of the door frame at tha center. This was beam element [1] i-end, for which 6

M(2) = 1.26 x 10 in.-lb .

Thebendingstressisa=f 6

, 1.23 x 10 (9.21) = 19,210 psi 589.7 (for50psipressure).

Yielding will be initiated at P= (50) = 93.7 osi .

D-7 NUREG-1037 APPENDIX B ,

l

The door frame deflections are:

Node Az ' Relative to Side 52 O.0625 0.0000-51 0.0688 0.0063 50 0.0705 0.0080

~

These values must be evaluated considering the door top deflections for-a 50 psi pressure loading. ,.

y f , , . . . ~ ~ .

N l , . 1 . .h rar nWir s.-

l=

2 2

4(x) = Wx(5t2 , 4x )

480Elt W=hx30=11,250lb-l t = 30 in. t s

6 E = 29 x 10 I

D-8 NUREG-1037 APPENDIX B

I Top of Door Properties o

vr > .-u M '" \ _.,

N E= ,

_ (7.19) 7jI9 + 1.5(0.25)

Y* = 3.02 in.

(7.19 + 1.5)

I x = 3.0(0.5)3

]

+ j(7 39)3 + 1.5(2.77)2 + 7.19(0.58)2 = 44.93 in.4 a @ x = 8 is 0.00288 in. (node 51) a @ x = 15 is 0.00389 in. (node 50)

Separation @ node 50 is 0.0080 - 0.00389 = 0.00416 in.

- @ node 51 is 0.00630 - 0.00288 = 0.00342 in.

For 50 psi, the leak area is:

4 0.00342(8) (0.00342 + 0.00416) (7) ,o,33 ,,,2 l

Extrapolating this leak area linearly to the initiation of yield in the ,

! door frame yields a leak area of l

9* (0.16) = 0.30 in.2 at 93.7 psi 5

D-9 NUREG-1037 APPENDIX B

i

~

.Beyond yield,'the leak rate will be assumed to increase with pressure 10 times as fast as before yield. This assumption requires further justification.

At 150 psi' the leak area is lo Il50;793 93 7) (0.30) + o.30 2.10 in.2 The leak area as a function of pressure is t,. _ . . . _ . _ _ . . . . . . _ - _ _ _ . . .

i T

i a

1 I i no-."E)

Since this leak area is based on the inner door being open and loss of seal at the outer door, this leak area should be used as an upper bound leak area only.

l 1 :

l i

! D-10 NUREG-1037 APPENDIX B i

' ~

[i; _

i

^

.c

Pages D-11 through D-31 contain Peach Bottom personnel airlock bulkhead analysis :

, computer printout; These pages have been omitted from this appendix but are

. available from:

~

-Mr.EGoutam Bagchi- .

Office.of Nuclear Reactor Regulation '

-U.S. Nuclear Regulatory Commission Washington, D.C. 20555 ,

s , ,

s

P 1

4 1

14 I

)

4 w

d b

e NUREG-1037 D D-31 APPENDIX B y

I DRYWELL HEAD SEAL l

Peach Bottom Drywell Head The drywell head has a 32 f t diameter, double seal in its flanged connection. Sixty-eight 2-1/2-in., bolts of A320-L7 steel hold the flange in place. Flange separation pressure, and leak area as a function of drywell pressure will be examined.

pp.96-101. Basically the method states that the flange bolt load is equal to a bolt-flange stiffness ratio times the bolt load; then the preload is added on.

For the bolt, N

1 P+F Fj= g1g2 0 where F)

=

total bolt' load K)

=

stiffness of bolts K

2

= stiffness of flange P = applied load (from pressure)

F, = bolt preload.

For the flange, D-32 NUREG-1037 APPENDIX B

p

~

K '

2 P-F, F2" K1+K2 where-F 2

= total flange load.

K j =h=68(w/4 )2(27.5E6) = 2.46E8 lb/in.

K 2

= AEC , w[(195.5)2 , (jg4)2(27.3E6) 27.31

= 1.835E9 lb/in.

K 1 2.46E8 K) + K2

2.46E8 + 1.835E9 = 0.118 K

2

= 0.882 K1+K2

. pressure load on bolts = pA head =p(w194)=p(1.182ES) preload in bolts, torqued to 850 ft-lb. . .

I2)

F, = h = ( (g 5) = 40,800 lb/ bolt 6

F, = 40,800(68) = 2.77 x 10 lb, total Flange separation occurs when F2=0 .

F2"K j K P - F, = 0 2

D-33 NUREG-1037 APPENDIX 8

~

, (Kj '+ K2)

P=F, g 1-0 P = (2.77 x 10 ) (2.46E8 i +

. 1.835E9)

m. '= 3.141 x 106 lb 3.

p=fa 82E

" 20*0 DSIQ After flange separation, 1.182E5 (p - 26.6) = k jx i where x = flange separation, in.

Leak area is given by AL = Cx L

where A

g

leak area, in.2 C

circumference of flange (inner edge) in.

so that l.

C 1.182E5 p - 26.6) 1 l'

1.182ES(p - 26.6) i Ag = 2 (194) 2.46 x 10 l

i AL = 0.5857 p - 15.579 NUREG-1037 D-34 APPENDIX B I

This area / pressure relationship is based on no sealing action. It represents an upper bound on the leak area if sealing action is at all effective.

Leak Area (in.2)

Pressure (psia) 0 0.00 26.6 0.00 78 30.20 116 52.36 126= 58.18 1 31 61.1 5 142 67.94 160 78.13 Structural failure of'the head seal is controlled by bolt yield. The containment pressure necessary to yield the bolts is given by:

A p=Sb {b where S = stress in the bolts, psi b

p = containment pressure, psig D-35 NUREG-1037 APPENDIX B

A = pressure" area of the drywell head (1.182E5 in.2)

\

= net section area of the bolts.

For an A-320-L7 steel, the yield stress is 105,000 psi from the current ASME code.

p , (105000)[ /4(21/2)2(68)] '

T.15zt3 p = 297 psi Structural failure is not likely. .

t I

h F

D-36

~ NUREG-1037 APPEN0!X B'

l

~

TWELVE-FOOT DIAMETER EQUIPMENT HATCH SEAL The equipment hatch is secured by twer,ty-four 1-3/4+ swingbolts.

There is a double rectangular seal design.

l_WMff24 dB )

WHA >=

\ f

-C u 9 1 9_L/ I W t..-

find

= = 4 I K

bolts 27.1E6 = 1.4222E8 lb/in.

The flange section has a 12 ft ID and is 3 in. thick find

- 144 Il 27.5E6 = 1.714E9 lb/in.

K fj,,,,=h=['/4Il The force tending to separate the seal with pressure is equal to

. containment pressure times the hatch door area F, = pA = p w/4(146)2 = 16741 p l

l i

D-37 NUREG-1037 APPENDIX B

3 After preloading the' bolts (torqued to 1900 ft-lb) the force in the

~

flange an'd the force in the bolts can be determined using the method of Appendix B.

I F

bolt bolts Kbolt + Kflange P 0 where F, = preload in' bolts F

p

= pressure load K

flange F

flange

= (p),p Kg+Kfj g p o.

The preload is determined for the 1900 f t-lb torque. . .

F o = h x 24 (Shigley, p. 310)

,1900(12)(24) = 1.563E6'lb 0.Z(1.75) so. . .

F f j',,g,= g ,g (16741 p) - 1.563E6 = 4 jl'hE9 1 422E8) p - 1.563E6 f

Fflange = 15458 p 1.563E6

'At flange separation, the oly force that remains in the flange is a compressive force (-) due *,o seal load.

D-38 NUREG-1037 APPENDIX B

L i

i i

F = EAc ='900 psi [2 -x 2w(73.5)(0.718)] (0.5) = 298424 lb seal Calculate the pressure at flange separation:

-298424 = 15458 p - 1.563E6 p = 82 psi After flange separation, the stiffness of the bolts controls. If the seal is assumed failed (no gasket) then seal force goes to zero, and pressure , 3x ,16741(p - 82)

K 1.422E8 hlts ax.= 1.177 x 10~4 p - 9.654 x 10-3 Leak area is given by AL * ' O '*

where A =

leak area (in.2) 0 =: minimum diameter of flange (in.)

Ax = flange separation as given above.

'A g = (144)[1.177 x 10~4 p - 9.654 x 10-3)

AL = 5.325 x 10-2 p - 4.367 NUREG-1037 APPENDIX-B

4 This area / pressure relationship is based'on no sea 11og action. It represents an upper bound on leak area.

~

Leak Agea (in. )

Pressure (psia) 0 0.00 82 0.00 116 1.81 120 2.01 126 2.34 j 3; 2.61 136 2.90 143 3.23

- 160 4.15 s

i l D-40

- NUREG-1037 APPENDIX B' l

V h ,

v APPENDIX D

t. i L - 1 E ;CRDtHATCHSEALANALYSIS _

7

{ --The CRD hatch has ' thirty-six-1-5/8+ bolts, .9;75 in. long..

^

K bolts'.= I = 6(w)(0.8125)$(27.5E6) 9.75 = 2.106E8 1

=K

~ ,

flan = from Shigley, p. 306, base flange stiffness on a cylinder with K

l dia. yof bolt for-ID and 3 times dia. of bolt for 00:

~

A = 36[w/4[3 x 1-5/8)2 - (1-5/8)2)) = 597.3 in.2 8

Kf = M = 597.3 (27.3E6[ 9.75 = 1.672E9 = K 2

.L

'Using the. method of Reference,15:

2 , 2'.106E8 1.6/ZE9 + 2.10bt8 = 0.112 K

g+K2 ,

, K 2

K) + K2 ,

Pressure l'oad to' bolt!s is a-function of hatch area !

.F,=PA=P{(37.4)2,r1098.6 P

~ The ' seal in this design is a 0.275+ 0-ring gasket of silicon, fit into' a . 0.192 in.- deep' groove.

/

' A3 ,,j '= 2 x [2w(18.7)(0.3)] = - 70. 5. i n.

-D-41 NUREG-1037- APPENDIX B-

F 3 ,,)

= EAc = 900(70.5)(0.275 - 0"1' ) = 19150 lb

, y5 Bolt preload (torque = 900 f t/?b):

F, = fg7 x 36 = 0 21 5 8) x 36 = 1196307 lb Flange separation pressure (Psep) is:

' -19150 = 0.888(1098.6 Psep) - 1196307 P = 1206 psi sep

- 1.e., no leak at accident pressure.

4 I

l i

I D-42 l- NUREG-1037 APPENDIX B

l t

I

\

i APPENDIX E LIMERICK LEAK AREA ESTIMATES l

l E-1 APPENDIX 8 NUREG-1037

Limerick Personnel Airlock Analysis l

~

The purpose of this analysis is to determine the leak area of the Limerick Containment personnel airlock as a- functon of containment pressure.

Limerick's personnel airlock as shown by the following figure is very nearly identical to Peach Bottom's personnel airlock. An analysis of Peach Bottoms's personnel airlock was previously performed for the same purpose as this analysis. That analysis was performed for both the interior and exterior doors and bulk heads. Due to the similarity of the

~

interior bulkhead design, an analysis of Limerick's interior bulkhead is not justified based on the following comparisons.

Peach Bottom Limerick s

design pressure 56 psig 62 psig barrel inside diameter 8 ft 4 in. 8 ft 7 in, bulkhead plate thickness 1-5/8 in. 2 The other significant design features are identical, such as, number and size of gusset plates, thickness of. barrel end and size of door opening.

Limerick's exterior bulkhead is similar in design to ' Peach Bottom's, however, not sufficiently identical to make leak area' estimates without performing additional analysis. Limerick's exterior bulkhead w'as analyzed using a plate-beam model for pressure loading. This analysis was performed using the structural computer program SAP IV. The following figure shows node point locations and element numbers used to model the bulkhead.

Quarter symmetrical boundary conditions were imposed on the model boundaries. All of the model input data is included with the attached microfiche copy of the SAP IV computer run output.

, NUREG-1037-APPENDIX B

z C

=

FT1 O

s 1 ~

o La N

I Limerick 12 ft. Dia. Hatch with Integral Personnel Airlock Conteenment Lines . e

. %\"-* .. ,

f 7

..g

<z u Z

'ZZ7EER 9- i . , . . 4

/

y,ip3 ,

. '\ '

r- , z.s , r2 b _j ,

( 4, Detail A

'. f. .

- ~ " ' "

Silicon-Rubber Gastet m.._ ~.

m_

a- - .,

3q,,,......,, --

,L,

r- ......

Detai A .. .

Gumdrop suicon-nubber I

j. 21&zztr 2 Geekete g j[

'. W

p7//yg xWm

= '_ .

fgbmp[g.-,~--...wr.ggaf a .

N *

.. ./

..M b.

m

'.. . . . ........ .. .. .... ... . . Detail

~~

B 16enes.e

h-n . ... . ,

40 39 35 37 36 35 34 33 32 31 I

- Z = = 6.0" 30

~

29 28l 27[ 25i 25 24 23 22 21 Z=0.0" 20

@l@l@l@i@

19i 18 17 16! 15

@~ @ .@ @

14 - Z= 2.0" 13 12 11

.G-@

~

10 9 5. 7 6 5 4 3 2 1~

Barrel Plate Elements ,, 2, 23 h -

42 2 .p . . . . . . . . . . . .. . . . . . . .4 3..................341 s

/^: b~, @ :

j ,

da 52 : 51 *

$0 2[5 [17] .

(16] ] (2). [1]

e 3] '

..'.5,,,

26 .. , , , , ,5,} . - - - f '. - . . . . . .-[58 27

/g:46@!54 @

i

'59 8 't )

~

63 (12] . [13] . [14] ; [15] Elements

@i-@l 8. ! @ c5)

$ !47 l55 !60 64

! ! [] - seam

@ ;@l @ l @ c6i Elements 2, .............a ....;a..........;e................. ,,

'a 1 a v .v: Us- - @

l 49 I57 ' I lt7, 30 - 62 66

[8 } - [9] (10] (11]

Bulk Head Beam & Plate Elements E-4 NUREG-1037 APPENDIX B i

- - - ~ '"

~_. . . . _ . . .

t ..

2

) ,

b 71 ' 70 59 68 67

[21] [20] -[19] [15]

r m n

[22) ) .

[28]

72 8' S3 79 78 ,

[23] i . ,

[29]

73 82 89 88 .

[24] ) i I [30) i' 74 83 SC 95 94

[34] .. [35] [36] . [37]

h-

[25]- ) I

[31]

75 84 21 96

99

[26; x v v v

[32]

i 76 85 92 97 100

[2?] ( .

i b' v (k!v -

[33]

f l 77 [se] 86 [39] 93 [40] 98 [4 t] t oi l

Door Elements TL8884-2 i

f E-5 APPENDIX B i f NUREG-1037 L

. .- , . - - . - - . -. - - .~ .,._z. .

m s.*

1 p 3.ias -

i L.i . .- L.

.l,..

1OF

"~ T ; i.,

-l - .- -

5" l

T:

~

I

,1 i

' Y

. a

~ f '-

3 t *

~

n.. ..

t Door properties A- AY IX 9 Y,. 99 d,=lY,-Fl A9 d,2 9 8.0 0.5 4.00 1.83 5.0 26.79 0.67 3.5 17.50- 1.17 6.84 2.92 6.69 10.42 19.53 4.36 55.51 0.46

-0.84- 7.06' -5.93 4.73 --18.79 -0.03 Totals 15.08 --

35.10 --

i 70.35 11.52 7 = 35 10:= 2.33 in.

15.08 l.

1 I, = 70.35 + 11.52 = 81.87 in.4-I i

l 1

l 1 i

- NUREG-1037 APPENDIX B

7- -

}=

Beam Properties- b

{- .- ._-

3

r s- =!

}_

v t-6 r

! 1 1

Y S S-ii T

@ 2.5-

_ es -- ---.4

.L .

JDoor frame top and bottom--beam elements 1-2, 16-17 2- Ig Ad A

3 Y, A,Y, d4= Y, - y gg j

$ 12.5 6.25 78.13 0.60 4.50 162.76

@' 16.25 -1.25 -20.31 6.90 773.66 8.46

@ '8.0 6.0 48.0 0.35 0.98 --

@ 13.0- 13.5 175.50 7.85 801.09 4.33 Totals 49.75 --

281.32 --

1580.24 175.55

.s A', Yi , 281.32 = 5.65 y_ {ZA 4 49.75 2 = 1580.29 + 175.55 = 1755.7 in.4 IX

EIX, + {A gg d Iy =2 5)3 .

12 )3 2.5 .5)3 ,.104.0 in.4 J=.f[6.5(3)3+12.5(1)3+6.5(2.5)3]=96.5in.4

- 'A = 12.5 + 16.25.+ 13.0 = 41.75 in.2 z

a. This moment of inertia is included with the plate elements.

'NUREG-1037 E-7 APPENDIX B

-Ay = 12.5 in.2-i l

1 A, = 13.0 + 16.25 = 29.25 in.2 !

--* t" e--

3 t 2.5- , ,

8" -

9 / i-s 'T t !

' l i

y -

s.s-a i

.I ,

7 t,Tj 2,:s-

_t

- a s- W Door frame sides--beam elements 3-7 Af Y, A,Y, d, = Y, - y 2 I A,dq X, 12.5 6.25 78.13 3.37 141.96 162.76 16.25 -1.25 -20.31 4.13 8.0 277.17 8.46 6.0 48.0 3.12 77.88 --

Totals 36.75 --

105.82 --

497.01 171.22.

y=

I i i = 36.75105.82 = 2.88 in.

{A, -

I, = 497.01 + 171.22 = 668.24 in."

y ,_12.5(1)3 + 2.5(6.5)3 y 12 12 = 58.3 in.4

a. This moment of inertia is included with the plate elements.

i NUREG-1037' E-8 APPENDIX B

I J = _f [6.5(2.5-)3 + 12.5(1)3] = 38.0 in.4 A, = 12.5 + 16.25 = 28.75 in.2

.. A x

= 16.'25 in.2 Ay = 12.5 in.2 r-Y',

y r a :

i !

fir I,.

s L , r ir i l

{

n ..

Bulkhead stiffeners--beam elements 12-15.

y = 6(3) - 16(0.5) _ 0.45 in.

22 E-9 APPENDIX'8 NUREG-1037.

I '*

x 2 3 + 6(2.$5)2 + 16(0.95)2 = 72.8 in.2 i I y =6 )3 = 0.5 in.4 J=6(j)3 = 2.0.in.4 Az = 6.0 in.2- {

Ay = 6.0 in.2 A-y = 0.0 in.2 Beam Element 8-11 values are 1/2 of these because of symmetry.

The model loading consisted of a 50 psig pressure, s

~

The maximum beam bending moment occurs at the center of top.'and bottom of the do'or_ frame, beam element 1, 1-end.

- M(2) = 1.48 x 106 in.-lb.

The bending stress _due to this moment is:

o=f=(1.48 (8.85) = 7,460 psi For a pressure loading of 140 psi the bending stress is:

o=h(7,460)=20,888 psi NUREG-1037 E APPENDIX B

n .

)

4 This calculated stress is considerably less than the material yield

' strength' of 36,000' psi .

Deflections for;the top of the door and door frame for a 50' psi pressure ,

are as follows: )

. Node .aZ (in.)'

50 .033805 ,

- 67 .033705 separation .00010 Node AZ (in.)

51 .033054

+

69 .032760

. separation .00029 ,

For these small calculated separations, this personnelf airlock leak area is Zero.

'. 1 i

i S

I APPENDIX B NUREG-1037 E-ll

. _ _ . ~. - . _ _ . . . . - - - - .-

1 f.

[ J

~

,1 Page E-12lcontains ' Limerick personnel airlock analysis microf,1che. It has been

omitted from this' appendix but is available from:

Mr.'Goutam Bagchi

' Office of Nuclear Reactor Regulation

_U.S. Nuclear Regulatory Commission.

Washington, D.C. 20555 t

'4 4

r 1

_4 4

h t

w u _

l l

1

" 1

- NUREG-1037' E-12 APPENDIX B l

i

' ~

i- LIMERICK LARGE OPENING PENETRATIONS' LEAK AREA ESTIMATES

/-

Limerick's-drywell. head,Jequipment hatches,' drywell head manhole

- hatch; control rod drive removal hatch, and suppression chamber access

- hatches are all external flanged penetrations. The bolt preload and seal geometries for the equipment hatches and drywell head are such that the relosure flange and' penetration flanges will have metal to metal. contact .

. with no_ preexisting flange separation. For flanged parts with a preload, the flange load is given~by:a ,

F2 = K2 p.p (K1 + KZ) where:

., :c F2 = flange load i

K2 = flange stiffness Kl = bolt stiffness P = . applied-load 4

F. = .preload.

Initiation of flange separation occurs when F2 becomes =zero. The required applied load for this' to occur. is given by:

p = (K1 + K2) F ,

K2

.i 1

.The drywell head flange and bolt properties are as follows:

l l

a. " Mechanical Oesign Analysis," by M. F. Spots. _

]

C-13 NUREG-1037 APPENDIX B 4

, - - , . -.y. . . - - . . --

--4 ,---.o _%w.~--,. ----,..w - ---,e ~~------,,4-,-,,.-------w-e--. r vm-,e-,,----- -- - -- _ .__

80 bolts - 2.75 in. diameter x 36.0 in. long .

The bolt material is SA 320 grade L43 with a modulus of elasticity 6

E = 29.9 x~.10 . psi. The bolts have a preload of 157,000 lb/ bolt.

F = SO x 157,000 lb .= 1.256E7 lb bolt area = 80(p 2.752 = 475.17 in.2 i

K1 = CAE . (475.17) 36 29.9E6 = 0.39E9 lb/in.

Flange area = (229.75 2 - 225.752) = 5723.98 in.2 The flange material is SA-516, Grade 70 with a modulus of elasticity E = 27.9E6 psi.

6 AE = (5723.98) 27.9 x 10 = 4.44E9 lb/in.

K2 = I 36 p , (0.39E9 + 4.44E9) 1.256E7 lb = 1.366E7 lb .

(4.44E9)

The drywell head has a diameter of 453.75 in. which the internal pressure acts on. This pressure area is A = '-(453.75)2 = 161,704.9 in.2 .

'4 Flange separation begins at a pressure of 1,366E7 lb = 84.47 ps19 .

161,704.9 in.

NUREG-1037- E-14 APPENDIX B

The flange separation due to bolt stretch.for' pressure loading from 84.47 psig to 140 psig is 3

aP . (140 - 84.47)(161,704.9) .= 0.023 in.

RT 0.39E9 The leak area for this flange separation is leak area = :(453.75) 0.023 = 32.8 in.2 This calculated leak area is based on the assumption that the flange seals have no resilience, therefore, this is an upper bound on the leak area.

The flange and bolt properties for Limerick containment's equipment hatches are:

Number of bolts 24 Size of bolts 1.75 in. diameter x 13.88 in. long Bolt E = 28E6 psi Bolt area 24 p (1.75)2 = 57.73 in.2 ,

AE . 57.73(28E6) = 116.5E6 lb/in.

K1 = I 13.88 bolt preload = 50,000 lb/ bolt F = 24 x 50,000 = 12E5 lb flange area = p150 2

1442 ) = 1385.4 in.2 i

flange material E = 29E6 psi 1

E-15 NUREG-1037 APPENDIX B i

1

~

~l K2 = (1385.4X29E6) 13.153

= 2.89E9 lb/in.

Force required for flange separation is:

'l p , (K1 + K2) p . (0.1165E9 + 2.89E9) 12E5 = 12.48E5 lb K2 (2.89E9)

The pressure area of the equipment hatches is A = T (146)2 = 16741.6 in.2 ,

The pressure required to separate the flanges is 12.48E5 lb = 74.5 psig 16741.6 in.2 The flange separation due to bolt stretch for pressure loading from 74.5 psig to 140 psig is a aP , (140 - 74.5) 16741.6 = 0.0094 in.

-KT 116.5E6 The leak area per hatch is w(144) 0.0094 = 4.26 in.2 ,

For both hatches the leak area for a containment pressure of 140 psig is 2 x 4.26 = 8.52 in.2 ,

Limerick'sremaininglargeopeningpenetrationhatchcovers[drywell head manhole hatch, control rod drive removal hatch, and suppression chamber access hatches (2)] all have an identical double tongue and groove 1

seal design. The seal groove is 5/8 in, deep by 0.469-0.484 in, wide with

NUREG-1037 E-16 APPENDIX B

a 1/2 in. wide by 1/2.in.-high seal. The tongue dimensions are 1/4 in.

. wide by 1/4 in. high with a full (1/8 in.) radius. In addition to having the same geometry, the operational procedures for tightening these closures are identical. -The procedure states " uniformly tighten the nuts until the gap between the cover and opening flange is approximately 1/16 in. or as required to obtain a leaktight closure for a test pressure of 62 psig between the two seals." Thus, this procedure specifies the maximum preexisting gap for these closures to be 1/16 in. The minimum preexisting flange separation is determined using the seal solid height for the seal gland maximum width.

Excess material volume is equal to seal and tongue volume less the gland volume. Using cross sectional areas the minimum flange separation is:

f . As + At - Ag Wg where As = seal cross sectional area At = tongue cross sectional area Ag = gland cross sectional area Wg = width of gland.

-As = 0.50 x 0.50 = 0.250 in.2 Ag = 0.625 x 0.484 =' 0.303 in.2 At = 0.25 x 0.125 + p0.125)2 = 0.055 in.2 i

af . 0.250 + 0.055 - 0.303 = 0.004 in. .

O.484 l E-17 NUREG-1037 APPENDIX B

I For these closures, the flange separation'due.t'o the pressure load stretching the bolts is determined assuming no preload. For this condition the flange separation is merely equal to the axial elongation of the bolts for the entire pressure loa ~ ding'. -

Pt.

af = AT where <

=

P closure pressure load for 140 psig t = length of bolts A =

bolt area for all the closure bolts i =

modulus of elasticity for the bolts.

Drywell Head Manhole Cover Pressure loading diameter = 24 in.

P=140(g)(24)2=63,334lb 16jindiameter-boltsx2.33in.long Bolt material is SA 193 Grade 87. E = 28E6 Yield strength Fy = 105,000 psi A=16(g)(0.625)2=4.91in.2 af = ( 63,334)(2.33) = 0.001 in.

(4.91)(zstb)

E-18 NUREG-1037 APPENDIX B

bolt a . P = (63.334) = 12,898 < Fy .

'K 4.91 The elastic assumption is valid.

Control Rod Drive Removal Hatch Pressure loading diameter = 37 in.

~ P.=140(p(37)2=150,529lb

'36 1in.diameterboltsx5in.long 8~

Bolt material is SA 193 Grade B7, E = 28E6 psi Yield strength Fy = 105,000 psi A=36(p(0.875)2=21.65in.2 6f = (150,529)(5.0) = 0.001 in.

'21.65)(28E6) bolt stress a = P . 150.529 = 6,953 psi < Fy .

A 21.65 I

The elastic assumption is valid.

Suppression Chamber Access Hatches L

l Pressure loading diameter = .53.5 in.

[

P=140(p(53.5)2=314,720lb t

E-19 NUREG-1037 APPENDIX B

35 j.in.diameterboltsx4.75in.long '

~

~ Bolt material is SA 193 Grade 87, E = 28E6 psi Yield strength Fy = 105,000 psi

.A=36(gj(0.625)2=.11.04in.2 af.= 314,720(4.75) = 0.005 in.

(11.04) 28E6 bolt stress a = XP = 314,720 11.04 = 28,507 psi < Fy The elastic assumption is valid.

e

?

h t

NUREG-1037 E-20 f ,

APPENDIX B

6 APPENDIX F GRAND GULF LEAK AREA ESTIMATES F-1 NUREG-1037 APPENDIX B

\

APPENDIX F

' GRAND GULF LEAK AREA ESTIMATES Personnel Airlock Analysis There are two personnel airlocks, not including those on the drywell.

These airlocks have double inflatable seal designs located on the door edges. The doors are forced against two 4 f t 6 in, door stops with metal to metal contact, then the seals are remotely pressurized. Details on the pressurization system will be forth coming. Reference 1 indicates that the double inflatable seal design may be inadequate for applications in a temperature environment of 330' and more. Plots of Grand Gulf accident temperatures and pressures from Reference 2 indicate that secondary containment could easily reach 200*F, but temperatures beyond that are as yet undefined. A major leak in drywell personnel airlocks could send secondary containment temperatures significantly higher, threatening the integrity of the^ personnel airlock seals there also. The leak area is calculated assuming that the inflatable seals have completely deflated due to temperature degradation at both ends of the lock.

l 1

F-2 NUREG-1037 APPENDIX B

b 1.o

.S d,,7

g. y- -

$- Y////////M N\

[ ] t -l ( f///As w s R^ hf -

,ps. f I w Ts\NNNN\[

'- N .NNN'N'.NNNNi 1 .

t' \\T see.% R-A y g NTs 6 k

.s-te g[

loor

' i

! [. - s, s.

/.2r v w w I

i <

7l . j E8 .

- g_ .

~. % 4-

. \ '~~ LakQ

> \. b N_

sa% t-B l E"

The resulting leak area would be 2

.A{ = 3/8 -(68-5412'+-(30)2 + 2 (6) = U ,14 in per door .

'if the seal was entirely blown r:A u, > Iditional 5/8 in. could conceivably be added to the 3/8 in. gap wio$ .

A{=.(3/8+5/8) (68 - 54)2 + 30(2) + 2r( ) = 125.70 in.2 per door .

F-3 NUREG-1037 APPENDIX B

The structural integrity of the bulkh ad, door frame, door stop, etc.

will be examined next to insure a gross structural failure will not occur due to accident pressures.

Stress Calculations for Personn'el Air Lock Serial No. 32877 c . it i t" -+ 6 l 4'3"'

a+and pms.ra ro ea

/ '

~

} /

/ k .

D $'N \

.I f 'i \

,v- if i ss , i

_eq. ._---. n - l ;,,

' ts.

i i

\[ h . .* T; 2 J!

-l Ig . . .

i 't n ,,

x.A e j ;

i Check Critical Components of Air Lock for Structural Integrity

1. Check Weld Stress @ Door Stop. Use p = 70 psi total load on_ door = Lg = Area (Door) x 0 psi AD s50.5 in. x 80 in. = 4040 in.2 LD = 4040 x 70 LO= 282800 %

F-4

'NUREG-1037 APPENDIX B

- each door stop-sees; = 28 = 141400 lb 2

I 1

=S" t-

-e 3.. - %

3,* \

je pl J MIS ' s 54' Weld Area ~= AWeld = 3/8 x 0.707 x 54 2

'AWeld = 14.32 in

- Calculate weld shear stress, T

~

weld

=141000 T

weld

14.32 '

T weld = 9850 psi per the AISC Code, Table 1.5.3,

.T allowable =.0.4 Fy -(Fy= 34.6 ksi @ 200'F, SA 516, Gr 70)

= 13.8 ksi T weld = 9850 psi < 0.4 = 13800 psi

'a.

F-5

+

NUREG-1037 '

2. Check Support of Door Frame Assume the following:

r.

Door and frame are very stiff

[

From p. F-4 LD = 282,800 lb.

2' *r an.*

3 Use the followin cross section

.- ss - : $

c~

% b =

} : 2.t?

4 l l '4 "

h 8" :lT calculate I x-

~A = 6(1.25) + 8(0.5) = 11.5 in.2 Ay = 6(1.25)3.5 + 8(0.5)'0.25 = 27.25 in.3-

y=h5=2.37.in.

Ix= $l +6(1.25)(3.5-2.37)2.8(ypd+8(0.5)(2.37-0.25)2 I

I, '= 22.5 + 9.58 + 0.08 + 17.98 F-6 L

NUREG-1037 APPENDIX B

Ix = 50.14 in.2 y s 10700 Ib calculate stress in ' beam

.? A I e MS z[I l

)

I'

.R R*P M,,x = 'Pl

=~ 70700 x 15 x 1/2

= 530250 lb/in.

o=f, y = 6 - 2.37 = 3.63

530250 (3.63) 50.14 o'b = 38390 psi Fy = yield stress = 34600 psi 0 200*F (SA-516, Gr 70)

The AISC allowable per-section 1.5.1.4 is 0.6 Fy = 20760 psi However, the AISC allowable is for a normal operating type event. The 70 psi considered here is a severe accident condition. Gross general deformations _are allowable and some loss of dimensional stability is

. acceptable. This criteria corresponds with a Level D event as defined in

.Section III of the ASME Code. Following the philosophy of the ASME Code, the' allowable primary stress for a Level'D event based on calculated elastic stress is 1.05 Su (Paragraph F-1323.l(b),Section III Appendices).

for primary bending type stress (1.5 x 0.7 Su = 1.05 Su). For primary membrane type stress.the' allowable value is 0.7 Su.

l F-7 NUREG-1037 APPENDIX B

s

~ Su'= 70 ksi 9 200*F SA-516 Gr 70 therefore: .

9-ob = 38390 psi < l.05 Su = 73500 psi Check shear stress in the weld about the section shown on page 2.

--A weld = (8.x 2 + 6 x 2) 0.707 x j = 7.42 in

= 9523 psi Tweld

7 4 T weld = 9523. psi < 0.4 Fy = 13800 psi 3.. Check stress in plates The largest-area of free plate is a 32 in. by 21 in. rectangle Assuming all edges are fixed, the fourth edition'of R. J.-Roark, Formulas for Stress and Strain, gives:

Case 41 a=.32,:b=-21,h=h=1.5238 1 o-= 0.4457, m = 0.0241 Max.S = S b 8'*D2 = 0.4557 70 (2 t (0.5) 4

.NUREG-1037, F-8 APPENDIX B

=.56270 psi 0 center.of long edge

.S b

Sb is a. local membrane' type stress and per ASME Code philosophy for a level 0 type event, the allowable stress is 1.5 x 0.7 S = u 1.05 S u Sb = 56270 psi < 1.05 Su = 73500 psi ,

4. Check Weld Shear Stress Around Total Section For the 3/8 in. throat' weld applied at the edge of a circle of radius 57 in.:

F weld = (57)2 x 70 = 715000 lb Aweld = 114 m ( ) 0.707 = 94.9 in.2 Tweld =

= 7530 psi

'T = 7530 psi < 0.4 F = 13800 psi weld y Equipment Hatch Evaluation This hatch has a typical flared flange containing two concentric-grooves which house the 3/4 in. diameter double 0-rings. These seals:are made from EPDH material, compound No. E603, which hasla durometer rating of I

55-65. Increasing containment pressure tends to increase the compression p seal. The only possibility for leakage would be a gross structural _ failure j of the hatch. door. Temperature degradation of the 0-rings would probably l

not produce a sizeable leak since metal to metal contact with considerable force would still exist. Assuming the gaskec degrades and there is a 0.0002 in, gap all around the 9.5 f t radius hatch, the leak area would be 1

l 1

NUREG-1037 F-9 APPENDIX B

1 - .

i fless than' O.15 in.2 Surface finishes of 63 'u in. on the metal to metal contact areas, along with axisymetric design and loading of the hatch door

itself support the assumption that gaps, even without a seal, will be very small..

The. collapse of- the hatch door shell is examined below. Door ;

thickness and radius of curvature were taken as typical from previously

analyzed plant designs.

, r ~ ~ %L o -ri 3s j n y 70 psl.

/ 7 cahi,mt pressw.

7e

'(/g aM .

N 2 a* %*h 3D x x leh.k slall 3

From Roark, p. 354, case T, the probable actual minimum critical buckling pressure is

P.

-* 0.365 Eh 2 "

-(0.365) (27E6) = 152(h)2psi er 2 a (223)2 From Timoshenko's " Theory of Plates-and Shells," p. 436, actual uniform :

membrane compressive stresses due to 70 psi external pressure are:

l I

NUREG-1037 F-10 APPENDIX B

m.

N = j -ap93 , = -2'23 1+ 31 = -8405 psi '(from p. 436 Timoshenko) 1 N, = ap'(1 + cos e - cos e) = -4975 psi These low stresses are for three times design pressure; accompanying deflections are small.

. Fuel Transfer' Tube Evaluation A cross section of the fuel transfer tube and closure hatch in the containment fuel pool are seen below.

v g Au!., J M.,'J.isiv~

, ,l

" '8

, %)v %L :l's nJa & M p**tAmd u co' c '

4 .g ( 22' A wJa .q.

L...-~ ~ },

(cus -

.g s;; g4 2,2 c.cy<a, w g, Asp p,a.Jua.

culmka d wit 31 pis osud frena <

%.y A ,Jand passe

~

f9f* F M %s% 6e The fuel transfer tube closure hatch assembly is a remotely operated design which clamps the hatch door to its seating flange at four locations around the perimeter of the 50 in. diameter door. A-double 3/8 in. o-ring

- seal design is used which is increasingly effective with rising internal containment pressure. The center line of the hatch is 22 ft below the .

! water surface, which adds sealing pressure.

l 1

NUREG-1037 F-ll APPENDIX B

4 lb !

. hydrostatic pressure'= ogh = '62.4 32.2 (22 ft) = 1373 psf 32 2 u-

.= 9.5 psi The transfer tube and closure assembly should not leak' under accident conditions for the following reasons:

1. Rising containment pressure increases sealing force

2. Hydrostatic pressure increases sealing force

3. Volumetric swell of o-rings from fluid environment increases sealing effectiveness i _ NOTE: Any leak that might occur would be scrubbed by fuel pool water in

-the-spent fuel building.

Check of the structural stability of the hatch door under uniform pressure:

A =w (25)2 = 1963'in.2 door Fdoor = 1963 (70 psi) = 137444 lb

= 2w } I = 875 psi, very low t

3744 from Timoshenko, " Theory of Plates and Shells":

'r,,x=f2 "

j, 2

= 22786 psi (Sy = 30000) l 4

~

NUREG-1037 APPENDIX B

' max

na"64700(25)4 " 64 (2472530) = 2.76E-4 in., low where 3

" Eh

27E6 (l'l 12 (1 - y2) 12 (0.9i) = 2472530 The plate will not collapse at accident pressures i.e., seals should remain functional; no leaks will develop.

I F-13 NUREG-1037 APPENDIX B

EVALUATION OF SUPPRESSION P0OL BY-PASS LEAK AREA Per G/G FSAR Section .3.8.3.7.2.2 the G/G drywell leak rate during the

-30 psig structural integrity test was 3200 scfm. The allowable is 84,000 scfm.

2

*Z 1= + + 2 + 'h P1-P2 = ah 1.5 v2

fj y 2 y 29 d 29 lb pg '- p2 = 30 psi = 4320 ft Y

air = 0.25 t 3

Assume f = 0.038, 1 = 48 in., d = 0.02, 0.01 in crack.

(0.038) 48 = 91.2 0.02 (l.5'+ 91.2) v 2 4320 2 (32.2) "M 1

- [(4320)2(32.2)l = ft 109.6 sec (625) 92.7 j 0

leak area = - = 0.49 fd = 70 in.2

( 6) l Other suppression pool by-pass leak paths include: (a) piping penetrations, (b) drywell head, (c) equipment hatch, (d) personnel air lock. Containment piping penetrations have not been a significant source )

NUREG-1037' F-14 APPENDIX B

of containment leakage; therefore, the drywell piping penetrations are considered to provide a negligible by-pass leak area. The drywell head is below the upper containment pool. The water in this pool should provide adequate cooling to the drywell head seals; therefore, by-pass leaking through the drywell head seal flange is considered insignificant.

A review of the drywell equipment hatch design was performed with the conclusion that it would not contribute to suppression pool by-pass 1Gakage. It has sufficient strength to resist anticipated drywell wall pressure differentials. It has double o-ring seals which have been dGtermined to be unaffected by hydrogen burns. Any leakage of this hatch cover should be included in the preexisting leak area determined above from the drywell wall structural integrity test.

The drywell personnel airlock is the same design as the containment personnel airlocks. It therefore will not leak as a result of pressure loading; however, it is subjected to the drywell temperatures which will very likely destroy the EPDM seals. Kris Parczewski of the NRC has performed heat transfer calculations of the drywell personnel airlock and has determined that the door seals for the wet well side of the personnel air will reach temperatures in excess of 500*F. The drywell side door seals are subjected to even higher temperatures. Based on this analysis it is assumed the pressure integrity of this personnel airlock will be lost at 5 hr or 300 min after the initiation of the severe accident. The upper bound leak area for the personnel airlock with the seals entirely blown out is 125.7 in.2, The suppression pool by-pass leak area is the sum of a linear leak area as a function of pressure differential across the drywell wall and a step increase in leak area at 5 hr. Since the pressure differential is limited by the suppression pool water column height of approximately 4 psi, the linear leak area will have only 9 in.2 contribution. The by-pass l leak area can be simplified to be time dependent only. From time 0-5 hr l the leak area is 9 in.2 and from 5 hr on the leak area is 135 in.2 ,

NUREG-1037 F-15 APPENDIX B

. - =- ..

REFERENCES

1. NRC Letter, " Potential Problems on Airlock. Seals in Containments and Drywells," Robert L. Baer, October 7, 1983.

2. Personal communication with C. Hofmayer, Brookhaven National Laboratory.

d J

=i f

i 9

4 1

I I

NUREG-1037 F-16 APPENDIX B i

APPENDIX G SEQUOYAH LEAK AREA ESTIMATES 4

G-1 NUREG-1037 APPENDIX B

APPEhDIX G SEQUOYAH LEAK AREA ESTIMATES l

Sequoyah Per sonnel Airlock The interior bulkhead and person 1el airlock door were modeled using the SAP IV finite element computer code. The model included enough of the airlock sleeve to get away from the cylinder edge effects. Short truss members were placed between door and door frame so that when pressure was applied to the door and bulkhead, any gap that developed between door and door frame would result in a truss under tension. This truss was then removed from the model and another computer run was made. This process continued until all remaining trusses were under compression (indicating the door was firmly seated at those locations).

The interior personnel airlock door tends to be forced against the door frame with increasing containment pressure. This increases the sealing capability of the design. But, as the SAP IV model showed, the differences in stiffness between the bulkhead and the door can lead to the development of leakage paths. This occurs when the door is stiffer than

^

the bulkhead; the door frame bows and the door does not. Usually, as in this case, gaps are within the normal sealing range of the rubber seal at the door edge. The capability of the design to contain pressure is therefore not a structural problem, bJt one of seal integrity With rising containment temperature.

The airlock sleeve is welded directly to the liner. A bellows connects the outer shield wall to the airlock sleeve and takes up any expansion that may occur in the liner due to temperature and pressure effects from the containment atmosphere.

i i

f 6

NUREG-1037 G-2 APPENDIX B

Figure G-1 shows the finite element model developed for the analysis.

Fif ty psi pressure was applied to the f ace of all containment-side olements. Once displacement results were obtained, the gap area was calculated from the extended length of the truss members between door and' d:or frame. It should be noted that this gap only exists if the seal is n:t resilient.

& - T. tit - @ - 4.ots - & - 4.340 - @

co.cossey 'mw2 t/@

,Xe A=f(0.002935)(4.39)+0.002935(4.016)+f(4.016)(0.004302-0.0029

+0.004302(5.858)+f(5.858)(0.004994-0.004302)

A = 0.0482 in.2 Total area (4 x area of quarter model) = 0.1928 in.2 Total area of two locks = 0.3856 in.2 The question still remains with regard to the survivability of the i

outboard door seal. If the inner door is considered failed, there may still be enough thermal lag to prevent temperatures from rising high enough to heat the outer bulkhead and door. The outboard seal would not degrade, and the airlock would remain sound. Further thermal analysis work is required in this area.

l NUREG-1037 G-3 APPENDIX B

4

/

\

(

\\ \

e

\/ 't/\/\

_t

,!,/j' Y l ~ ,

j ,/ N

.'. \;\ ,M N, / '. /'\\

DM/ ,s

/

~

\,

, /

x

! y .

'\ \

b ,/

N, / -

\ D N ,

% __./ ,

1 ,

\\i

\ f '., ,.- .

4

/ \ . ,

~ . .

\. ,, ,r

\v / s, w

\ \

[, ,

i /i I

(\. /

i Figure G-1. Personnel airlock model.

l NUREG-1037 G-4 APPENDIX B

At the 50 psi pressure, all elements remained within the elastic range of the material. The large bulkhead stiffeners take up most of the applied Icad.

Sequoyah Equipment Hatch This hatch is a 20 f t diameter shell'with a 20 f t spherical radius.

I Its edge. ring has a 2 x 2 in cross-section. The barrel flange has a double gumdrop seal design which is compressed with rising containment pressures. Even if the seals degrade with temperature, the mating faces are machined and they would remain in intimate contact. Leak areas will remain negligible as long as the hatch shell does not collapse from pressure, and the clamping hardware holding the hatch to the flange remains operable. Figure G-2 shows some details of the equipment hatch.

From Roark, p 354, case T, the probable actual minimum buckling pressure for the 20 ft diameter spherical shell is P=[1-0.00875(+-20*)][(1-0.000175f)(0.3E)(f)]

= [1 - 0.00875 (30 - 20)] (1 - 0.000f75 (2 )[0.3(29E6)(h)]=73 psi This is above the expected failure pressure for the containment.

As the hatch shell is " flattened" by containment pressure, its circumference (where it contacts the penetration flange) increases. If the diametral growth is too large, the swing bolts and clamping hardware could be overstre'ssed. If enough of the clamps failed, the hatch could lose its seating and leak paths may be created. The radial growth of the hatch due to a 50 psi pressure acting on its convex side is examined below.

l l This analysis was developed from Reference 1, Article 131.

NUREG-1037 G-5 APPENDIX B

V.

c .:*',, * .a ..

c.mm -

-t >

FD -C] .Md- -+4,h; r

C.

- +

,* I .c v -

~ *===*s .

-- ~ et!^< ^ CC ma :* - - ~ * = , = . = ,

A I -

e d ##J '#f' a .p .< . e. s. ,, .e e.

t .., .t

.. , Lc

, s.

,g , ,,,,,,. - ..- ,

..,. ....... f .x 4, ..

se .

g wr.

a s-; ...vs_.,, -_

. . . .. . ~ . .... .

. . .- ~ --

?'Q \

i=== x{. . ,E' :

C *

8. .

!id !!

=a -

U. 'ar* ===r y.') :

O%

=.g.

1 .t _ __. .

a

/ a -

s 1 .

, -\

s

., ,psee. .

/ O

\ame.s y_ p.e -

./ \g m sm .

s .a.o .-

,, a e .e '

{i tl o

j. e .e a rs* Q g .,,

eh .. - Ic A ',

l?

\ / w

.e.G..- e:.

s

,~.-e. .-

_- Z *,.*.*", , ,,, y- *

. a s 9 y . e~,

.( ', $,f*,7*?,. a

,. h s

,,.s ._ _s L1. nr yn;;;;;?" '*12n

. A, .-* g- -- *** ? *e e5=::= _' \4s. -

.~ ":~.*.::.". '

.J -- ---- -

B=0 I, ,

1;tvatica n.A SLUo0N &3 Figure G-2. Equipment hatch.

G-6 NUREG-1037 APPENDIX B

D r=k Lhi s k.att

elj ' g,.

r ay'

- 6.a m N.^ d

r,- su ,

J a .

Le m When.the shell is symmetrically loaded with pressure, membrane forces N,, N, develop. .

i , ,

-aq , -(240 in.) (50 psi) , -6431 lb

, 1 + cos a 1 + cas 30 in. circ

-3961 N

3 = ag (j , fg3 , - cos a) = n. circ These forces cause an increase of the radius, rg , equal to 6, = (N, - r N,) = 29E6 0.75) [-3961 + 0.3 (6431)] = -0.01121 in.

This displacement is accompanied by a small rotation of the edge tangent, N , which will be neglected here.

A thrust. Hg , exists in the horizontal direction at the edge.

I Hg = -cos a (N,) = -cos 30 (-6431) = 5570 ,

NUREG-1037- G-7 APPENDIX B

__ -_.- . . _ =_. . . _ . __ _

u..

, The'correspon' ding tension force in the-e'dge ring is H og r and the

]

elongation is , ;

.Han r

e " T6T

~

The increase of the radius r, due to H 9 is H,r[

5570 f120?

2 ;

F

'l " *e o " Ebd 29E6 ;2) J,2) = 0.6914 in.

1 ..

t The edge ~ deformation of the shell must be identical to that of the edge-

-ring.' The uniformly distributed (on the edge and ring circumferences) radial forces are denoted by H, while the couples are denoted by M,.

2 2

,,2A sin a g, , 2 a A sin a H = horizontal displacement of shell edge Eh Eh

, 2 L23)2 sin 30 g ,2 (240)'23 (sin 30)2 H 29E6 (0.75) 29E6 (0.75) -

4

. = 2.432E-S Ma - 1.26897E-4 H j The rotation of the shell_ edge is'given by

-4 A 3 2

'"W**2A sin a Eh

-4 (23)3 H = 29E6 (0.75) (240) 29E6 (0.75) sin 30

""

2 (23)2 H

= -9.323E-6 Ma + 2.4322E-5 H ,

\

N' ,

, c I = D g

~

-)

r i

NUREG-1037 G-8 APPENDIX B ;

The action of Ma and H upon the ring is statically equivalent to the combined action of the overturning couples:

.T = Ma + He and of forces H applied horizontally to the centroid of the ring section.

The latter cause a radial displacement of the ring equal ~to 2

Hr 2

= .2414E-4 H 62

Ebd " 29 6 (2) ,

The rotation of the transverse section of the ring, due to T, is given by 2

,2

IH" + N8)

Ebd

= 12 (120)2 29E6 (2) (2)3 [Ma + H (1)] = 3.724E-4 Ma + 3.724E-4 H Now, the total horizontal displacement of the shell must be equal to that of the edge, and the same goes for the rotation 6, + 6 = 6 ) + 62 + '2'

+=$ 2 Substitution of the calculated values into these equations yield:

-0.01121 + 2.432E-5 (Ma) - 1.2687E-4 (H) = 0.6914 + 1.2414E-4 (H)

I

+3.724E-4 (Ma) + 3.724E-4 (H) i NUREG-1037 g.g APPENDIX B

r. ,

and

-9.323E-6 (N) + 2.4322E-5 (H) = 3.724E-4'(Ms) + 3.724E-4 (h)

Simplifying the two equations:

-3.4808E-4 (Ma) - 6.2341E-4.(H) = 0.70621

-3.8172E-4 (Ma) - 3.4808E-4 (H) = 0 (Ma) + 1.791 (H) = -2028.9

(% )~+ 0'.9119 (H) = 0 0 + 0.8791 (H) = -2028.9 I

H = -2308 I . of cire.

I"*

Ma = -2104.7 n.

Now, the total horizontal radial displacement of the hatch shell ring due to " flattening" from containment pressure is:

4,+ 6 = -0.01121 + 2.432E-5 (-2104.7) - 1.26897E-4 (-2308) = 0.23 in.

The implications of this displacement are examined below, i

NUREG-1037 G-10 APPENDIX B

,. .- loM

l o. J sM 23: ,; -

w 2.x 2. .a A

$ ( riS ens ln n-"(

'l I ,.

^

g d

a . c, e- harr4( punkmS dl*$*

l 14*A u .x u u %-

(u)3 The swingbolt design allows the horizontal movement to occur without bending of the bolt. If all the deflection is taken up by the shank of the bolt it will lengthen by AL, 0.23' n

11.75" tan-1 0.23g = 1.12 0

3

, = 11. 5_25 AL = 0.00225 in, 2

c = h = 0. ] ;5 =0.00019149=f,o=5553 psi G-il NUREG-1037 APPENDIX B l

-This stress is in addition to the preload on the torqued bolts. It assumes rigidity of the eyebolt attachment hardware and is conservative.

l w

The force in the bolt'is oA = 5553 (1.25 [ = 6820 lb.

The lh e eyebolt pin is ok by inspection Theeyeboltpinsupportplatesarefin.x4.5in.andstraddletheeyebolt.

A p =2(f)4.5=4.5in.2,t = = 1515 psi Stresses in the clamping hardware due to the shell edge ring displacement are low. Yield of the material is at 50 ksi with an ultimate strength of 90 ksi. Failure of the latch clamping system is unlikely.

4 9

NUREG-1037 G-12 APPENDIX B

, - - , n ----.-,.-....-,e-.v . - -n---,- , - , - - - - - - - - - - - - - - - --

le-

+ s

.. ; SEQUOYAH EQUIPMENT HATCH SEAL LEAKAGE L. Greimann, F.'Fanous, D.'Bluhm July 1984-71 Summary fThefollowingsummaryconstitutesanabstract.of.a.

socn-to-be-published report by the same title in which complete details of

- the analysis ~are presented.

Background and Objective .

As part of the Containment-Performance Working Group study, an

- analytical effort was undertaken to investigate the leakage characteristics of the.Sequoyah equipment hatch seal during severe accidents. The effects of containment deformation and hatch postbuckling on the distortions at the

' sealing surface were to be. included.- Only-leakage at the seal surface'is considered in this work. Leakage through a crack in the steel plate was not considered.-

2 Equipment Hatch Model f

The equipment hatch is one of several penetrations in the Sequoyah containment vessel. .If penetration effects are neglected, gross yielding J of the .1/2-in. 'shell plate near the springline will occur at a pressure of

'50 to 60 psig[ ' A three-dimensional finite element model of the equipment

' hatch was developed which includes: shell elements for the hatch, sleeve,

' containment jlate and stiffeners; prestressed bar. elements for the swing bolts; and friction elements for the seal' interface. One-quarter symmetry of the assembly was enforced. Geometric and material nonlinearities were

. included. A single lobe imperfection was included in the hatch shell

~ model. The imperfection amplitude was equal to the ASME tolerance

' allowable and the wave length was equal to the elastic buckling length.

i 4

i

- G-13

. NUREG-1037 APPENDIX ~B

P i llie? ,

O

[g-Loading which corresponds to pressure inside the containment was applied in increments. Analyses were made with two coefficients of friction for the

~

steel at the seal interface: 0.3 and 0.6.

~

Results The maximum' strain of about 1.4 percent occurred in the top of the sleeve, near its_ intersection with the containment. The maximum relative sliding and rotation at the seal interface was 0.9 in, and 1-1/2 *,

respectively, for the 0.3 friction coefficient. Hatch buckling occurred in the range of 85 to 90 psi.

Conclusions

The Sequoyah equipment hatch seals should not leak before very large

't strains develop in'the 1/2-in. containment shell plate near the springline, which occurs between 50 and 60 psig. In the unlikely event of hatch ,

buckling, postbuckling deformations would not introduce leakage.

Fuel Transfer Tube There are four components comprising the pressure boundary of the fuel transfer tube, as shown in the figure below. Each component will be

~

evaluated on the basis of nominal section data. Since this is an~ASME code design, welds are assumed as strong as the nominal section. The tube is connected to all structures other than the containment pressure boundary with bellows rated for a 1 in, axial deflection. The maximum containment radial deflection predicted in Reference 8 is also on the order of 1 in. so that bellows loads can be neglected. Thus, only the pressure load is significant. The maximum pressure load is based on the section area of the sleeve.

NUREG-1037 0"I4 APPENDIX B

' F e pA 2

, pw D 4

, (50) w (24)2 4 .;

= 22619 lb

1. Sleeve--the sleeve is a 24 in. 00 by 1/4 in. thick cylinder.

Hoop stress is given by ,.

o=f

=(50)h

= 2400 psi Although the material of the, sleeve .is not well specified .it has been identified as a stainless steel, so that minimum yield stress can be established at 20.2 ksi (304L 9 250*F, ASME). The failure pressure is given by

= 421 psi Pp l= 50 (22

2. Flange--the flange is a disk 23-1/2 in. 00, 20 in. ID and '

1-1/8 in. thick. Shear stress in the' flange is given by o = F/A .

1 A= Dt l

G-15 NUREG-1037- APPENDIX-B

l

= v20 (1-1/8)

= 70.69 e = 22619/22.5 o = 320 psi i

The flange is of A240, 304 stainless steel (Sy = 23.7 ksi 9 250*F,ASME). Failure of the flange will occur at:

,F , (50) (23700) 320

,-3703 psi 3.- Trans'fer tube--the fuel transfer tube is a 20 in. OD by 3/8 in.-

, thick tube of 304 SST (Sy = 23.7 ksi 9 250*F, ASME). Hoop stress is given by PR a=7 >

10)

=(50)(f)

(

-= 1333 psi

. Failure of the tube will occur at pF , (50) 1333 (23700) ,888 psi

, 4. Blind Flange--the blind flange is a 20 in., 150 lb flange. Seals in the flange are compressed by containment pressure, so that I leakage through the seals will be negligible. Since this is'an ASME code design, the~ ANSI B16.5-1977 pressure temperature ratings are applicable. This source gives a flange pressure NUREG-1037 G-16 -

APPENDIX B

,-,,-w

rating of 220 psig at 250*F. It represents a conservative failure pressure but it demonstrates such a large margin that the detailed analysis required to raise the failure pressure is not indicated.

Py = 220 psig The pressure capability of the fuel transfer tube is controlled by the blind flange sealing the end of the tube. This flange will fail at some pressure above 220 psig.

Sequoyah Containment Vessel Shell Analysis ,

at Personnel Airlock Penetration The model developed for this analysis is seen in Figure G-3 along with the boundary conditions used to solve the problem. Fifty psi p.ressure was applied normal to surfaces exposed to the containment environment. In the vertical direction a load was applied corresponding to the force that holds the dome in place.

The liner code, SAP IV was used for this study. The model was made'up of 231 nodes (with 1217 d.o.f.), 216 quadrilateral shell' elements and 6 boundary elements. It included an 8.33' slice of the containment shell.

At least 4 ft of vessel wall was modeled all around the personnel airlock

~

barrel.

It was found that materials at some locations would go beyond yield at

- a containment pressure of 50 psi. This was most notable at the top of the airlock barrel to containment shell junction. The first few concentric rings of elements around the barrel also surpassed yield values (32 ksi).

Obviously, since the code is only linear, the displacement and stress results obtained are unreliable, but some important things were learned from this scoping study model.

1 G-17 APPENDIX,B NUREG-1037

l l

b p es &

e55 a

e c.

o'i I

7 '^ , # N, .

( )p* Pt [hh b'T*=

,, 7 \ )

L

.d c s' rem fnie.rtt0 r$T7 . fK4y'

/) P(ass ur s M1tnd c.fg.J sw s g,,, gre, sw(oces

./ (

butary \

a w.nu 6 .ec,-ptbh ~

i ci' f* r*J#

- a' l

j /

I

\\ __ \

/ '

)!y.,s N em J..n=

+ +< h 4 A

@bj 3 X

NN -

i

.M6' l

1

\

l Figure G-3.

SAP-IV finite element model of containment and personnel airlock barrel.

NUREG-1037 G-18 APPENDIX B

\__ _ _ ._ __. ____ _ _ _ _ _ _ . - - - - . -. -

The stress concentration factor around large openings is significant and warrants further investigation. Such an investigation might lead to lower estimates of containment vessel first yield pressure.

At pressures above 50 psi (some' spikes exist as high as 78 psi in prGdicted containment pressure curves), the rupture of the vessel could be

-controlled by behavior of the shell in the large penetration region.

Yielding propagates along a vertical line running through the center of the large opening and if rupture occurs, leak areas could be significant.

Sequoyah Personnel Airlock Penetration--Containment Shell Interface Nonlinear Analysis A nonlinear analysis of the containment shell-personnel airlock interface was performed to investigate strain concentration for a severe accident pressure of 50 psi. A quarter segment of this interface location was modeled by taking advantage of symmetry of the penetration. The model used to perform this analysis was planer, consisting of two dimensional plane stress and truss nonlinear elements. The computer code ADINA was used to perform this analysis. The analysis was performed for a loading corresponding to a containment pressure of 50 psi. Truss elements were used to model the personnel airlock barrel. The first radial 12 in. (two rows) of shell elements represent the reinforced shell region (t = 1.5 in.). The model was extended an additional 50 in. in the radial y and z directions. The first four rows of elements were eight node two-dimensional plane stress elements. The fifth row was five node two-dimensional plane stress elements and the remainder of the elements were four-node with a few three-node elements. -The material properties used to perform this analysis are: yield strength = 45,000 psi, elastic 6 psi, plastic modulus = 5.0 x 10 5 modulus = 28 x 10 psi. Since the model is planer, the loading was applied by imposing a stress equal to the hoop stress on the z = 107 in, edge and the corresponding axial stress to the y = 107 in, edge. The hoop stress value is.given by l

i i

6-I9 APPENDIX B NUREG-1037

c h 0.69 50,000 psi

and "a*

2( ) = 25,000 psi.

Figures G-4 and G-5 show the model nodalization, boundary conditions,

and imposed loading.

l The results of the. analysis are shown below. The maximum stress calculated was 52,800 psi, which has a corresponding strain of 1.7%. These results suggest the penetration opening is adequately reinforced and strain concentrations are relatively low so-that leakage is not likely to result for this loading. The penetration ovalization is shown below by the calculated displacements.

M w

\ -

Y ,.. .. ~

Qs 4S***

q.- a, h*n.' ,

7 sy,

.26' y,n , g /~. 7 gm s s 4C 1" P' l

w )'

~ ~ '

sys . for l .

4

~NUREG-1037 APPENDIX B

i G: 2]~poo Ps?

i-t c _

a &

y fr=io x 4 e

\

4 e

. E:. W 5

4 u i

-(

_ d t:: / 6'-

$ T' v- r-p>~

g'

. C *'N re ,s s.. elenh T~ %% _

~ %

k a >> p p x e .

y*

i-Figure G-4.- Personnel airlock / containment shell model.

G-21 APPENDIX B NUREG-1037

tg

.r 9" L -

_ . si

G I

to

> p, G

y ;,-. ,

N '&08 7

x f5 -

E b Qa 4

,ri f, :T All' g <> 4

n. -; -3 in< >gog tr

- - f ,:3 y

.M _fp [6 '4) go

1lC

.s g-

,32',

, psg g ,,97 g h h<>31'g g32fg .w &

h m

- e g4m2mg ak" y M v

33/

$m r

,e

~ t m y tw 1

Figure G-5. Model detaffs at symmetrical bound aries.

NUREG-1037 G-22 APPENDIX 8

REFERENCES.

1. ~S. P. Timoshenko, Theory of Plates and Shells, McGraw-Hill Co., 1959.

un k

I G-23 APPENDIX B NUREG-1037 L.

l 4

Pages G-24 printout. through G-92 contain Sequoyah airlock penetration analysis compute from: 'These pages have been omitted from this appendix but are available l Mr. Goutam Bagchi.

Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission -

Washington, D.C. 20555 P

P l

NUREG-1037 G G-92 APPENDIX B

,5 4 - 4 -- n - M- &L<n> A. 2 3 = - *, A = ,asA e y S

V e

, APPENDIX C l

' PROBABLE LEAK AREAS IN CONTAINMENT ELECTRICAL PENETRATION ASSEMBLIES FOR SEVERE ACCIDENT STUDIES

f P

l l-t 6

h t

l.

NUREG-1037 i

PROBABLE LEAK AREAS IN CONTAINMENT ELECTRICAL PENETRATION .

ASSEMBLIES FOR SEVERE ACCIDENT STUDIES PURPOSE The purpose of this study is.to determine probable leak areas in containment electrical penetration assemblies (EPAs). This review will. provide information !

which.can be used in determining the impact of EPAs as a potential leak source

~

in containment buildings. This study will attempt to focus on two problem areas: -[

the potential for leakage under severe accident conditions,and the potential !

for leakage with conditions of aging. The potential for leakage under operating

~

i conditions will be included in the Brookhaven National Laboratory (BNL) study !

(incorporated into this report, NUREG-1037).

BACKGROUND Studies'are currently under way to understand the functional failures of contain-ment. Each containment building has a large number of penetrations, therefore',

.there are a large number of potential leak paths.- There are four'NRC programs that are concerned with containment integrity beyond' design conditions, the -

Containment Integrity Program (NUREG/CR-2549), the Isolation Valve Program (at INEL), the Integrity of Containment Penetrations Under Severe Accidents Program, and.the Electrical Penetration Assemblies Program (NUREG/CR-3234).

The severe accident studies have provided a major impetus to an in-depth inves-tigation into the physical integrity of containment buildings and penetrations.

One such study, a severe accident analysis conducted on a Mark I BWR-(NUREG/

- CR-2182) indicated very high temperatures in the drywell area, which is the location of the majority of electrical penetration assemblies.' Because of the high temperatures, sit was postulated in NUREG/CR-2182 that the sealants would

fail.and all the electrical penetration assemblies would leak before structural >

' failure would occur. Since other containments have similar electrical penetra-tion assemblies, it was concluded that all containments would experience the

same type of failure. The follow-on study (NUREG/CR-3234) concluded that many .

electrical penetration assemblies have a low potential for leakage, because at least one set of EPA seals is exposed to a much lower temperature. These seals i

l are outboard of the containment, which provides thermal lag during an accident ;

p progression. However, other types of penetrations with just inboard seals could

[ have a high potential- for leakage if they are sealed with organic materials.

l PROBABLE. LEAK-AREAS UNDER OPERATING CONDITIONS The' main source of information concerning probable leak areastin EPAs can be L

derived from the Licensee Event Reports (LERs) and Type B Integrated Leak Rate l Tests (ILRTs). The LERs are an important documented source of information, but l

do not represent the entire history of experiences. The data listed in Nuclear -

l : Safety Information Center (NSIC) material through February 1983 indicated 265 L leak type of failures for all penetrations (letter, Feb. 1983). Of the 265 L.

total, 18 are related to EPAs. . Fifteen of the 18 are leaks through only one of '

l the double seals. The remaining three are probable through-leaks. The word

probable: is applied to these three cases, because the complete details are not l NUREG-1037 C-1 1

l

l i

i

included in the NSIC listing. An estimate of several possible pre-existing

' leak areas can be derived. The three cases with probable through-leaks occurred at Hatch 1 (1975), FitzPatrick (1979), and Yankee Rowe (1981). The Hatch 1 incident was an initial leak rate of 50 cc/ min through two coaxial cables.

This problem was fixed by an application of epoxy sealant at both ends. The FitzPatrick incident occurred during an outage, when new cables were pulled

.through a penetration and not sealed. This large hole would have prevented the containment from passing the ILRT. The incident at Yankee Rowe involved five EPA flanges that were found loose. When the flanges were retightened the leakage was reduced to an acceptable level.

Before further discussion on the development of probable leak areas, a brief introduction into the conditions and parameters.important to flow should be reviewed. A variety of flow conditions can exist in most leaks; however, one

" type _of flow usually predominates. According to R. C. McMaster (1982), the flow ris viscous when the flow rate (Q) is greater than 10 8 Pa m 3

/sec* (or 10 2 cc/sec), which is the situation for the leaks under discussion. In addition, ~

viscous flow is divided into two types of flow, laminar and turbulent. Turbulent flow occurs at the condition where the upstream pressure _is more than twice the downstream pressure (McMaster, 1982; Baumeister, 1978). The critical flow is then dependent on only the upstream pressure and temperature. The flow rate

, : equation for air at approximately ambient temperature is (Baumeister, 1978)

Q = 0.311 PAL.(in ft8/sec) ,

where: p = upstream pressure, psia

A = le k' area, in.2 L

4 or: Q = 8800 pA (in sec/sec)**'

Applying this equation to the Hatch 1 leak of 50 cc/ min (0.833 cc/sec), the cal-culated equivalent hole diameter is 0.00142 in. (2 x 10 6 in.2). This is small compared to' a hole that is needed for a leak rate of 0.1% of the containment volume per day, i.e., a 1-mm-diameter-hole (0.040 in.). The probable leak area in the Yankee Rowe< incident is derived solely on the Dasis of flange dimensions .

Assuming that'the~five flanges that leaked used a flange for a 12-in.-diameter nozzle, is: then the-0-ring groove is 17~in. in diameter and the total seal length L = (n)(c) = (5)(3.1416)(17) = 267 in.

If.the gap between the-loose flanges is assumed to be 1/2 mil (0.-0005 in.), then the assumed probable leak area will equal 0.0267 in.2, _

The maximum allowable operating leak rate may vary from plant to plant. The maxi-mum operating leak rate will depend on various parameters, which include con-tainment volume, number of EPAs, and permissible leaks in other types of pene-trations. The vendors design and fabricate EPAs according to IEEE Standard 317.

' .The most recent issue of the standard stipulates that the assembly can not have a leak rate greater than 10 3 cc/sec of dry nitrogen to qualify for shipment.

3

' *The units Pa m /sec are in SI units and are defined as the pressure in Pascals, volume in cubic meters, and time in seconds.

- scc /sec = standard cubic centimeters per second, t NUREG-1037 C-2 i

, o 9

'I

.The designjleak rate,after a DBA/MSLB accident exposure can not exceed 10 2 cc/sec-

-of dry nitrogen at design pressures and temperatures. -Table C.1, indicates a

~ comparison of permissible containment leak rates vs. expected maximum EPA 1eak rates.

O Table C.1 Comparison of containment and EPA operational leak rates

" Manufacturer ' Design- Expected EPA

.& number'of volume, Conditions. ~1eak rate **,

Plant. .EPAs (n) ft3x108 ILR*, cc/sec cc/sec

! Zion U'itsn 1 and 2 0'Brien (56/55) 2.7 885 0.56/0.55 Surry Units 1 and 2 Amphenol (32) ~1.2 393' "0.53 0'Brien (11) i Conax- ( 9)

' Peai:h Bottom' Units 2 . Physical ~ 0.3- '98.3 (?)

and 3- -

Sciences (?)

General-Electric (?)

Conax (?)

- Grand Gulf. Unit 1 Westinghouse (47) 0.3 98.3 0.47 Sequoyah Units 1 and 2 Westinghouse (47) 1.2 393 0.47

ILR - Integrated leak rate for containments. Assumed to be a 0.1% of contain-ment volume per day.

- This columnLis the number of EPAs (n) times the maximum allowable leak rate per

' day per assembly, which is 10 2 cc/sec of dry nitrogen at DBA pressures and tem-peratures.

PROBABLE LEAK AREAS UNDER SEVERE ACCIDENT CONDITIONS

-The potential for leak paths through EPAs under severe accident conditions has been studied in NUREG/CR-3234. Severe accidents.as used in that study refers to temperatures, pressures, and time durations that exceed the design-basis accident (DBA) conditions. The severe accident scenarios studied ~for EPA leak -

age.are those in:which both the fan coolers and sprays are not: functioning.

~This type of severe accident leads to a continuous increasexin pressure, which will eventually cause a loss of structural integrity of the containment building.

lThe ultimate capability of the containment building provides an upper bound to the maximum pressure loading to which EPAs'can be exposed.

_The-main conclusion of NUREG/CR-3234 was that properties of some of the materials used,in fabricating EPAs can degrade with temperature and time. 'Thus, if a pressure load occurs during a severe accident,.there is a potential.for-leak-

-age. The method of minimizing potential leak paths is to utilize materials.

with high temperature capabilities. The main 1 source of-potential leak paths

' occurs in those materials most sensitive to temperatures and these were deter-mined to be-the organic materials used in sealants,. gaskets', and 0-rings.

NUREG-1037 C-3 e

These organic materials or elastomers are also typically more sensitive to '

thermal and radiation aging, and other environments.

{

The probable leak paths in EPAs under severe accident conditions have only a !

very limited amount of experimental verification as is discussed in NUREG/CR-3234.

The only analytical efforts applied to this problem treat the thermal response of the EPAs under severe accident conditions. The thermal analysis is important because it is.used to calculate the response at the outboard end of an EPA to the inboard temperature and pressure profiles. If the assembly has sufficient length and mass, the outboard. temperature will lag considerably behind the inboard temperature. If at least one set of the EPA seals and/or sealants are at or below the design temperature, then the potential for leakage is expected to be low.

The temperatures and pressures have been calculated for a number of different severe accident progressions (NUREG/CR-3234; NUREG/CR-2182). One very important trend was noted for short duration, high temperature spikes: steam explosions

, and hydrogen burns do not impact EPAs. Another very important trend was noted for FWRs: the long-time temperature approaches a maxiium of 350 F. This is significant, because it suggests an upper limit to the temperature exposure.

The Mark III BWRs appear also to have the same temperature limit. However, for the Mark I (like Browns Ferry) and Mark II BWRs the inboard temperature can exceed 500 F.

In order to illustrate the contribution of EPAs to the potential for leakage, several containment buildings will be discussed. Although it would be desir-able to generate overall results on a generic basis, this is not possible because of differences in EPA designs that have been used in different contain-ment buildings. For example:

Zion 1 and 2 The EPAs installed in Zion 1 and 2 are all of the early D. G. O'Brien design.

These are canister type of EPAs and are welded to the containment nozzles. The containment nozzles are 12-in.-diameter schedule 40 steel pipe, which are welded in the liner plate. The EPA canisters are fabricated with 10-in.

schedule 20 steel pipe with lengths of 53 to 55 in. The canister designs are all double-ended; i.e. , a primary seal at each end of the canister, one at the inboard end and the second at the outboard end. The canisters have internal ports, so that pressure monitoring is conducted within the nozzles.

l' There are 56 EPAs in Zion 1 and 55 in Zion 2. Zion I has nine medium voltage power (MVP) canisters and Zion 2 has eight. The remaining 46 EPAs are for low voltage power, control, and instrumentation (LVP, LVC, LVI). The MVP conduc-tors use a ceramic insulation assembly which is brazed to the copper conductor (1000 MCM) and to another assembly at the header plate. The brazed connection is the primary seal for the MVP assemblies and is expected to have a long-term temperature capability.

The LVP, LVC, and LVI assemblies all have a similar type of construction. They are all double-ended and the primary seal for each conductor is provided with a glass-to-metal seal. These assemblies are in turn welded to'a 1/2-in.-thick

( header plate. All the conductors are additionally supported by potting materials j on aach side of the header plate. One side of the plate has 2 in of GE RTV 511 I

NUREG-1037 C-4 l'

and 3-1/4 in. of Emerson and Cummings 2651 potting compound. This potting pro-vides secondary sealing. Since the primary seal is glass, the aging effects will be minimal. Glass also has a temperature capability above the 350 F ex-pected in a severe accident and therefore no through-leakage is expected.

, Since the primary seals for all the Zion EPAs are non-organic, they are expected ta be affected very little by aging and will have a high temperature capability.

Surry 1 and 2 The Surry EPAs are designed to be mounted in an 8-in.-diameter schedule 40 nozzle. The nozzles are 59.75 in. long and are welded to a 1-in.-thick insert plate. There are 91 nozzles in the containment building, 90 located in one region between elevations 15 and 25 and one nozzle located at elevation 55.

(Not all the nozzles contain EPAs.) Three different manufacturers have sup-plied EPAs to Surry: Amphenol (1971-72), D. G. O'Brien (1972-73), and Conax (1981-82). Amphenol has been the major supplier: 64 EPAs, 32 to each contain-ment. The Amphenol design consists of a double-ended canister with connector sockets brazed into the header plates. The design has a hermetic seal in the header plate and at both ends of the assembly (i.e., a double-ended seal).

These assemblies are bolted to the flanges welded in each nozzle. Two 0-rings of silicon rubber are used for sealing the flanges. The remainder of the assem-blies are of welded construction.

D. G. O'Brien supplied 22 EPAs,11 each to Surry 1 and 2. This design uses glass-to-metal seals in the modules and should have a good long-term and high-temperature capability. The assemblies are bolted to the nozzle flanges and use double 0-rings as a primary seal. These 0-rings are silicon rubber, which has an estimated useful lifetime of two years at 400*F (Chivers and Hunt, 1978). This is above the long-term maximum temperature of 350*F expected in a severe accident.

Conax supplied 18 MVP assemblies to Sarry 1 and 2 (9 each) in 1981-82. These units were installed to improve the medium voltage service to the reactor cool-ant pumps. These EPAs have an odd configuration. They are bolted to outboard flanges and change from an 8-in. diameter to 12-in, diameter with 26-1/2-in.

extension. At the 12-in diameter, Cenax has three MVP modules mounted in the 12-in, blind flange. These modules utilize a ceramic bushing insulation that is brazed to a coupling assembly. The coupling assembly uses 0-rings to provide a seal to the header. These 0-rings are of an organic material (believed to be Viton). The seals are all outboard of the containment (i.e., single-ended).

In the event of any severe accident condition in the containment, the outboard location of the seals, provides some mechanism for thermal lag and cooling.

Peach Bottom 2 and 3 No information regarding EPAs was supplied by the utility and therefore no estimates of capability are possible at this time. The only information i

located to date has been a reference in the Franklin Institute Research Labora-

[ tory (FIRL) data bank. The data bank information states that the vendor was Crouse-Hinds and its operation was taken over by Westinghouse in the early 1970s.

(Telecon information, on October 12, 1984, with Philadelphia Electric staff corrects this information; we are told that Peach Bottom 2 and 3 contain Physical Sciences, General Electric, and Conax EPAs.)

NUREG-1037 C-5

Grand Gulf 1 The Mark III BWR type of containment of Grand Gulf 1 has a reinforced concrete containment building with a 3/8-in. steel liner plate. The concrete wall is 42 in. thick and all the EPA nozzles are 48 in. long. The nozzles are welded to the steel liner plate, which has been thickened locally to 1/2 in for the EPA nozzles. There are 47 Westinghouse EPAs in Grand Gulf 1. Three are 18 in.

in diameter and 44 are 12 in. in diameter. Two of the 18-in.-diameter EPAs supply medium voltage power (to the reactor coolant pumps) and contain three modules, each with a 1000 MCM conductor. The EPAs consist of a canister 51-1/4 in. long and three modules that incorporate a ceramic bushing which is brazed to an assembly that in turn is welded to the header plate. The canisters are welded to the nozzle at the outboard end. The MVP EPAs are double-ended and have a ceramic bushing at both the inboard and outboard end. The pressure is maintained and monitored in the interior of tha canister.

There are forty-four 12-in.-diameter EPAs and all contain three standard Westing-house modules. These modules are 5 in. i- diameter and 12 in, long. The con-ductors pass through two glass-epoxy ple that are mounted in the midplane of the modules. .These two plates have a sn I separation, so that each conductor in the interior of the module can be r 'onitored. The remainder of the modules are filled with an epoxy sealant. epoxy sealant also serves as potting material for restraining cables and conductors. The potting material is contained by a stainless steel can (shroud) that is welded to a ring at the midplane. The ring contains four 0-ring grooves and a small retaining flange.

The flange is clamped with lugs to the header plate. A passage is provided between the 0-ring grooves, so that the internal module pressure can be moni-tored. The inrer 0-rings are EPDM rubber, while the outer 0-rings are silicon rubber. This same passage connects to pressure ports in the header plate, sc that all three modules can be mnnitored concurrently.

There is one additional EPA that is 18 in. in diameter and contains five modules as described above, instead of three. The design and fabrication is identical to the three-module EPA and any conclusions regarding those nodules would apply to this EPA.

In this particular Mark III BWR, the suppression pool is located below all the EPAs.

Therefore, in a severe accident the suppression pool will appreciably reduce the temperatures exposed to the EPAs, even though the drywell may experience high temperatures.

The EPAs also have at least one primary seal located outboard, so that any tem-perature response of the outboard seals will lag behind the inboard environment.

The design and fabrication of the Westinghouse EPAs should have a substantial margin to survive the maximum severe accident temperature and pressure (*350 F and 55 psig). The ultimate pressure capability of the containment building has been estimated to be 55 psig, and therefore, provides a limiting pressure on the EPAs.

The limiting parts of the EPAs are most likely to be the 0-rings, because they are fabricated from organic materials. These materials will creep with age (and NUREG-1037 C-6

temperature) and increase the probability of leakage with age. (Some of these same EPAs are stored in a warehouse, where pressure monitoring is maintained and some of the EPAs have experienced pressure reductions during this period of storage. However, the cause of these leaks has not been identified or no estimate of the'1eak rate is available. The volume being monitored is small, approximately 238 cc, so that even at the permissible leak rate the pressure would drop in a few weeks, unless the pressure is maintained.)

Sequoyah 1 and 2 The Sequoyah plants are.a PWR containment with an ice condenser type of suppres-sion system. This type of design typically has a large steel containment shell.

The steel shell is approximately 5 to 6 ft within a thick concrete building that functions primarily as a biological shield. The nozzles in the containment shell are typically short, in this case only 2 ft long. The primary vendor for the Sequoyah EPAs was Westinghouse. The EPAs were supplied to Sequoyah in 1973.

These EPAs were of a canister design (the Westinghouse modular design was developed in 1975). There are 47 EPAs in each plant; 4 for MVP and 43 for LVP, LVC, and LVI. The MVP assemblies are welded into 18-in.-diameter schedule 80 steel pipe. The other 43 assemblies are welded into 12-in.-diameter schedule 80 pipe. The MVP modules incorporate s ceramic bushing at each 6nd of the canister. The LVP, LVC, and LVI canisters have a complex mixture of primary sealants. Some used epoxy and some were hermetically sealed. The hermetic seals consisted of connector shells welded to the header plate. The EPAs with only epoxy seals have a low-temperature capability because the formulation used by Westinghouse in the mid-1970s had a low-temperature capability. The EPA temperature test specification was 15 minutes at 300 F and 15 psig; then 250 F and 12 psig for 10 days. The 15 minutes at 300 F, because of thermal lag, would have little effect on the sealants. The peak temperature experienced by the sealants would be 250"F.

The Sequoyah plants also have one EPA that is a Westinghouse modular design, but details are not available. In addition, tha personnel airlocks contain a Conax module type of penetration seal.

In a severe accident, the containment atmosphere will tend to be in a saturated steam state. This means that the temperature and pressure are not independent and can be established by knowing either value. Since the Sequoyah containment 1 shell has one of the lowest estimated ultimate pressure capacities (60 psig), the corresponding temperature will be limited too. At 75 psia saturated steam, the temperature is approximately 308 F. This is above the qualifying temperature and the EPAs with the epoxy primary sealants will have a limited useful lifetime and may result in a higher potential for leakage.

A brief description of several containment buildings and their EPAs was presented above as background information. The next problem is'to develop " probable leak areas" for EPAs under severe accident conditions. At this time only an estimate based on engineering judgment can be provided,.since the most realistic approach to this question is to measure leak rate in an appropriate experimental program.

The " probable leak areas" for EPAs which are based on a pressure criterion are not appropriate. The behavior we expect from the EPA sealants, gaskets, and 0-rings is to be sensitive to temperature and time. In general, the EPA seals are separated far enough from the deformations in the containment wall, so that NUREG-1037 C-7 m . J

during a severe accident, minimal changes are transmitted to the EPA seals and hence do not change with pressure. In addition, 0-rings are used extensively

.in header, plates and a good 0-ring design should be capable of 1000 psig pres-sure load. This is considerably higher than the ultimate capacity of contain-ment buildings (*60 to 200 psig).

As noted earlier, the EPAs are designed to have a leak rate less than 10 2 cc/sec after a DBA. The equivalent area for this leak rate is 1.7 x 10 8 in.2 (0.00015-in.-diameter hole), which is very small. To be conservative, we can assume that a probable leak area is much larger, for example 10 5 to 10 7 in.2, After the EPA sealants reach their limiting temperature, we can assume that the probable leak areas will increase two orders of magnitude, 10 3 to 10 5 in.2, because of the degradation of the organic materials. The EPAs with glass-to-metal seals are assumed to have the lower leak area values, the organic seals will have the larger leak areas. To arrive at the total leak area, these values would be multiplied by the number of EPAs in each containment. Table C.2 con-tains estimated probable leak areas per penetration.

Table C.2 Probable leak areas in EPAs Estimated area per EPA

Sealant / Limiting gasket & temp. of Before, After, Plant (no.) of EPAs seal, *F in.2 in.2 Zion 1 & 2 Glass / welded 500 10 7 10 7 (56 & 55)

Surry 1 & 2 Glass /sil (32) ? 10 5 10 3 Glass /sil (11) 500/400 10 5 10 3 Polysul/vit (9) 340 10 5 10 3 Peach Bottom Epoxy (?) 250 10 5 10 3 2&3 Glass / metal (?)

Grand Gulf 1 Epoxy /sil (47) 350 10 5 10 8 Sequoyah 1 & 2 Epoxy / welded 250 10 5 10 3 ;

(47)

Based on engineering judgment.

The estimates of " probable leak areas" in Table C.2 can be used to estimate the total values under severe accidents. These results are tabulated in Table C.3 and it assumes that all EPAs contribute equally to the flow.

The holes listed in Table C.3 are quite small and their response to flow with particulates is expected to be substantially different (Morewitz, 1982).

I NUREG-1037 C-8

Table C.3 Probable leak areas for the reference containments

_ Estimated total probable Containment leak area, in.2

-failure pressure,

Plant psia

~

Low (1) Mid (2) High (3)

Zion 1 & 2 149 0.00015 0.00056 0.0012 Surry 1 & 2 100 0.00015 0.052 0.104 Peach Bottom 2 & 3 119 - - -

Grand Gulf 1 70 0.00015 0.047 0.094 Sequoyah 1 & 2 75 0.00015 0.047 0.094 (1) Low - Area based on assumed area before an accident is initiated.

(2) Mid - Area based on assumed area after cn accident is initiated.

(3) High - Assumes that this area is twice the mid value.

REFERENCES Baumeister,.T., Ed., Marks Standard Handbook for Mechanical Engineers, Eighth Edition, McGraw-Hill Book Company, 1978, pp. 4-47 & 4-48.

Chivers T. C., and R. D. Hunt, " Scaling of Cas Leakage From Static Seals,"

8th International Conference on Fluid Sealing, September 11-13, 1978, paper G3.

Letter from NSIC to W. A. Sebrell, SNL, dated February 1983.

. McMaster, R. . C. . Ed. , " Leak Testing," Nondestructive Testing Handbook, Vol.1, Chapter on Mathematical Theory of Gas Flow Through Leaks, pp.89-100, American Society of Metals,1982.

Morewitz, H. A., " Leakage of Aerosols From Containment Buildings," Health Physics, Vol. 42, No. 2, February 1982, pp. 195-207.

U.S. Nuclear Regulatory Commission, NUREG/CR-2182, D. H., Cook et al., " Station Blackout at Browns Ferry Unit One - Accident Seq"7nce Analysis," ORNL/NUREG/

TM-455/V1, Oak , Ridge National Laboratory, Oak Ridge, Tennessee, November,1981.

-- , NUREG/CR-2549, T. E. Blejwas et al., " Background Study and Preliminary Plans for a Program on the Safety Margins of Containments," SAND 82-0324, Sandia National Laboratories, Albuquerque, New Mexico, May 1982.

-- , NUREG/CR-3234, W. Sebrell, "The Potential for Containment Leak Paths Through Electrical Penetration Assemblies Under Severe Accident Conditions,"

SAND 83-0538, Sandia National Laboratories, Albuquerque, New Mexico, June 1983.

NUREG-1037 C-9 a

{ 3

,sRC PORM 335 - U S. NUCLE A A 4 ElutATORY COMutf4'ON t v tsOwT NuustR #Ampnadat TIDC. ### Foi N 9, d enyt hb'df BIBLIOGRAPHIC DATA SHEET NUREG-1037

$f E INSTRUCTIONS ON THE RivtR54 -

2. TITLE AND SU5flTLE 3 LE AVE BLANK .

C nttinment Performance Working Group Report, Draft Report for Comment 4 OATE REPORT COMPLETED MON T a's DEAR

. Av7 ORis, May 1985

. Containment Performance Working Group ,o,,,, ,,AN May b 1985

7. PE ~PORM1NG ORGANil ATION NAME AND MalLING ADDRE55 uac s

va le Cast 8 PROJECTiT A$K WORK uN#I NUMgtR Divit, ion of Engineering Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission Washington,.D.C. 20555 10 SPON50 RING ORGANil ATpON N AME ANO MAsLINu ADDRES$ ifactwele taaes 11. T YPt OF REPOR Y Division of Engineering Technical Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission =PE-*CO eEOa ~ m a Washington, D.C. 20555 12 SUPPLEMENT ARY NOTES t1 A8ST,.ACT fAIO we*as er sesss Containment buildings for power reactors have been studied to estimate their leak rate as 'a function of increasing internal pressure and temperature associated with severe accident sequences involving significant core damage. Potential leak paths through containment penetration assemblies (such as equiprent hatches, airlocks, purge and vent valv s, and electrical penetrations) have been identified and their contributions to leak area for the containment are incorporated into containment leak rate and pressure /-

temp:rature response as a function of time.

~

Because of lack of reliable experimental data. on the leakage behavior of centainment penetrations and isolation barriers at pressures beyond their design conditions, an analytical approach has been used to estimate the leakage behavior of components found in sp cific reference plants that approximately characterize the various containment types.

i.oOCu .NTANA6,...... mORoio..c , TOR. ,A.,. 31 Leak rate Un1imited Cdntainmentperformance .

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ML20126C960

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