ML20030D622

Mark II Containment Program Evaluation and Acceptance Criteria

ML20030D622
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Issue date:	08/31/1981
From:	Anderson C Office of Nuclear Reactor Regulation
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References
REF-GTECI-A-08, REF-GTECI-CO, TASK-A-08, TASK-A-8, TASK-OR NUREG-0808, NUREG-808, NUDOCS 8109140048
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MARK ll Containment Program Load Evaluation and Acceptance Criteria Generic Technical Activity A-8

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Manuscript Completed: July 1981 Date Published: August 1981 C. Anderson Division of Safety Technology Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission Washington, D.C. 20666 f.

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ABSTRACT This report, prepared by the staff of the Office of Nuclear Reactor Regulation, arovides a discussion of LOCA-related suppression pool hydrodynamic loads in

) oiling-water reactor (BWR) facilities with the Mark II pressure-suppression containraent design.

This report concludes NRC Generic Technical Activity A-8,

" Mark II Containment Pool Dynamic Loads," which has been designated an

" Unresolved Safety Issue" pursuant to Section 210 of the Energy Reorganization Act of 1974.

On the basis of large-scale tests conducted in 1979, the Mark II Owners developed improved condensation-oscillation and chugging loads for the suppression pool boundary and lateral loads for the containment downcomers.

The staff has reviewed these proposed loads and concluded that, with a few specified changes, these loads provide conservative loading conditions.

In addition, the staff has conducted a study which confirms that the lead plant pool-swell loads, adopted by the Mark II Owners as the final load specifications, are conservative.

This study used the results of full-scale multivent Mark II tests conducted in Japan.

The staff acceptance criteria for pool-swell loads from the Lead-Plant Program and new criteria for steam loads developed in the Long-Term Program have been consolidated and constitute Appendix A of this report.

ill

CONTENTS Page ABSTRACT........................................................

iii FOREW0RD........................................................ ix ACKNOWLEDGMENTS................................................ xi NOMENCLATURE....................................................

xiii 1

INTR 000CTION..............................................

1-1 1.1 Problem Definition...................................

1-1 1.2 Lead-P l ant P rogram..................................

1-3 1.3 Long-Te rm P rog ram....................................

1-3 2

HYDRODYNAMIC-LOAD EVALUATION..............................

2-1 2.1 Pool-Swell Loads.....................................

2-1 2.1.1 Background.................................

2-1 2.1.2 Evalution of Pool-Swell Loads..............

2-2 2.1.2.1 Submerged-Boundary Loads During Vent Clearing.............

2-3 2.1.2.2 Air-Bubble Pressure..............

2-3 2.1.2.3 Pool Velocity....................

2-4 2.1.2.4 Pool Acceleration................

2-4 2.1.2.5 Submerged-Boundary Pressure During Bubble Expansion..........

2-5 2.1.2.6 Pool Elevation...................

2-5 2.1.2.7 Diaphragm-Floor Reverse Pressure.

2-7 2.1.2.8 Wetwell-Airspace Pressure........

2-8 2.1.3 Eva l ua ti o n S umma ry.........................

2-9 2.2 Steam-Condensation and Chugging Pool Boundary Loads..

2-9 2.2.1 Condensation-Oscillation Load..............

2-10 2.2.1.1 Background.......................

2-10 2.2.1.2 CO Load Description..............

2-10 2.2.1.3 CO Load Evaluation...............

2-14 2.2.1.4 Evaluation Summary...............

2-20 2.2.2 Improved Chugging Load.....................

2-20 2.2.2.1 Background.......................

2-20 2.2.2.2 Chugging Load Description........

2-21 2.2.2.3 Chugging Load Evaluation.........

2-24 2.2.2.4 Evaluation Summary...............

2-33 v

M 2.3 Downcocer Lateral Loads..............................

2-33 2.3.1 Background......................

2-33 2.3.2 Lateral Load Description...................

2-34 l

2.3.2.1 Single-Vent Load.................

2-34 2.3.2.2 Lateral Load on Multiple Vents...

2-35 2.3.3 Lateral Load Evaluation....................

2-36

' 3.3.1 Data Base........................

2-36 1.3.3.2 Single-Vent Loads................

2-37 2.3.3.3 Lateral Load on Multiple Vents...

2-41 2.3.4 Eval uati o n S umma ry.........................

2-44 2.4 Submerged-Structure Drag Loads.......................

2-44 2.4.1 Background.................................

2-44 2.4.2 LOCA-Water-Jet Ring Vortex Model Evaluation. 2-45 2.4.3 Evaluation Summary.........................

2-47 1

3 RESO LUTION OF THE ISSUE...................................

3-1 3.1 Implementation of Guidel ines.........................

3-1 3.2 Recommended Changes to the Standard Review Plan......

3 3.3 Recommended Development of a Regulatory Guide........

3-1 4

REFERENCES................................................4-1 APPENDIX A - NRC ACCEPTANCE CRITERIA FOR PARK II LOCA-RELATED P0OL DYNAMIC LOADS APPENDIX B - MARK II OWNERS GROUP SUPPORT PROGRAM l

APPENDIX C -

SUMMARY

OF MARK II LOCA-RELATED POOL DYNAMIC LOADS i

e i

vi

l LIST OF FIGURES Figure Page l

2.2-1 4TCO Test Facility.....................................

2-11 l

2.2-2 JAERI Multivent Test Facility - Elevation View.........

2-14 2.2-3 JAERI Multivent Test Facility - Plan View..............

2-16 2.2-4 Wetwell Pressure Spatial Distribution..................

2-17 l

I LIST OF TABLES r

1 l

Table P_ age 1.1-1 Listing of Domestic BWR Facilities with the Mark II Containment System.....................................

1-2 l

l 2.1-1 JAERI Pool-Swell Tests.................................

2-2 2.2-1 4TC0 Test Matrix.......................................

2-12 2.2-2 JAERI Test Matrix......................................

2-13 2.2-3 4TC0 Time Periods for Basic C0 Load....................

2-14 2.2-4 4TCO Time Periods for CO Load with ADS.................

2-14 i

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vii 1

I

FOREWORD NUREG-0808 is being issued to describe Mark II LOCA-related pool dynamic loads that the NRC has reviewed and found acceptable to meet the requirements of General Design Criterion 16 in Appendix A to 10 CFR Part 50.

NUREG-0808 is not a substitute for the regulations, and compliance is not a requirement.

However, an approach or method different from the loads referenced herein will be accepted only if the substitute approach or method provides a basis for detereining that the above-cited regulatory requirements have been met.

ix

E-1 L

ACKNOWLEDGEMENTS l

A-8 Review Team The following individuals participated in the Mark II containment Long-l Term Program Safety Evaluation and contributed substantially to this report:

C. Anderson, USNRC, Division of Safety Technology (A-8 Task Manager) l J. Lehner, Brookhaven National Laboratory l

G. Maise, Brookhaven National Laboratory l

T. Ginzbury, Brookhaven National Laboratory A. Sonin, Massachusetts Institute of Technology R. Scanlan, Princeton University G. Bienkowski, Princeton University I

l l

t l

l l

t i

xi

t NOMENCLATURE ADS Automatic Depressurization System BCP Bottom-Center Pressure BNL Brookhaven National Laboratory BWR Boiling-Water Reactor CO Condensation Oscillation DFFR Dynamic Forcing Function Report EPRI Electric Power Research Institute FSI Fluid-Structure Interaction GE General Electric Company GKM Grosskraftwerk Mantiheim GKSS Gesellschaft fur Kurnenergieverwertung in Schiffbau und Schiffahrt IWEGS Inhomogeneous Wave Equation Green's Function Solution, computer program for cylindrical geometry IWEGS/ MARS Similar to IWEGS with Modified Annular Region Solution JAERI Japan Atomic Energy Research Institute JMARS Similar to IWEGS with Modification To Solve a 20' Sector of an Annulus KWU Kraftwerk Union LLL Lawrence Livermore Laboratory LOCA Loss-of-Coolant Accident LPP Mark II Lead-Plant Program LTP Mark II Long-Term Program MSP Mean Square Power FOP Peak Overpressure FSAM Pool-Swell Analytical Model PSTF Pressure-Suppression Test Facility QSTF Quarter-Scale Test Facility RMS Root Mean Square SRP Standard Review Plan SRV Safety / Relief Valve 4T Temporary Tall-Tank Test 4TCO Tempoiary Tall-Tank Test - Condensation Oscillation viii

l 1 INTRODUCTION l

l Pursuant to Section 210 of the Energy Reorganization /tet of 1974, the capability of the boiling-water reactor (BWR) Mark II containment to withstand loss-of-I coolant accident (LOCA)-related pool dynamic loads, which were not considered l

in the original containment design, was designated an " Unresolved Safety Issue" l~

-(Task Action Plan (TAP) A-8).

This report, along with three previous NRC reports,t 2.s describes the generic hydrodynamic loads to be used to evaluate i

l BWR/ Mark II facilities.

The NRC and its consultants have reviewed the applicable l

exp rimental and analytical programs, and have concluded that the proposed LOCA hydc.idynamic loads referenced in Appendix C, as modified by the requirements sec, forth in Appendix A ("NRC Acceptance Criteria for Mark II Containment"),

will provide a conservative evaluation of the containment structures, piping, l

and equipment for suppression pool hydrodynamic LOCA loading.

These loads constitute the resolution of TAP A-8.

t 1.1 Problem Definition l

In the United States there are 11 BWR facilities in various stages of construction which have the Mark II containment system.

About half of these are currently l

scheduled for operation by the end of 1982.

A listing of the domestic BWR facilities with the Mark II containment system is provided in Table 1.1-1.

The original design of the Mark II containment system considered only those loads normally associated with design-basis accidents.

These included pressure 6nd temperature loads associated with a LOCA, seismic loads, dead loads, jet impingement loads, hydrostatic loads due to water in the suppression chamber, overload pressure test loads, and construction loads.

However, since the establishment of the original design criteria, additional loading conditions have been identified that must be considered for the pressure-suppression containment-system design.

In the course of performing large-scale tasting of an advanced design pressure-suppression containment (Mark III), and duriig inplant testing of Mark I containments, new suppression pool hydrodynamic loads were identified that had not been included explicitly in the original Mark II containment-design basis.

These additional loads result from dynamic effects of drywell air and steam being rapidly forced is'.o the suppression pool during a postulated LOCA and from suppression pnol r esponse to safety / relief valve (SRV) operation, which is generally associateu with plant transient operating conditions.

Because these new hydrodynamic loads had not been considered, the NRC staff determined that a detailed reevaluation of the Mark II containment system was required.

The Mark 11 containment design was based on the experimental technology obtained l

from testing performed on a pressure-suppression concept for the Humboldt Bay

{

Power Plant and from testing performed for the proposed Bodega Bay Plant concept.

The purpose of these initial tests, performed during 1958 through 1962, was to demonstrate the viability of the pressure-suppression concept foi reactor containment design.

Tests were designed to simulate a LOCA with l

various equivalent piping break sizes up to a break approximately twice the i

cross-sectional size of the design-basis LOCA.

The tests were instrumented to obtain quentitative information for establishing containment design pressures.

Data from these tests were the primary experimental bases "or the design and the initial staff approval of the Mark 11 containment system.

l 1-1

i 1

Table 1.1-1 Listing of Domestic BWR Facilities with the Mark II Containment SysteA Plant Name Applicant Bailly 1 Northern Indiana Public Service Co.

Chesterton, Indiana WPPSS-2 Washington Public Power Supply System Richland, Washington LaSalle 1 and 2 Commonwealth Edison Company Chicago, Illinois Limerick 1 and 2 Philadelphia Electric Company Philadelphia, Pennsylvania Nine Mile Point 2 Niagara Mohawk Power Company Syracuse, New York Shoreham Long *!sland Lighting Company Hicksville, New York Susquehanna 1 and 2 Pennsylvania Power and Light Company Allentown, Pennsylvania Zimmer Cincinnati Ges and Electric Company Cincinnati, Ohio During the large-scale testing of the Mark III containment system design in the period 1972 through 1974, new suppression pool hydrodynamic loads were identified for the postulated LOCA.

General Electric (GE) tested the Mark III containment concept in its Pressure Suppression Test Facility (PSTF)4 These tests were initiated for the Mark III concept because of the geometrical configuration differences between the previous containment concepts and the Mark III design, principally in the utilization of horizontal vents.

(Steam had been ejected vertically downward into the suppression pool in the previous BWR containment designs, whereas the Mark III design ejects steam horizontally into the suppresion pool.) More sophisticated instrumentation was available for the Mark III tests, as were computerized methods for data processing.

It was from the PSTF testing that the short-term dynamic effects of drywell air being forced into the pool in the initial stage of the postulated LOCA were first clearly identified.

In addition to the information obtained from the PSTF data, other LOCA-related 5 for similar dynamic load information was obtained from foreign testing programs pressure-suppression containments.

It was from these foreig:i tests that steam 1-2

condensation loads on the vent system downcomers and suppression pool boundaries during the later stages of steam-vent flow were first identified.

Consequently, In April 1975, the NRC sent letters to each of the domestic utilities having BWR facilities with Mark II containment system designs re-questing that they provide information demonstrating the adequacy of their containment design.

These letters reflqcted NRC concerns about the need to evaluate the containment response to the newly identified dynamic loads associated with postulated design-basis LOCA.5 The domestic Mark II containment owners formed an ad hoc Mark II Owners Group to develop responses to these NRC requests.

They developed a two part program consisting of the Lead-Plant Program (LPP) and the Long-Term Program (LTP) to accommodate the licensing needs of the lead and the following Mark II plants.

These programs are described below.

1. 2 Lead-Plant Program Licensing activities for certain Mark II lead plants (Zimmer, Shoreham, and LaSalle) were originally scheduled to precede completion of the entire Mark II containment program.

Consequently, the LPP was develope!d to demonstrate that suf ficient information about and understanding of the pool dynamic phenomena of interest existed to establish conservative loads for the lead plants.

Because of the LPP emphasis on developing loads consistent with the licensing requirements of the lead plants, a bounding interpretation of the available test data was utilized for many of the pool dynamic loads.

This was done to ensure that conservative loads were available for the lead plant evaluations.

The NRC staff reviewed the Mark II owners' lead plant program and identified acceptable pool dynamic loads to be used in the evaluation of Mark II pressure-suppression containment designs for these plants. These loads were discussed in the staff report, " Mark II Containment Lead Plant Program Load Evaluation and Acceptance Criteria," (NUREG-0487)1 issued in October 1978.

Supplement I to this report' was issued in October 1980.

It addressed acceptable alternative loads to the original NUREG-0487 lead plant loads.

Supplement i also identified problems relating to the lead plant condensation-oscillation (CO) and chugging loads.

Large-scale testing applicable to the Mark II design which was conducted in 1979 indicated deficiencies in the lead plant CO and chugging loads.

As a result, the Mark II owners developed interim CO and chugging loads to reflect the results of these tests.

The staff's evaluat. ion of these loads was provided in Supplement 2 to NUREG-04878 Supplement 2 was issued in February 1981.

It completed the Lead-Plant Program.

1. 3 Long-Term Program The objectives of the LTP were (1) to provide justification, by tests and analyses, for refinement of selected lead plant bounding loads, and (2) to provide additional confirmation of certain loads used in the LPP.

The Mark II owners elected to adopt most of the pool-swell loads developed during the LPPt.2 An exception was the diaphragm-floor upload specificatior.

The Mark II owners revised the LPP floor upload specification to reflect a deficiency in the earlier specification.

Since completion of the LPP, the 1-3

a Japan Atomic Energy Research Institute (JAERI) has planned an extensive series of full-scale tests using a 20* sector of a Mark II containment, which to date have been partially completed.

The NRC and its consultants have utilized the results of these tests to confirm the Mark II LPP pool-swell acceptance criteria.

The staf f'n evaluation of the Mark II owners' revised diaphragm-floor upload specification and the results of the staff pool-swell-loads confirmation study are provided in Section 2.1 of this report.

During the LPP, the Mark II owners concluded that steam condensation pool dynamic loads should be developed that were more rigorous than the bounding loads used in the LPP.

These loads included the CO and chugging loads on the pool boundary and the chugging-induced lateral loads on the containment down-Large-scale tests were conducted in 1979 to provide additional informa-comers.

tion related to these loads.

These tests formed the basis for the LTP steam loads.

The staff evaluation of these loads is provided in Sections 2.2 and 2.3 of this report.

Ths Park II owners proposed an alternate load specification for the LPP LOCA-ralated submerged structure drag load in the form of a ring-vortex model.

This ring-vortex model considers the drag loads that occur during the water-jet and the air bubble periods following a LOCA.

The staff's evaluation of this model is provided in Section 2.4 of this report.

The LOCA-related suppression pool nydrodynamic acceptance criteria for the Mark II containment design are provided in Appendix A.

This appendix contains the applicable LOCA-related criteria from the LPP and the revised pool-swell and steam-load criteria resulting from the LTP.

Appendix B contains a list of the reports transmitted to the NRC during the course of the lead plant and long-term programs.

A summary of the LOCA pool dynamic loads acceptable to the NRC is provided in Appendix C.

This report concludes the generic Mark II containment program.

The loads referenced in Appendix C, as modified by the acceptance criteria provided in Appendix A, constitute the resolution of Unresolved Safety Issue A-8.

1-4

2 HYORODYNAMIC LOAD EVALUATION 2.1 Pool-Swell Loads 2.1.1

Background

In the first few seconds following a postulated LOCA a mixture of air and steam is carried through the containment downcomers into the suppression pool.8 The loads associated with the transfer of the air / steam mixture are referred to as pool-swell loads.

Pool-swell loads for the Mark II containment design were developed during the Mark II LPP.7 As a result of this work, the NRC formulated acceptance criteria for Mark II lead plant pool-swell loads 'z, 1

These conservative loads developed during the LPP are to be adopted.by all Mark II plants. With only one exception, the Mark II LTP excluded further work on pool-swell loads.

The Mark II owners conducted an additional investigation into diaphragm reverse pressure loads.

Since completion of the Mark II LPP, JAERI has planned an extensive series of full-scale tests using a 20" sector of the Mark II pool.

At this time the tests have been partially completed.

Because of similarities among all Mark II plants, these new tests provide an opportunity for additional verification of the lead plant pool-swell loads.

NRC consultants at Brookhaven National Laboratory (BNL) have conducted a verification study of the lead plant pool-swell load"*.

The study used the results of six separate pool-swell tests of the JAERI test facility.s-ta These are listed in Table 2.1-1.

Test No. 1101 is a liquid blowdown test, while the remaining five are all steam blowdowns.

In conducting these comparisons, it was decided to examine all pool-swell-related parameters in the NRC acceptance criteria and to compare them against the pertinent JAERI data.

The general approach taken in comparing the JAERI data to NRC criteria was the following:

The JAERI test apparatus was treated as if it were an actual plant.

The pool-swell' containment loads (and other parameters) were calculated by BNL according to the specifications in the NRC acceptance criteria.

Next, the corresponding loads (and other parameters) were determined frte direct measurements in the JAERI tests.

The leads were then compared.

If the loads from the NRC specifications exceeded the measured values, the acceptance criteria were confirmed under conditions simulating a Mark II LOCA.

Some of the pool swell parameters, specified in the NRC acceptance criteria, are based on drywell pressure histories that are calculated with the GE Containment Analytical Model.84 This computer program calculates conserva-tively high pressure histories in the containment drywell.

The code is propri-etary and was not available for use by BNL.

The drywell pressure history is needed as input to the Pool Swell Analytical Model (PSAH)7 and is also used in the specification of maximum pool elevation.

Pecause the GE code was not used to predict drywell pressure histories for the specific JAERI test runs under

• J. Lehner, " Mark II Containment Load Evaluation sport, "USNRC NUREG/CR repcrt in preparation, 2-1

l Table 2.1-1 JAERI Pool-Swell Tests

• l l

Test Type Nozzle Submergence T(Pool)

No.

Blowdown Dia (mm)

(m)

(*C) 1101 Water 200 3.63 29.9 1201 Steam 200 3.33 52.3 1202 Steam 240 3.24 53.5 1203 Steam 220 3.34 53.9 l

1204 Steam 220 3.34 18.7 l

1205 Steam 220 3.35 53.5 I

i

• All tests have nominally the same vessel conditions t

(P=6900 KPa; T=286'C).

sxamination, a different approach is needed to determine pool-swell parameters corresponding to the NRC criteria.

The following approach was adopted:

The actual measured drywell histories from the JAERI tests (instead of GE model calculations) were used as input to the pool-swell model and elsewhere in the NRC specifications.

Because these drywell pressure histories are not as severe as those produced by the GE model, one may assume that, if the NRC specifications bound the pool-swell parameters using taasured drywell pressures, they will also bound them if the GE-model pressurization is used.**

On the other hand, if the measurements exceed the specifications obtained in this canner, as was the case for maximum pool height, a different, better, estimate tpproach is needed.

2.1.2 Evaluation of Pool-Swell Loads This section includes the staff's evaluation of the pool-swell-related parameters in the NRC lead plant acceptance criteria.tez These parameters include:

submerged boundary loads during vent clearing air-bubble pressure pool velocity submerged boundary pressure during bubble expansion pool elevation diaphragm-floor reverse pressure wetwell-airspace pressure

• • The diaphragm floor upload specification is an exception to this generalization.

z-2

This evaluation was conducted by BNL.

The evaluation consists of a comparison of the lead plant acceptance criteria against data from the JAERI pool-swell tests.

2.1.2.1 Submerged Doundary Loads During Vent clearing The vent-clearing phenomenon following a LOCA results from the clearing of water from the main-vent downcomers due to drywell pressurization.

As a result of the phenomenon, pressure loads are produced on the containment basemat and the submerged-wetwell wells.

stipulate that the overpressure (above local 2

The NRC acceptance criteria hydrostatic) on the basemat and walls below the vent exit is determined as follows:

First, defir.e a parameter:

F = (mhl)/[(A /A )V 3 p

DW

- mass flow (1b/sec) m 8

V

- drywell volume (f t )

DW h

- enthalpy - (Stu/lb)

L

- submergence (ft)

A E

7

- pool area to vent area ratio V

now:

P

= 24 psi when F < 55, and P

= 24 + 0.27 (F-55) when F > 55 On the pool walls between the pool surface and the bottom of the downcomers, the overpressure is stipulated to vary linearly, with zero overpressure at the pool surface.

These criteria were used to calculate the JAERI boundary loads for the tests listed in Table 2.1-1.

Use of the criteria leads to an overprediction of the boundary pressure loads measured in JAERI by a factor that ranges from 1.5 to 2.5 for submerged boundaries below the vent exit plane.

The margin is reduced above the vent exit elevation.

This comparison confirms that the vent-clearing acceptance criteria are conservative.

2.1.2.2 Air-Bubble Pressure l state that the air-bubble pressure, during pool The NRC acceptance criteria swell, is to be calculated using the PSAM code.

The BNL equivalent of PSAM was, therefore, used to calculate these pressures.

2-3

Experimental measurements of air-bubble pressure are not directly available from the JAERI data.

However, if the bubble pressure is interpreted as it is in the PSAM code (that is, the pressure pushing up a piston-like slug of water), then the wall pressures at the elevation corresponding to the ends of the downcomer are, in fact, equal to the bubble pressure.

Consequently, the average of the maximum of the four wall transducers at the 12-foot elevation was taken to be the air-bubble pressure.

These values of bubble pressure were compared to the NRC specifications with the result that the calculated pressures are approximately equal to or slightly higher than the measurements.

This confirms that the air-bubble pressure criterion is conservative.

2.1.2.3 Pool Velocity i

The NRC acceptance criterial require that the pool velocity be calculated l

using the PSAM code.

The velocity thus obtained is then multiplied by a factor of 1.1.

The drywell pressure history, required as input to the PSAM code, is to be calculated using the GE containment model.14 BNL has produced a pool-swell code which duplicates the GE PSAM code.

This code was exercised for the conditions of the JAERI tests.

Because the l

calculated drywell pressure histories were not available, the measured histories were used as the driving force for the pool-swell code.

The use of measured histories as input, instead of calculated, is expected to lead to lower pool l

velocities.

The measured pool velocities from the JAERI tests were obtained from the level probe readings.

This involved numerical differentiation of experimental data, with the inherent amplification of experimental errors.

To avoid this problem, l

the data were first smoothed analytically.

The position of the surface, as a j

function of time, was represented by a third-order polynomial, which was fitted to the data using a least-squares fit.

Because the pool surface was found to have significant curvature (it rose faster in the middle than at the inner and outer walls), two fits were obtaineu for each test run.

One con-sidered all the points which led to a geomet I: ally averaged pool velocity; the other considered the level probes near t'.0 center of the pool (level probes at stations L2 and L8) which yielded the center pool velocity.

l Haximum pool velocities were calculated for each of the tests in Table 2.1-1.

A comparison of the calculated PSAM pool velocity to the measured maximusc center pool velocity shows calculated values abort 10 to 40% higher than the measured values.

An even greater margin is observed for the comparison with the average pool-surface velocity.

This comparison confirms that the pool velocity acceptance criterion is conservative.

2.1.2.4 Pool Acceleration 1

The NRC specification for pool acceleration stipulates that the acceleration be determined by application of the PSAM code.

It is dif1' cult to determine the pool acceleration from the JAERI level probe measurements because it involves e.americal evaluation of the second derivative of experimental data.

If there is analytic smoothing, before differentiation, the value of acceleration becomes too artificial because it is very dependent on the type of smoothing that is selected.

2-4

One can, nevertheless, draw a general conclusion about the adequacy of the NRC specification.

Because the velocity specification is bounding (compared to JAERI measurements), and the PSAM code predicts the pool swell on a realistic time scale, the accelerations produced by this code are conservative as well.

2.1.2.5 Submerged Boundary Pressure During Bubbla Expansion The NRC specification for loads on submerged boundaries, as stated in Reference 1 and modified for different pool elevations in Reference 2, stipulates the following:

(1) for the basemat:

uniform pressure equal to the maximum air-bubble pressure plus hydrostatic head corresponding to vent clearance from basemat (2) for the containment walls below vent exit:

maximum air-bubble pressure plus hydrostatic head corresponding to vertical distance from vent exit (3) for the containment walls between vent exit and maximum pool elevation:

linear variation between maximum air-bubble pressure and maximum wetwell-airspace pressure Because this specification is appropriate to the time when the pool is at its maximum height, the JA kI pressure measurements on the walls and the basemat are compared at this time.

The comparison between the NRC acceptance criteria and the observed submerged boundary pressure during bubble expansion shows higher calculated pressures at all pool elevations.

The pressure margins range from about 10 to 20% for the tests in Table 2.1-1.

2.1.2.6 Pool Elevation 2

The NRC specification for maximum pool swell height stipulates the greater of (1) or (2) as follows:

(1) 1.5 times vent submergence (2) the elevation corresponding to the time of drywell floor uplift AP = 2.5 psid.

The pool surface elevation corresponding to the maximum wetwell-airspace compression will be calculated assuming a polytropic process with an exponent of 1.2.

The application of (2) requires knowledge of the drywell-pressure history.

Normally, this would be calculated using the GE model.24 However, as discussed earlier, these calculations were not performed by the Mark II owners and were therefore not available for the specific JAERI tests.

Again the measured drywell pressures were used instead.

The uplift differential pressure that is added to the drywell pressure is 2.5 psi.

This dif' rential pressure corresponds to the minimum reverse pressure differential for a large break capable of producing significant pool swell.

This method was used to calculate maximum pool-swell heights.

The result was an exceedance of the acceptance criteria for maximum swell height (including froth activity).

The test results showed some fluid activity at the highest level probes (that is, approximately 25 feet above the pool surface).

2-5

-_. - _. ~ ~.

1 4

some fluid activity at the highest level probes (that is, approximately 25 feet above ttu pool surface).

j The question of interest is whether the activity is produced by the impact of a solid slug or whether the probes are merely wetted by spray.

This question 4

was resolved by looking at the timing of the probe activity.

If the level j

probe was actuated before the wet @l1 airspace reached a maximum pressure, the wetting was caused by solid impact.

If it occurred later, when the pressure 1-was decreasing and the slug "as a whole" was descending, the contact was made by froth or spray.

Froth impact is much less significant from the loading j

j standpoint.

Information related to timing of the probe activity was used to determine the maximum elevation reached by the pool in a bulk mode.

A com-i parison of the maximum bulk-mode elevation of the pool with the above acceptance l

criteria still indicated 0.5 - 105 exceedance of the maximum pool-swell elevation l

criteria.

It should be recognized that this is a very conservative comparison, because the drywell pressures used.for the NRC criteria were the measured j

JAERI pressures rather than the (higher) values one would obtain with the GE j

analytical model.

Consequently additional calcult ions were performed to

}

account approximately for this difference.

f' Although the GE analytical model is not available to generate the drywell pressure j

histories under the JAERI conditions, some pressure histories under somewhat different conditions are available.

In Reference 15, some CONTEMPT-LT/022 l

calculations for drywell-pressure histories for the reference plant are provided.

The CONTEMPT code was exercised in a mode that simulates the assumptions of the GE analytical model.

These calculated drywell pressure histories were used to construct the JAERI pressure history that would correspond to the GE

]

analytical model.

}

This information was then used to recalculate the maximum bulk pool-swell l

height. The resulting calculations using the acceptance criteria yield results j

that bound the maximum pool elevation in the bulk mode.

However, observed j

froth activity exceeded the criteria for maximum pool elevation in each of the i

tests in Table 2.1-1 by about 5 feet.

i In summary, with regard to the maximum bulk pool-swell height, the BNL calcula-i j

tions of the JAERI tests confirm the conservative nature of the current l

acceptance criteria for the maximum pcol elevation in the' bulk mode. The exceedence of the criteria in the.95RI tests as a result of froth activity is j

discussed below.

As mentioned above, froth activity was noted at an elevation of about 5 feet above the point of maximum bulk pool-swell height in each of the JAERI pool-swell tests in Table 2.1-1.

This is in contrast to the small (that is, about I foot) froth zone in the 4T tests.

The results of small-scale tests were reviewed to examine the Mark II owners' belief that froth activity in JAERI resulted from a downcomer bracing system l

that was located 2.5 feet above the initial pool surface.

l l

EPRI conducted 1/13-scale single-vent air tests with a typical Mark II geometry.

{

The influence of different bracing configurations on the gross pool-swell phenomena was studied.

Two different bracing cross sections (that is, square and i

1 2-6 i

1

- - - - ~. -

.n

- -. -. ~... -.-..-. - -.... - - - -

round cross sections) located at each of three different containment wetwell elevations (vent exit plane, initial pool surface, and three vent diameters above the pool surface) were ustJ in the tests.

Additional tests were per-Formed with braces located at a distance approximately equal to 3/4 of a vent diameter above the pool surface.

The results of these tests showed that there was no significant change in the gross pool-swell phenomenon for those tests conducted with the bracing system at either the vent exit or near the initial pool-surface elevation.

However, for those tests with the bracing located at an elevation three vent diameters above the initial pool surface, changes in the gross pool-swell phenomena were observed. This elevation corresponded to approximately the elevation of maximum pool-swell velocity.

Impact of the pool surface on the elevated bracing system resulted in significant froth activity above the maximum bulk pool-swell elevation.

Based on the EPRI tests, the staff has concluded that the froth activity in the JAERI tests probably resulted from the elevated location of the bracing system.

All domestic Mark I'I facilities with bracing utilize a system that is either submerged or 1cca;ed at an elevation near the initial pool surface or substantially above the region of maximum pool-swell height.

As a result, the staff does not foresee the potential for significant froth loads in domestic Mark II plants at elevations above the maximum bulk pool-swell elevation, nor does it foreste a need for the development of a froth load.

However, all Mark II owners shall conduct a review of their plants to establish that there are no significant obstructions, such as bracing in the wetwell region of maximum pool-swell velocity, that could ?ead to significant froth activity.

In addition, fragile equipment located in the potential " froth zone" (a zone extending to about 5 feet above the elevation of maximum bulk pool swell) should either be removed or protected from potential froth activity.

2.1.2.7 01aphrage-Floor Reverse Pressure The h C specification for diaphragm pressure differential, APUP is given as:

3 APUP = 8.2 - 44 F (psi) 0 < F < 0.13 APUP = 2.5 (psi)

F > 0.13 F = AB AP VS VD - (AV):

wher.r:

AB = break area AP = net pool area AV = total vent area VS = initial wetwell airspace volume VD = drywell volume The basis for this APUP specification is a correlation based on the GE 4T measurements.te.tr The details on how this correlation was derived from the data are described in Lehner's report (see above).

This specification was applied to the JAERI test facility after the JAERI test results were adjusted to the nominal Mark II Grywell and wetwell pool-operating 2-7

temperature conditions using influence coefficients (Lehner).

Under these 1

conditions the acceptance criteria bound the JAERI measurements.

Although it was determined that the acceptance criteria bound the JAERI measure-ments, a deficiency was noted in the NRC acceptance criteria for test conditions not included in the JAERI test matrix.

Unlike other pool-swell loads, the maximum reverse pressure does not neesr.sarily occur with the largest break area.

This becomes evidstt when it is reemenized that the reverse pressure increases as F decreases, and that, for any particular plant, F is directly proportional to the break area.

An invest.igation of 4T test data (Lehner) clearly shows this increase in reverse pressure as F is reduced in the range of F trom about 0.06 to 0.3.

As F is reduced further, the reverse pressure decreases. At F = 0, there is no upward reverse pressure on the diaphragm; rather a downward pressure of approximately 5 psi enrresponding to an approximate 12-foot vent rubmergence would result.

Because a spectrura of break areas ranginC from zero to the double-ended break is possible, the problem becomes one of determining just where the reverse pressure maximizes, and using the reverse pressure as the basis for the load specification.

As a result of the above deficiencies in the diaphragm-floor reverse pressure specification, the Mark II owners have revised this load specification.as They reviewed the results of the recent 4TCO tests, as well as the original 4T tests, for a range of simulated break sizes.

They observed a maximum upload of 4.14 psid after adjustments for initial drywell and suppression pool temperatures were applied.

In addition, they added a 1.36-psid. margin to cover uncertainties.

They arrived at a single 5.5 psid reverse pressure load specifi-cation for all Mark II plants covering the total range of potential break sizes.

NRC staff consultants have conducted a similar evaluation of test data (Lehner) from a variety of related test facilities.

The study shows that the proposed load specification is conservative.

In summary, based on the above information, the staf f has concluded that the 5.5 psid reverse pressure' load specification for the diaphragm is conservative and acceptable.

2.1.2.8 Wetwell-Afrspace Pressure The NRC acceptance criteriat stipulate determinatica of maximum wetwell-airspace pressure as the sum of the drywell pressure calculated with the GE code,14 and the maximum diaphragm-floor reverse pressure differential specified by NRC.

The previous specification for the maximum diaphragm-floor reverse pressure differential was a minimum value of 2.5 psid.

This value, rather than the revised value of 5.5 psid (see discussion in Section 2.1.2.7 of this report),

was used in the criteria for the maximum wetwell airspace pressure.

Because the GE drywell pressurization model was not available, the staff again used the JAERI measurements for the drywell histories. Adding 2.5 psi, the reverse pressure differential to the peak drywell pressure from JAERI measurements, the staff obtained the value for the NRC specification of the wetwell-airspace pressure.

This specification for wetwell-airspace pressure bounds the JAERI measurements with a margir. for the tests in Table 2.1-1 ranging from 13 to 20%.

2-8

2.1.3 Evaluation Summary The Mark II owners plan to use substantially the same pool-swell loads for all domestic Mark II plants as were used by the original lead Mark II plantste2, BNL, NRC consultants, evaluated the load plant pool-swell loads by comparing them against the JAERI full-scale test data to confirm these loads.

As a result of this study, the staff concluded that, with one exception, the tests confirm that the pool-swell acceptance crite.-ia are conservative.

The one exception relates to the maximum pool-swell height.

The maximum pool-swell l

height in the important " bulk mode" does not exceed the acceptance criteria.

However, there is evidence of froth activity several feet above the maximum l

criteria.

The NRC does not consider this froth load to be significant.

Nevertheless, it is the staff position that the Mark II owners should review their containments for equipment in the wetwell-froth zone to ensure that potential froth activity can be accommodated.

Preferably, equipment in this zone should either be protected or removed.

Although not directly from the evaluation of the pool-swell acceptance criteria against the JAERI test data, results of a review (Lehner) of the 4T test showed that the lead plant acceptance criterion for the diaphragm-floor reverse pressure should be revised.

The criterion was revised by the Mark II owners from a previous minimum value of 2.5 psid, which varied as a function of break t

size, to a new constant value of 5.5 psid for all plants, independent of break size.

The staff finds this new value conservative and acceptable.

2.2 Steam Condensation and Chugging Pool Boundary Loads After the initial pool-swell transient resulting from a postulated LOCA in a BWR, steam with decreasing amounts of air is vented from the drywell into the wetwell of the pressure-suppression system.

The purpose of this venting is to condense the steam in the wetwell pool and limit the pressure buildup in the containment.

During such steam venting, condensation-driven oscillations have been observed in related experiments.

Two types of condensation-driven oscillations occur.

The first type, called condensation oscillation (CO), occurs during the earlier portion of the blowdown.

C0 is characterized by approximately sinusoidal pressure oscillations in the l

drywell and wetweil system.

These condensation oscillations are followed by the second type nf condensation-driven oscillations, called chugging.

Chugging is characterized by discrete bursts of pressure oscillations in the wetwell pool with quiescent periods between them.

The pressure oscillations during chugging are associated with the rapid collapse of the steam bubble at the vent exit and typically exhibit a pressure spike, followed by a damped ringout which has predominant frequency components at the vent and pool natural frequencies.

Both of the condensation-driven oscillations produce oscillatory pressure loads on the containment surfaces.

l The condensation phenomenon involves an unsteady, turbulent, two phase flow.

The complexity of the phenomenon does not allow a completely analytical descrip-tion of the various chugging phenomena, Furthermore, because of the apparently random element in the condensation phenomena, there are significant questions whether condensation loads can be derived from scale tests.

Consequently, load definition relies primarily on a data base taken from full-scale experiments 2-9

which cover the range of conditions in Mark II plants.

For this reason, the Mark II condensation-driven CD and chugging loads are based on the results of tests conducted in the 4TCO test facility.

This facility is a full-scale single-vent test facility typical of domestic Mark II plants.

The 4TCO facility arrangement is depicted in Figure 2.2-1.

Twenty-eight tests were conducted with parametric variations as shown in Table 2.2-1.

Data to establish multivent effects were obtained from JAERI full scale tests.

The wetwell in the JAERI facility is a 20" annular sector with seven vents, typical of Mark II plants.

An elevation view of the JAERI facility is shown in Figure 2.2-2, and a plan view of'the vent arrangement is shown in Figure 2.2-3.

The test conditions are listed in Table 2.2-2.

2.2.1 Condensation-Oscillation Load 2.2.1.1

Background

The condensation-oscillation load definition, as described in the " Mark II Containment Dynamic Forcing Function Information Report (DFFR)," Revision 3,18 was based on the 4T Test Series 5101.

These early tests were designed and conducted with an emphasis on pool-swell phenomena investigations.

A later 4T test series was conducted to investigate the C0 phenomena.

These tests are referred to as the 4TCO tests. The 4T facility was modified for these tests.

The major differences in these tests from the earlier tests are the prototypical vent length and the emphasis upon recirculation line (liquid) break conditions.

On the basis of the results of a new test series, TS5200, documented as NEDE-24811-P,20 it was determined that Cie DFFR load definition 18 was not bounding ct all frequencies.

In addition, the amplitude of the DFFR load was exce,eded.

Consequently, the Mark II owners developed an interim load specification to reflect these observations. This interim load was developed to meet the licensing needs af the lead Mark II plant *.

The staff and its consultants reviewed the interim load and found it acceptable.

The final generic CO load definitionat is based on the same methodology as the interim load specification.

2.2.1.2 CO Load Description The CO load specification consists of two load cases:

the first is the " basic C0 load," and the second is the "20 load for combination with ADS."

The latter load specification is related to the conditions prevailing at a Mark II plant when actuation of the automatic depressurization system (ADS) takes place. A minimum ADS time delay of 90 seconds folloring a LOCA results in relatively low vent-mass-flux conditions.

The 4TCO test data shows that CO loads are small during low vent-mass-flux conditions; hence a reduced CO load is justified for CO loads in combination with ADS.

For both CO load cases, the specified loads in the drywell and in the suppression pool are to be applied simultaneously to a structural model of the drywell and suppression pool of a Mark II plant.

The basis for the load specification for

• The lead Mark II plants are LaSalle, Shoreham, and Zimmer.

2-10

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P Table 2.2-1 4TCO Test Matrix I t,1 t.

Init.

Nominal Wet-Dry-Vent Wet-well Init. well Sub-well Pool Free-Blow-Venturi Vessel Metal Mer-Initial Temp, space Run down Dia Press.

Temp gence Target Actual Temp Vent No.

Date Type (in)

(psia)

(*F)

(ft)

(*F)

(*F)

(*F)

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1-7-80 5

3.00 1046 275 11 70 69 60 No 2

10-15-79 L 3.00 1041 280 11 70 76 83 No 3

10-19-79 L 3.82 1045 270 11 70 74 72 No 4

10-24-79 L 3.82 1047 278 11 70 74 73 Yes 5

11-27-79 L 3.00 1044 278 11 80 79 60 No 6

11-3-79 L

3.82 1045 275 11 70 71 80 No 7

11-28-79 L 3.00 1042 280 11 90 93 70 No 8

11-5-79 L

3.82 1054 275 11 110 111 79 No 9

11-7-79 L

3.00 1045 277 11 110 114 79 No 10 11-29-79 L 3.00 1047 277 9

70 73 68 No 11 11-30-79 L 3.00 1047 278 13.5 70 74 79 No 12 11-9-79 L

2.50 1049 274 11 110 109 77 No 13 11-12-79 L 2.125 1052 271 11 110 109 72 No 14 11-19-79 L 2.125 1045 278 11 70 70 60 No 15 11-16-79 L 2.125 1046 280 11 70 71 69 Yes 16 10-4-79 5

3:00 1050 298 11 70 75 76 Yes 17 12-5-79 5

3.00 1051 270 9

70 71 63 Yes 18 12-10-79 S 3.00 1048 273 13.5 /0 71 65 Yes 19 12-21-79 5 3.00 1050 270 13.5 70 71 66 Yes 20 12-27-79 S 2.50 1045 270 11 70 68 59 Yes 21*

1-4-80 S

2.50 1047 275 11 70 68 61 Yes 72 11-21-79 L 3.00 1051 267 11 110 109 72 No 23 1-10-80 L

3.82 1045 278 11 110 100 69 No 24 1-29-80 L

3.82 1047 265 11 110 111 67 Yes 25 1-31-80 L

2.50 1046 271 11 110 111 69 Yes 26 2-2-80 L

3.00 1045 273 11 110 111 80 Yes 27 2-14-80 L

3.00

.045 270 9

110 110 71 Yes 28 2-15-80 L

3.00 1047 270 11 110 110 78 Yes "Jat deflector removed 2-12

TABLE 2.2-2 JAERI TEST MATRIX 0I50448G1 Cosettion5 luf fl AL CON 0lll0N5 lest

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fil JW1_l'sLJL _ril J=L 0001 meter 100 2.805 none ca.600 22.5

3. N2 0002 meter IOP 2.105 mene 7015 186 6.18 24.6 3.861 0003 watee 100 2.105 anne 4916 286

?.35 27.8 3.807 0004 meter 200 f.105 mene 7005 786 3.39 9.9 3.8523 water too F.305 none 77 1020 786 7.17 29.9 ab" i{

r105 meter 74 P.109 none M3 6966 287 7.99 14.2

3. N 5 3801 meter 74 2.105 23 F98 688F 784 7.79 1%.9

3. M ?

3102 water 200 2.105 98 68 69 %

F87 573

33. f 3.622 1701 stese 700 9.105 mene 85 6894 746 SJI 52.3 3.327

=

170ll stese 240 9.105 none 89 6976 286 5.44 53.5

3. M2 1703 s teau F20 9.105 none 51.9 6914 2M 5.06 53.9 3.340 1204 stese

??0

%.105 none 89.9 6978 286 5.32 18.7 3.338 1705 steam 720 9.105 none 94.5 4966 287 5.77 53

• 3.350 the drywell is the measured 4TCO drywell pressure, and that for the wetwell is the measured 4TCO bottom-center pressure (BCP).

The pressure for the drywell as measured in 4TCO is to be applied uniformly through the plant drywell.

The load configuration for the wetwell is shown in Figure 2.2-4; the load is to be applied uniformly on all Mark II plant-suppression pool surfaces below the vent exit.

The dynamic pressure amplitude is to be decreased to zero at the poo' surface (that is, the multiplier on the BCP values varies linearly from one at the vent exit elevation to zero at the pool surface).

The pressure-time histories used in the generic CO load definition are selected from the 4TCO Test Series 52002o and are given in Tables 2.2-3 and 2.2-4.

The parameters which are important to C0 loads, break size, pool temperature, and break type, (that is, steam or liqub break), were varied over the range of anticipated condition in the 4TCO tt.t series.

The selection rule for these time series was to bound the maximum power spectral density (PSD) observed in the BCP throughout the C0 period in all 28 4TCO test runs, in approximately 2-second blocks, up through 60 Hz.

If so desired, with the approval of the NRC, a specific Mark II plant can exclude from this data base of 28 runs those portions of the runs where the pool temperature exceeded the bounding temperature applicable for this plant.

For LaSalle, where the bounding temperature is 140*F, this exclusion resulted in a slightly different seTection, as expected.

The pressure-time histories for each of the time intervals listed in Tables 2.2-3 and 2.2-4 are applied separately to strucDaral models of the specific Mark II plant to obtain the structural response.

The application of these 2-13

time histories fully retains the dynamic character of the pressures as measured, including the phase relationship between frequencies.

These dynamic pressure signals are applied as rigid wall losds.

Table 2.2-3 4TCO Tims Periods for Basic CO Load Run Number Time (sec)

Run Number Time (sec) j 3

13 to 15 15 31 to 48 4

10 to 12

'22 13 to 21 5

19 to 21 23 5 to 7 8

5 to 7 24 12 to 14 9

10 to 23 25 32 to 42 10 28 to 30 26 16 to 24, 12' 21 to 25 32 to 36 14 25 to 31 27 16 to 24 Table 2.2-4 4TCO Time Periods for CO Load with ADS Run Number Time (sec.)

13 50 to 59 14 50 to 59 I

2.2.1.3 CO Load Evaluation The methodology presented in Reference 21 and summarized in Section 2.2.1.2 above haF been reviewed by the staff and its consultants.

The resulting evaluation is provided below.

2.2.1.3.1 Data Base The data used by the owners to arrive at the new load specification for CO are documented in the Final Test Report, NEDE-24811-P,2o referred to as Test Series 5200.

This test series was designed explicitly for a parametric investigation of CO.

The choice of the initial pool temperature and the break size was made

]

in order to cover the complete range of pool temperatures and steam-mass flow j

rates expected in Mark Il plant LOCAs.

The combination of initial temperatures in the range of 70*F-100*F with the 2.1-inch, 2.5-inch, 3.0-inch, and 3.8-inch diameter venturi sizes did cover the complete range of mass flux expected.

The number of test runs conducted in accordance with the test matrix provided sufficiently dense experimental data points to construct the contour maps presented in Reference 20.

The staff and its consultants have reviewed the final test report. The staff concludes that the accumulated data of the TSS200 tests cover the range of conditions encountered in all domestic piants with Mark II containments.

The staff finds that those tests provide sn appropriate data base for a conservative load specification for CO.

2-14

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I 2.2.1.3.2 4TCO Pressure Distribution The selection rule for the 4TCO pressure-time histories was as follows: The PSD* at each of the 2-second intervals for almost all the tests were superimposed to generate a bounding envelope over the frequency spectrum from 0 Hz to 60 Hz.

Every PSD that contributed to the bounding envelope was identified, and the time interval for which it was generated was determined.

The results are conservative because each pressure-time history is applied independently and directly to the structure.

It should be noted that the envelope does not represent a PSD of a time series of a hypothetical LOCA; rather, for any LOCA which is determined by some initial conditions, the PSD of its pressure-history will be bounded by this envelope.

2.2.1.3.3 Amplified-Response-Spectra Analysis In Appendix A of Reference 21, a pressure-response spectrum is presented comparing the response spectrum of all the 28 4TCO runs and that of the group selected using the PSD envelope as a selection criterion.

A response spectrum selected using an acceleration-time history as an input is normally used for evaluating a system tesponse for an acceleretion acting at the system support.

Similarly, when the excitation is a pressure-time variation, a pressure-response spectrum is of ten used with the pressure signal replacing the acceleration-time history.

In the camparison, the Mark II owners demonstrate that the selection rule which is based on PSD considerations results 4: a data set which essentially bounds the response spectrum of the 28 runs.

2.2.1.3.4 Comparison of JAERI and 4TCO The Mark 11 cwners have investigated CO loads observed in the JAERI test facility.22 This investicition consisted of a comparison of the CO loads in the JAERI and the 4TCO tests.21 The comparison was made for test runs in the two facilities having comparable mass flow.

The comparison shows that the 4TCO PSDs are significantly higher than the PSDs for the comparable JAERI test at all frequencies.

4 It may be noted that although the comparison is made between runs of comparable vent mass flows, these runs in themselves are bounded by the PSDs corresponding to the pressure-time histories of Table 2.2-3.

Thus, additional conservatism is available with the present load definition when comparea to JAERI.

• Power Spectral Density, PSD, is a representation of the energy at various fre-quencies in a signal. A PSD may be thought of as a decomposition of a time history signal into an infinite nurber of individual sinusoidal components of varying amplitude and phases.

In practice, the PSD consists of a finite number of evenly spaced frequencies representing a finite number of sinusoids (the number and spacing of the frequencies is dependent on the data digitization rate and the duration of the particular history).

A PSD value is calculated at each evaluation frequency.

The energy at each of the evaluation frequencies can be determined by examining the PSD.

2-18

i 2.2.1.3.5 Exclusion and Credits The 4TCO test facility geometry was designed te produce a bounding response te the blowdowns which simulate the LOCA to make certain that data from the 4TCO i

tests are represtative of all Mark II plants.

Similarly, the range of test parameters in the tests bounds the range for all plants.

Because specific Mark II plants have a much larger pool area per vent than the 4TCO facility, the loads in these plants would be expected to be smaller than those observed in the tests.

Also, the range of pool temperatures tested exceeds the range of pool temperatures expected in specific plants.

The base generic CO load does not take credit for these conservctisms.

However, specific plants may elect to take credit for these conservatisms using the methods previously used 8

by the LaSalle plant.

2.2.1.3.6 Load Conservatisms No credit is taken in the C0 methodology for either multivent effects or fluid-structure interaction (FSI) effects.

The conservatism implied by not taking credit for these two items is discussed below.

Multivent Effects There are two distinct aspects of the multivent effects:

one concerns amplitude, a spatial effect; the other concerns phasing, a timing effect.

(1) Amplitude:

If all vents are oscillating in phase, then the amplitudes of these pressure oscillations at the various vents may still differ in amplitude.

(2) Phasing:

If all the oscillatt ns at the various vents are of equal magnitude, then the timing may still be such that the oscillations are out of phase.

When dealing with a data base derived from single-vent experiments, credit for the amplitude effect might be obtained by taking a time average of a sequence of condensation oscillations in lieu of the spatial distribution anticipated for a multivent facility.

Credit for phasing can be effected by reducing the single-vent experimental data by a factor reflecting the effective increasa in cell volume in the multivent facility with out-of-phase vents.

Preliminary observations of the JAERI full-scale Mark II tests indicate a multivent factor of approximately 3

two.

In the current methodology, the bounding time series is selected at each frequency, and the load is applied directly to the structural models.

No credit is taken for either amplitude averaging or phasing; hence the conservatism results.

Fluid Structure Interaction The analysis submitted by the Mark II owners as Appendix B of Reference 21 states:

2-19

a i

(a) In the low-frequency range (f s 30 Hz), measured pressure amplitudes are essentially unaltered by structural flexibility; and (b) In the high-frequency range (f > 30 Hz), measured pressure amplitudes--which by definition are obtained in a flexible wall facility--are higher and, therefore, constitute a bound for pressure amplitudes anticipated at those frequencies for a rigid-wall facility.

One potential FSI effect is a shift in the frequencies of the 4TCO pool acoustic modes induced by flexibility of the steel boundary.

The study in Reference 21 indicates that this effect is not significant for the 4TCO wall pressures.

A second potential FSI effect is the introduction of structural resonances.

The structural modes examined included (1) bouncing of the entire facility on its l

concrete / soil foundation, (2) vibration of the baseplate, and (3) vibration of the sidewall.

Calculations performed in support of the Reference 21 study show that the net effect of these modes is to produce more high frequency j

response in the pressure signals than would bc seen in a rigid facility.

i Therefore, it is conservative to use the measured pressures directly for plant l

evaluation.

Details of the 4TCO FSI evaluation are presented in Reference 21.

l The staff agrees that the above assessment is correct and, therefore, the application of the measured pressures directly to a rigid plant for evaluation will produce a conservative result.

2.2.1.4 Evaluation Summary i

The staff and its consultants have reviewed the CO loads proposed by the l

Mark II owners in Rcference 21.

As discussed in the previous sections, the I

staff finds that the CO load is conservative and acieptable for use in the i

evaluation of the Mark II plants. The 4TCO test dr.a referred to in Tables 2.2-3 and 2.2-4 constitute the generic CO load.

T.'is load was proposed for the Mark II plants as a bounding load.

The generic i.:m' does not take credit for the potential reduction in the CO load resulting frc n the larger size of the Mark II unit cell versus the 4TCO cell size, or does it take credit for the load reduction resulting from excluding high-temperature 4TCO data.

For credit to be taken for geometry or pool temperature limits by a specific Mark j

II plant, information similar to that provided in Reference 23 should be l

provided to the NRC.

l 2.2.2 Improvea Chugging Load j

2.2.2.1

Background

l If a LOCA-type of blowdown should occur in a Mark II containment, the bulk of the drywell air is first carried over into the wetwell, which esults in pool-swell phenomena.

Then, essentially pure steam is forced through the downcomers into the suppression pool where it condenses.

During the early stages of this steam blowdown phase, local steam-flux rates are high, and the steam-water interface near the vent exit is relatively stable.

As the steam flux continues to decrease, the steam-water interface takes on an oscillatory character.

Eventually, a random and erratic motion results from the growth 2-20

and rapid collapse of steam bubbles.

These collapse events, referred to'as chugs, vary in intensity and ducation and cause associated pressure loads on the containment.

The lead plant chugging load specification was based on a conservative pressure time-history obtained from the original 4T (Temporary Tall Test Tank) tests conducted by GE for the Mark II owners.

The lead plant specification involved the direct application,f these 4T pressure traces to the Mark II containment walls.

Details of the lead plant specification can be found in NUREG-04873 The Mark II owners felt that because of enhancement of the pressure signals in the 4T tests by fluid-structure interaction (FSI) and some other considerations, the direct application of 4T pressur e traces as specified in the Lead-Plant Program was overly conservative.

Therefore, the Lead-Plant Specification was regarded as a temporary one, and the owners initiated en improved chugging task which they believed would lead to a reduced chugg"ng specification for the Long-Term Program.

The evners at first used the stme 4T data that was the basis for the Lead-Plant Specification for this task.

llowever, when the owners decided to resolve questions concerning vent-length effects during the condensation-oscillation phase of steam condensation by conducting a new set of tests in the 4T facility, the NRC staff requested that they monitor chugging data as well.

The staff's concern was related to peak pressure amplitudes recorded in several foreign full-scale tests, which exceeded the peak amplitudes of the original 4T tests.

The new 4T test series (referred tc cs the 4TCO tests) included chugging events with peak pressures significantly greater than those observed in the original 4T tests.

The new test series had a vent geometry prototypical of Mark 11 plants in which the drywell could be charged by a two-phase flashing mixture for a liquid break simulation.

These tests demonstrated that certain combinations of mass flow, air content, and temperature during the blowdown could produce higher chugging loads than were produced under other conditions.24 Therefore, the NRC and the Mark II owners agreed that the generic improved-chugging-load task would use the 4TCO data for the development of a refined chugging load, and that the containment of all plants, lead plants included, would have to be assessed for this new load.

In July 1980, the Mark II lead-plant owners proposed using a simplified approach (which they believed was bounding) to deal with the new data from 4TCO for assessment of their containments.

This simplified interim approach is outlined and reviewed in Supplement 2 to NUREG-04878 This interim approach was not, however, a substitute for the load specification developed through the generic improved-chugging load task.

Accordingly, each of the lead Mark II plant owners stated that he will assess the plant containment with the loads examined here.

2.2.2.2 Chugging-Load Description In the generic improved-chugging-load methodology, the Mark II suppression poo,1 is treated as an acoustic medium with each chug represented as a point source located at the vent exit.

The 4T and Mark II containment pools are modeled acoustically, and a vent exit source is inferred from the test facility data. The resultant design source is then transferred to a Mark II containment.

The acoustic model is thoroughly described in Reference 25, but the test data used in this reference are still the original 4T data.

Reference 26 describes the final methodology and plant application with the design sources based on the 4TCO data.

2-21

2.2.2.2.1 Design-Source Description Ten acoustic sources, referred to as design sources 801 to 810 in Reference 26, are used to define the load. These sources were inferred from actual chugs occurring during the 4TCO tests.

Details of the " inferring" of sources from actual chugs are discussed in Reference 26.

Each source consists of a triangular impulse and a series of sine terms.

Sources 801 to 807 are used to generate responses from 0 to 100 Hz.

These seven sources are an average of the sources inferred from two adjacent chugs.

Sources 808 through 810 are inferred from single chugs only (the higher of the pair of chugs for 803, 804, and 805) and are used only to generate responses from 50 to 100 Hz.

Table 2-1 of Reference 26 gives the numerical parameters associated with the design sources.

These consist of impulse amplitude, duration, and shape; sinusoidal amplitudes and frequency; pertinent damping values; s:sociated pool acoustic speed; and the 4TCO run number during which the chug or chugs occurred from which the source was inferred.

2.2.2.2.2 Design-Source Application The design sources are inferred from the wall pressures recorded dur ing the 4TCO tests by using a computer program called IWEGS, which solves the inhomogeneous wave equation in a cylindrical geometry.as The next step is generating Mark Il containment wall pressures by applying the inferred design sources at the exits of all the vents in a particular Mark II containment.

These wall pressures can then be applied to an appropriate fluid-structural model of the containment to obtain the structural response's.

The 10 sources 801 to 810 are applied separately to the plant to obtain the structural response.

The envelope of the responses from all 10 sources is then used in the plant evaluation.

2.2.2.2.2.1 Transfer from Point Sources to Wall Pressures Reference 26 details two methods for determining the wall pressures from the point source loads:

one is a rigid-wall method; the other allows for flexible walls in a limited way.

In the rigid-wall procedure, wall-pressure responses to design sources are calculated as though the containment had perfectly rigid walls.

To do this, a computer code called IWEGS/ MARS is employed.

This code solves the inhomogeneous wave equation in an annular geometry.

The resulting rigid-wall pressures are then applied to a coupled model of the actual flexible containment structure and the enclosed fluid.

Structural responses are calculated from this coupled model.

In the rigid-wall procedure, the effect of 4TCO wall flexibility is removed by adjusting acoustic speed and damping values in the IWEGS code used to infer the design sources.

Pool acoustic speeds and damping values appropriate for rigid suppression pool boundaries are then used in the IWEGS/ MARS code to calculate the containment-wall pressures.

In the approximate flexible-wall lotd procedure described in Reference 26 the design source acoustic speed and damping values are adjusted to account for differences between 4TCO and Mark II wall flexibility and thus compensate directly for fluid-structure interaction.

The advantage of this method is that the model of the containment us1d for structural analysis need not be fluid coupled and is, therefore, simpler and less expensive.

2-22

2.2.2.2.2.2 Chug Desynchronization The design sources are applied one at a time. That is, for one pool chug, all the vents have source 801 applied at their ends; for the next pool chug all vents use 802, and so on, through 810.

However, during a pool chug, a given source is applied desynchronized; it does not start at all vents at exactly the same time.

The individual precise chug-start time for each vent is deter-mined as follows:

A uniform distribution of start times over a 50 ms interval is assumed.

One thousand Monte Carlo trials are conducted to select 1000 sets of chug-start times.

(For a 100-vent plant, one set of start times consists of 100 time values in the 50 ns interval.) That set of the 1000 having the smallest variance is used for the actual start time application of all 10 snurces.

The selected chug-start times are randomly assigned to the vents in the containment.

2.2.2.2.2.3 Symmetvic-Load Case Two loading cases have been identified by the Mark II owners: A symmetric-load case and an asymmetric-load case.

The symmetric-load case is simply defined as the desynchronized application of each of the 10 design sources in turn at every vent exit location in the Mark II wetwell.

Ten independent pool-boundary-pressure distributions corresponding to the 10 design sources will result from the IWEGS/ MARS calculations.

Each of the 10 pressure distributions will then be used for structural analysis of the containment.

2.2.2.2.2.4 Asymmetric Load Case The definition of the asymmetric load case is more complicated than the symmetric case.

First, a design moment axis is identified. This is done on purely geometrical grounds. Any diameter of the suppression pool can be considersd a moment axis.

If L is the perpendicular distance from the ith vent to such a 9

diameter, then, in general, N

L 0

i=1 because of the asymmetry in vent locations.

The design moment axis is the one that maximizes this summation.

Plus and minus sides of the axis are chosen to make the summation nonnegative.

A source strength, S+ = (1 + a)S, where S is the design source strength used for the symmetric load, is applied to all the vents on the plus side, and S- = (1 a)S is appiled to all the vents on the minus side.

The parameter a can be expressed in terms of the root mean-square which, in turn, can be defined using the expected value, variance, moment, M and desigb"Ia,lue of a statistical measure of the chug strength.

In other words, a can be chosen to correspond to desired p'robability values for not exceeding a certain res moment.

Details of the evaluation of a can be found in Reference 26.

The S+ and S-design sources are again applied with the same set of chug-start times as the symmetric case.

Each design source, modified l

to become S+ over half the pool and S-over the other half, is, in turn, applied at the vent exits.

This results again in 10 independent distributions of pool-boundary pressure corresponding to the 10 design sources.

2-23

2.2.2.3 Chugging-Load Evaluation 2.2.2.3.1 Data Base Although many experiments to evaluate and understand chugging have been conducted both domestically and internationally, the Mark II owners' specificat*on is based on data from two facilities:

the 4TCO tests conducted by GE for the owners and the JAERI tests conducted at Tokat-Mura, Japan by JAERI.s s>27-so 2s The single-vent, full-scale 4TCO tests provide the fundamentals for the specification of sources amplitudes, frequency content, damping and pool acoustic speed).

JAERI data (accumulated in a seven-vent, full-scale facility) is used for the desynchronization of chug-start times and for some of the statistical quantitites involved in the asymmetric load application procedures.

The 4TCO test series was designed for a parametric investigation of condensation oscillation.

The ranges of break size (as well as break type), mass flow, and pool temperature over which the tests were conducted covered the entire Mark II range.

Therefore, the parameters important to chugging--such as mass flow, air content, and bulk pool temperature, and their combinations--were also distributed over the range of interest for Mark II chugging.

Although the investigation of the C0 phenomena concentrated on relatively large steam and liquid breaks, substantial chugging data were accumulated from all blowdowns except those with the largest break size.

Past observations have indicated that the highest amplitude chugs occur when the vent-masb flow is relatively high anc 'he steam nearly air free.

Because these are the conditions toward which the 4TCO tests are biased, the staff concludes that the chugs observed in this test serios provice a conseiyatO.e data base.

Over 300 chugs observed during the course of 28 blowdowns provide an appropriate basis for load assessment.

The JAE.".I Mark II full-scale steam tests are containment pressure suppression tests sponsored by the Japanese government in support of Japanese plants with Mark II containments.

This 5 year test program is being conducted in a seven-vent, full-scale facility that is prototypical of a 20 sector of a Mark II containment.

This test facility represents the most realistic model of a multivent Mark II containment of the various tests condu-ted.

Data from these tests are the most appropriate for assessing multivent effects in the Mark II design.

Other test programs have also contributed to the understanding of the chugging process.

The most important of these include:

the single-vent full-scale, prototypical plant-unique tests condec+.ed in GKM-II-M for Pennsylvania Power i

and Light (PP&L)81 the three-vent, fall-scale but nonprototypical GKSS tests conducted in Germanysa; and the scalea, multivent tests conducted for the Mark II owners by Creare, Inc. of New Hampshire.88'84 In general, these experiments have helped to confirm the results of the 4TCO and JAERI tests and have provided added assurance that all significant aspects of the chugging phenomena were included in the 4TCO and JAERI observations.

The scaled multivent Creare tests also aided in verifying the validity of the acoustic modeling of the Mark II suppression pool.

i 2-24

2.2.2.3.2 Selection and Sourcing of Chugs The <TCO test series consisted of 28 blowdcens of either liquid or steam.

Venturi diameter, initial drywell air centent, submergence, and pool temperature all were varied to span the Mark II range, Detailed information on the test conditions'can De found in Table 2.2-1.

Data were reduced for 26 of the 28 runs, resulting in 297 chugs from which sources could be selected.

Six of the runs produced the 15 largest chugs based on mean square power (msp) and peak overpressure (POP).

Ccnsistent with observations made at other blowdown experiments, the largest chugs in 4TCO occurred in those runs where the vent steam mass flux remained high after the drywell air was pu ged.

These conditions occurred during chugging intervals between periods of CO or soon after the end of CO.

Chugs from which design sources would be developed were selected based on two requirements.

First, the largest chugs in terms of msp and POP would be included and second, the chugs had to be representative of the frequency content observed in all the 4TCO data.

Table 4-3 of Reference 26 lists the 15 largest chugs in order of msp and shows the 7 chugs which were selected as key chugs for design source purposes.

The seven key chugs selected include the three highest and six out of the top seven based on msp.

They also include the five highest with respect to POP.

All six runs with the highest chugs are represented among the key chugs.

The chugs chosen also cover all three break sizes, both types of blowdown (liquid or steam), all three initial pool temper-atures, and the inclusion as well as the exclusion of a vent riser.

A PSD (Power Spectral Density) comparisen between the seven key chugs and the 4TCO bounding PSD envelope shown in Figure 4-22 of Reference 26 confirms that the seven key chugs are very representative of the frequency content observed in all 4TCO data.

In order to infer design sources from specific 4TCO chugs, the IWEGS.nmputer code is used to generate a transfer function between point source strength at the vent exit and pressure at the botton center of the 4TCO baseplate.

A general source function is assumed which includes both the impulsive bubble collapse and the sinusoidal components in its form.

The sinusoidal components are primarily due to vent response.

Acoustic speed and damping are used in the trans'er function to account for the effects of air content and wall flexibility.

All the adjustable parameters in the source expression (amplitudes, frequencies, duration, and damping) are chosen so that the calculated bottcm center pressure response to the source agrees with the observed pressure at the same location in the 4TCO tests.

Details of the procedure can be found in Secticn 4.3 of Reference 26.

The staff agrees with the owners that the seven key chugs selected for design source derivation represent a conservative subset of chugs from the 4TCO data.

The staff has concluded that the way sources are inferred from the key chugs is conservative with respect to amplituda and power because the msp of the source parameters is made to mat or exceed the msp of the measured chug pressures for all cases except for the higher of the two sources used for design source 807, which achieved 95%.

2-25

t

'2.2.2.3.2.1 Design-Source Amplitudes 3

Observations from all multivent test. facilities show that although system conditions do affect chug strength from an overall statistical viewpoint, there is still a large variation in pressure amplitudes among vents during any one pool chug.

Therefore, the Mark II owners determined that applying the seven sources j

inferred from the key chugs, which represent the highest amplitudes obstrved,

)

at all vents, was overly conservative.

Any source applied at all vents fo.' a pool chug should have an amplitude more representative of the average during '

particular set of system conditions rather than the highest recorded.

Therefore, to account for the vent to vent amplitude variation within a multivent pool I

chug, each key chug source was averaged with the source obtained from its larger adjacent chug. To develop average sources for the key chug and its larger adjacent chug, sources for the individual chugs were generated and then averaged at eac5 Lignificant frequency.

The acoustic speed for the average i

source was also the average of the acoustic speeds used for the individual sources.

The significant frequencies were usually averaged at integer values j

and limited to 10 or fewer different frequency terms for each average design source.

Details of the averaging procedure and justification fer its mathematical correctness (due to the linearity of the acoustic equations) can be found in Section 4.4 of Reference 26.

In addition to the seven averaged sources, the key chug sources used in the I

generation of average sources 803, 804, and 805 are used full strength but only for containment response above 50 Hz.

These sources are referred to as sources 808, 809, and 810.

1 l

The staff agrees that *he a$aplitude variation from vent to vent observed

^

during a pool chug in a multivent facility is a real effect. The staff con-cludes that the owners' method of proceeding (averaging amplitudes with the largest chug adjacent to the key chug) is acceptable and provides a conservative

}

method of accounting for this amplitude variation in deriving the design sources.

In addition, the choice of sources 808, 809, and 810 enhances the high frequency response because the key chugs from which these sources are 4

4 inferred show some of the highest power at 50 Hz or above among all the 4T00 q

l chugs.

]

l 2.2.2.3.2.2 Frequency Content of Design Sources i

l Tuere are a maximum of 10 sinuseidal termi in each design source, with several l

sources having fewer terms.

The frequcnetes of the sinusoidal terms are generally below 30 Hz, so any higher frequency response must come from the triangular impulse terms. Averaging of key chugs with adjacent chugs to infer i

design sources actually enhances the frequency content of the sources because I

the average of the two chugs will contain the frequencies from both, although j

at reduced amplitude.

Therefore, the design source inferred from the combination of the key chug and an adjacent chug will, in general, be richer in frequency j

content than if it had been inferred from the key chug only.

The use of discrete frequencies for the sinu,oidal terms in the source definition may seem limiting because the observed chugs obviously have no such restriction.

l However, the triangular impulse provides response over a wide range. The best j

way to judge the frequency content of the source specification is by comparison of the power spectral density (PSD) of the source induced pressures with the i

I 2-26 L

i l

PSD from observed pressures at the same location.

Such comparisons are shown in Reference 26 for both individual sources and the corresponding key chugs, as well as for PSD envelopes of the design sources compared to PSD envelopes of all the 4TCO chugs. A PSD comparison between design sources and data from the JAERI facility is also shown in Reference 26.

Based on these comparisons, the staff concludes that the frequency content of the design scurces is adequate and acceptable for load specification.

2.2.2.3.3 Acoustic Modeling As stated in Section 2.2.2.2.2 above, the design sources are inferred from 4TCO boundary pressures and then applied at all the Mark II vents, to produce boundary pressures on the Mark II containment.

To accomplish both of these operations in a convenient and cost-effective manner, chugging is taken to be acoustic in nature.

This means that the conservation equations that describe chugging in a pool can be reduced to the acoustic wave equations with acoustic point sources at each vent l'ocation. The justification for this treatment is presented in considerable depth in Reference 25.

The development of an acoustic chugging medel is based on two assumptions, both of which are supported in Reference 25:

(1) The linear wave equation applies to chugging.

(2) The vents are not acoustically coupled to the suppression pool; that is vent natural frequencies are independent of pool pressure oscillation frequencies.

These as;umptions lead to a set of acoustic equations whose solution is obtained in a straight forward manner by the Green's function method as illustrated, for instance, by Horse and Ingardas. To calculate numerical results from these analytical solutions, two related computer codes, IWEGS and IWEGS/ MARS, are used.

The codes differ only in boundary conditions for which results are obtained.

IWEGS is used for the cylindrical geometry of the 4T facility and IWEGS/ MARS for the annular Mark II containment geometry.

The staff finds that the acoustic modeling of chugging as described in Reference 25 to infer sources and apply them in Mark II containments is accept-able. The staff's finding is based on several considerations.

First, the physical arguments presented in Reference 25 to justify the simplifying assump-tions needed are sound.

Second, as shown in Reference 25, the two-dimensional, two phase flow, nonlinear finite difference code K-FIX was used by the Mark II owners to assess the impact of the assumptions and show their acceptability.

K-FIX was also used to verify some of the solutions obtained from the simplified acoustic Grsen's function IWEGS approach.

Finally, the observed 4TCO boundary pressures can be matched on an individual chug basis closely with the acoustic-model boundary pressures.

This is true both in the time and frequency domain.

Acouatic modeling was also used successfully to predict boundary pressures observed in several of the scaled multivent Creare chugging experiments.

2.2.2.3.3.1 Fluid Structure Interaction Fluid-structure interaction (FSI) in the chugging context refers to the difference in the boundary-pressure field of the fluid which is obtained using 2-27

rigid boundaries and the pressure field when the boundaries are ilexible.

Because truly rigid boundaries are almost never realized in practice, FSI to some degree is clways present in the chugging process.

F1 depends on the boundary flexibility and flexibilities differ among Mark II containments and between Mark II containments and chugging test facilities, as well as among test facilities themselves.

Therefore, transferring chugging results obtained in the test facilities to Mark II containments for structural assessment requires the inclusion of the effects of FSI in the analysis.

Two different methods of obtaining wall pressures which include FSI effects are considered in Reference 25:

a rigid-wall method and a limited, flexible-wall method.

The rigid-wall method, so called because the acoustic model equations are solved with rigid walls as a boundary condition, relies on the separability of the FSI problem.

First the acoustic-fluid equations are solved for the case of rigid boundaries; that is, where the normal component of fluid velocity is equal to zero on the boundary.

The resulting rigid-wall pressures are then input to a set of coupled fluid-structure equations allowing for boundary displacement, and an additional pressure field generated by the motion of the boundary is obtained.

The total pressure is the sum of the rigid-wall pressure and the pressure due to wall motion.

In other words, the boundary pressure which would be obtained in a flexible container from a forcing signal applied within the fluid is equated to two components:

the boundary pressure found in a rigid container with a forcing signal applied within the fluid Pr, plus the boundary press,ure from a flexible container with Pr as the forcing si,gnal applied on the boundary.

One of the staff's consultants (Lehner) demonstrated that this equation is valid if some very reasonable assumptions hold:

(1) Boundary displacements are small.

(2) Density perturbations also remain small.

(3) The period of the boundary oscillations is short compared to the characteristic time of the blowdown process.

A further proof of the validity of this method and a demonstration of its application was given by the Mark II owners in Reference 25, Section 5.1, where the results of several NASTRAN* calculations are presented:

First a triangular impulse was applied at the vent exit in a NASTRAN model of the flexible-wall 4T facility and the PSD of the resulting bottom-center accelera-tion computed.

Next the same impulse was applied to a NASTRAN model of 4T with rigid walls, resulting in the rigid-wall pressure.

This rigid-wall pressure was used as the input to the flexible-wall NASTRAN model without the vent source, and the PSD of the bottom-center acceleration was again computed.

As shown in Reference 25, the acceleration PSDs obtained from both calculations are identical, again verifying the rigid-wall method.

In addition, the method was used successfully to infer sources from 4T and 4TCO experimental data and then to match the observations in both the time and frequency domain.

The staff finds the rigid-wall method acceptable for calculating FSI effects in Mark II containments.

2-28 m

The flexible-wall method treats fluid-structure interaction without consider-ing the structu a per se.

FSI effects are accounted for in the acoustic model by shifting of frequencies and changes *n the damping parameter.

When using the rigid-wall method to finally obtain the structural response, the designer must utilize a structural computer program with compressible fluid elements because both the containment and the pool must be modeled when the rigid-wall pressures are appifed.

Such proqrams are complex to implement and relatively expensive to execute when compa'.ud to strictly structural programs..Therefore, if a flexible wall solution of the acoustic equations could be obtained, it would be very desirable from the designer's point of view because it would represent a substsntial cost saving in both reduced analyst manpower as well as reduced computer cost.

R3ference 25 demonstrates that a solution to the acoustic wave equation with flexible boundaries can be obtained with little more difficulty than for rigid walls, provided that certain assumption hold.

The principal assur.otion involved is that the conteinment boundary walls are locally reacting.

This means that the various parts of the containment surface are not strongly coupled to each other and the motion of a particular portion depends only on the acoustic pressure incident on that portion of the boundary.

If the walls are assumed to be locally reacting, then a specific acoustic admittance can be assigned to each point of their surface.

The acoustic admittance is a function of fre-quency as well as location. The solution to the flexible-boundary wave equation can be found in *erms of the acoustic admittnnce provided that this admittance is sufficiently small to allow a perturbation solution and sufficiently uniform to allcw the evaluation of certain surface integrals which occur.

(For details see Section 5.2 of Reference 25.) In practical terms, Shese assumptions are reasonable for Mark II containments with the rigid boundaries typical at reinforced or prestressed concrete containment vessels.

To verify that the flexible-wall method is a reasonable engineering approach for the Mark II chugging problem, the owners show in Reference 25 that flexible-wall eigen frequencies calculated with this method agree with calculations done by other means as well as with experimental results.

The fundamental flexible wall eigen frequency of the 4T facility is computed by the above methoo as well as by a structural approach given in Apper B of Reference 25 and by a NASTRAN calculation.

The computations all agree within a few percent.

Results are also compared to the experimentally obtained frequencies from the ANAMET 4T bell jar and impact tests.

Again, the agreement is good.

Next, the flexible-wall eigen frequency of a Mark II suppression pool is calculated by the acoustic Flexible-wall method and compared to a NASTRAN calculation.

Results are within 3%.

Finally pressure-time histories and PSDs from a chug source in the 4T facility are obtained with the flexible-wall acoustic theory and compared with NASTRAN calculations.

The results, illustrated in Figure 5-9 of Reference 25, show good agreement.

The staff concluded that flexible-wall method is valid for use on concrete Mar'k II containment vessels because the assumptions needed for the method are f

• NASTRAN is a nationally available computer program which allows the direct cair'ilation of flexible-wall pressures including CSI in a single step.

2-29

Justified for those staff structures.

Additional work will be needed to support the use of the flexible-wall method for steel containment vessels. The verification calculations provided in Reference 25 to justify the method demonstrate the applicability of the flexible-wall method.

However, the acoustic speed and damping, whose values influence the numerical outcome of the calculations, while reasonably chosen, are assumed values because specific data do not exist.

These two parameters are the only ones which show the FSI influence in the acoustic flexible-wall method.

For Mark II containments the damping value is fixed by a regulatory guide.

Therefore, the staff finds that when the flexible-wall method is applied for Mark II containment assessment, calculations should be made over a reasonable range of acoustic speeds because any particular value of sound speed in an actual containment is based on a number of assumptions. The range of acoustic speeds implied by the 10 chugging design sources (801-810 in Reference 26) should be used.

This range of acoustic speeds reflects those anticipated in Mark Il plants because the sources reflect the 4TCO parametric variations anticipated in Mark II plants.

In summary, the staf f finds the flexible-wall method acceptable for concrete Mark II containment vessels provided that the calculations are performed over the range of acoustic velocities indicated by the 10 chugging sources 801-810.2s 2.2.2.3.3.2 Sonic Speed and Damping Table 4.6 of Reference 26 lists two damping values and an acoustic speed for each design so9rce.

The sinusoid or vei.s damping is chosen as 6.5 or 6% for all sources, waile pool and wall damping ranges from 8 to 124.

Acoustic speeds vary from a low of 490 m/sec for source 801 to 1276 m/sec for source 805.

Reference 26 states that the acoustic speeds and damping used in Table 4.6 were adjusted to remove the effect of 4TCO wall flexibility. The chosen values were selected to reflect the air content in the pool.

The staff has concluded that there are too many parameters which can affect the acoustic speed and ine damping to infer values for specific chugs with great accuracy.

However, the staff finds that the damping values chosen for the design sources are conservative and the range of acoustic speeds applied i

with the 10 sources is wide enough to cover Mark II conditions.

2.2.2.3.4 Desynchronization of Chug-Start Times As described in Section 2.2.2.2.2.2, during a pool chug, a particular design source is not applied at exactly the same time at all vents but rather in a desynchronized manner; that is, the start times are randomly selected from26 within a 50-ms window.

The desynchronization of start times is borne out by 22 If some expanded time traces of the vent-exit pressures in the JAERI tests characteristics of a chug, such as the underpressure preceding the first significant overpressure, are chosen as a means of pinpointing chug starts, the five instrumented JAERI vents always show the characteristic feature occurring at slightly different times for different vents.

In order to justify their choice of a 50-ms time window, the Mark II owners calculated the average chug-time variance of 25 pool chugs from five JAERI tests for which the chug time could be unambiguously determined.

The result-Ing average sample variance corresponded to a 61-ms uniform window. Various statistical tests were conducced to establish that desynchronization does not 2-30

depend on plant conditions.

Details can be found in Section 5 of Reference 26.

To be conservative, the owners chose not the 61-as window but a smaller 50 ms one.ss The rest of the desynchronization application then proceeds as outlined in 2.2.2.2.2.2.

The staff finds that the desynchronization methodology as presented by the owners in Reference 26 is acceptable.

The staff agrees that scaled and full-scale multivent data show desynchronization of chug times and that the JAERI data base is the most appropriate one to use for quantification of this desyn-chronization.

Further, the method of arriving at the 50-ms value presented by the owners is acceptable.

The staff has concluded that using a chug-start-time window derived from data of five instrumented vents for a plant of 100 or so vents represents a substantial conservatism.

In other words, confinir.g the chug-start times from 100 vents to lie within a 50-ms time window which was observed from data of only five vents is conservative.

An additional con-servatism is the use of 1000, Monte Carlo trials to select the set of start times with the smallest variance for plant application.

In light of these various conservatisms, the staff finds acceptable the methodology for vent desynchronization as detailed in Reference 26.

2.2.2.3.5 Comparisons with JAERI Data In Section 6 of Reference 26, the Mark II F.ners compare the pressure response in the JAERI facility calculated with the 10 design sources with directly observed JAERI chugging data.

The design-source pressure predictions are made using a computer code called JMARS, which is based on the same governing equa-tions and uses the same solution technique as IWEGS and IWEGS/ MARS but is tailored to the JAERI geometry.

Thus, the acoustic methodology is also tested in this verification.

Comparison of the design sources is made with data from the eight largest chugs observed in JAERI Test 0002, which include the largest chugs measured under prototypical blowdown conditions in JAERI.

For each of the eight chugs, the spatially averaged pressure history is computed at two elevations (corresponding to where the most accurate pressurs data was obtained).

JAERI computed the spatially averaged pressure histories, the PSDs of the average pressure histories for each of the eight chugs at two elevations 1

(corresponding to where the cost accurate pressure data were obtained), and constructed the maximum envelope over the eight chugs.22 The owners compared the maximum envelope with the design source predicted pressure response.

The design-source PSD envelope is constructed by applying a given design source from the set of 10 design sources referred to in Section 2.2.2.2.1 to each of the JAERI vents, with start times selected from the 50-ms uniform probability distribution.

PSDs of the pressure-history average response at the same two elevations where the data were obtained are computed.

This calculation is repeated eight times with different sets of chug start times to obtain a PSD envelope.

The eight-chug simulation is repeated 20 times to get an average PSD envelope for one source.

The above calculation is repeated for each design source and the average PSD envelopes are again enveloped to obtain the envelope to be compared with the data PSD envelope discussed in the previous paragraph.

The calculatea responses to the desynchronized design sources bound the multi-vent data with substantial margin for all frequencies below 50 Hz.

In the i

high-frequency range (50-100 Hz), where the signal power is quite low, the calculated response, although not bounding at every frequency, is comparable to the multivent data..

2-31

While some arguments could be entered into whether the above method of comparison is the best one, the staff feels it is an acceptable one and is important in confirming the conservative nature of the Mark II owners' generic improved chugging specification.

2.2.2.3.6 Plant Application 2.2.2.3.6.1 Symmetric-Load Case The symmetric-load case should provide a conservative representation of the net vertical force on tne basemat and the net pressure acting on the contain-ment wall during a pool chug.

The staff finds that the symmetric load specific-ation proposed by the Mark II owners does fulfill this requirement.

Both the data selection for and development of the design sources are performed in a conservative manner.

The applicetion of the sources to a Mark II containment is also done conservatively with the selection of a 50-ms time window for chug-start times and the 1000 Monte Carlo trials for selecting the start time set with the minimum variance.

The Mark II owners performed sensitivity studies to confirm the trend of decreasing boundary loads with increased start-time variance.

They also verified the insensitivity of pool boundary pressures to vent assignment patterns.2e Both results demonstrate the conservative nature of the symmetric load specification.

2.2.2.3.6.2 Asymmetric-Load Case The asymmetric-load specification should provide a conservative upper limit to the possible imbalance in the pressure-load excitation of the containment structure due to variation in chug strength among vents during a pool chug.

l In Reference 25, the Mark II owners support their claim that the specification as outlined in Section 2.2.2.2.2.4 of this report fulfills this requirement.

In Reference 25 justification is given why the rms moment is an appropriate quantity for the asymmetric load bases.

There are also some interesting sensitivity studies reported in Reference 25 on the effect that random non-synchronization has on the asymmetric effect at small probabilities of not exceeding a certain lateral force.

However, the final asymmetric-load speci-fication as outlined in Reference 26 differs from that of Reference 25 in an important aspect:

The Reference 26 case applied a chug-strength variance under the assumption that chug strengths observed at any time during a blow-down could occur simultaneously at all vents.

The present asymmetric-load case utilized instead the variation in individual chug strengths within pool chugs from the JAERI data base.

This is a correct assumption in the staff's view.

In assessing the asymmetric-load specification, it is very difficult to accurately predict a realistic maximum asymmetry which a containment may experience.

There are many variables with counter balancing effects which tend to make analytic predictions very difficult.

On the other hand, it is unlikely that there are large asymmetries in a process which has so many variables and such a high degree of randomness as chugging.

The staff concludes that while the approach taken by the owners may not be the best one from strictly statistical constaerations, it is adequate in that it provides a reasonable measure of asymmetry. The application of the 10 sources in the asymmetric specification provides an additional 10 boundary pressures for plant structural assessment.

The staff finds the asymmetric-load specification acceptable.

2-32 u

2.2.2.3.7 Alternate' Improved Chugging Load An alternate to the generic improved chugging load was proposed by the owners in Reference 36.

This alternate consists of a minor variation of the basic generic load of 10 sources, wherein only the chugs in sources 801 through 807 are used without averaging of the key chugs with the adjacent d.ug.

All other features of the generic improved chug methodology discussed above are identical with the basic generic load (that is, desynchronization, acous-tic modeling, and FSI considerations).

In support of this specification, a PSD envelope of the seven key chugs was compared to a PSD envelope of seven chugs used for averaging 37 In addition, a PSD comparison was made between these sources and the data from the JAERI freility.88 Based on these comparisons, the staff has concluded that this alternate cieugging specification is conservative and acceptable.

2.2.2.4 Evaluation Summary The staff has reviewed the 4TCO and JAERI chugging data as well as the various documents provided by GE and the Mark II owners to describe and justify their improved generic chugging-load definition.

Based on this review, the staff finds ine generic chugging-load definition as described in References 25 and E6 to be acceptable for Mark II application.

In addition, we find the alternate chugging load specification described in Reference 36 acceptable.

2.3 Downcomer Lateral Loads 2.3.1

Background

After the air-dominated initial pool-swell transient from a postulated LOCA in a BWR pressure-suppression system, steam is vented to the suppression pool.

l During this steam venting, condensation-driven oscillations may occur.

Two types of oscillations have occurred.

The first type occurs during the earli-est portion of blowdown.

The CDs are followed by the second type, called chugging.

During both types of condensation, a downcomer will experience intermittent lateral loading.

However, the 4TCO tests 2o have shown and the staff agrees, that the magnitude of the lateral loading in a Mark II facility during the C0 period is relatively small compared to chugging-induced lateral loads.

Because chugging lateral loads are controlling, the following discus-sion is restricted to an evaluation of these loads.

The load may be the result of asymmetric steam-bubble collapse or the result of the impact on the vent caused by rapidly inflowing water.

In either case, the loads occur near the downcomer exit and have been observed to be impulsive in nature and random in both magnitude and direction.

The stochastic nature of the loads appears unaffected by the proximity of other structures such as containment walls, or another downcomer, as close as three vent diameters away.

Duration of an individual lateral load is typically less than 10 ms.

Lateral loads must be assessed for their effect on individual downcomers as well as for the effect that groups of laterally loaded downcomers have on the diaphragm floor and other containment cemponents.

For the Mark II containment Lead-Plant Program, the Mark II owners proposed a lateral-load specification for lead plants in terms of a static equivalent 2-33

load observed in a foreign licensee test on a single prototypical downcomer.

The 8.8 kibf static equivalent load for a single downcomer corresponded to the maximum observed during foreign licensee tests on a single prototypical do.inc omer. 5 Upon review, the staff determined that the magnitude of this lead plant lat-eral load had to be adjusted according to the downconer natural frequency for a certain frequency range.

Beyond that range, or for downcomers braced in a nonprototypical manner, the staff required a dynamic structural analysis of the downcomer response on a plant-specific basis.

The lead plant multivent load consisted of a static equivalent load. The magnitude of this load was determined from the single downcomer load by adjusting the single-downcomer load with a factor which accounts for the number of vents and the stochastic nature of these loads with respect to force magnitude, direction, and time.

Details of these lead plant load specifications can be-found in Reference 1.

Both the staff and the Mark II owners regarded the lead plant specification as a temporary specification which was to be supplanted by a more realistic dynamic load specification developed by the Mark 11 owners during the course of the Long-Term Program.

In Appendix D, Section B.1 of NUREG-0487,1 the staff stated that it would require a dynamic evaluation of downcomer response to lateral chugging loads for all plants fc11owing completion of the Long Term Program.

2.3.2 Lateral Load Description 2.3.2.1 Single-Vent Load The Mark II Owners have proposed a single-vent, dynamic, lateral-load specifica-tion.88 This load has been proposed for downcomers with up to and ircluding a-24-inch diameter.

The load is represented by a half-sine wave of durations ranging frem 3 to 6 ms for high and low intensities, respectively. The maximum load amplitude ranges from 10 k1bf to 30 k1bf These loads may be considered to be uniformly distributed over 1 to 4 feet of the downcomer end.

The mathe-matical expression for the transient load function can be represented by:

]

F(t) = A sin 1, lateral load (k1bf) where:

A = 10 k1bf = maximum amplitude with I (period) = 6 ms and A = 30 k1bf with t = 3 ms They have also proposed a modification to this load applicable to 28-inch diameter downcomers.40 The modified load is substantially the same as the basic load for 24-inch downcoreers; however, the range of maximum amplitudes is increased by a factor f = 1.34 to account for diameter dependence.

2-34 y

f 2.3.2.2 Lateral Loads on Multiple Vents While the probability that a number of im)ulsively loaded downcomers exper-ience the load in the same direction at t1e same instant is extremely small, a number of the downcomers may have a fraction of the load acting in the same direction at roughly the same time.

Therefore, the total load experienced by a group of downcomers in a certain direction is required in order to evaluate the structural response of the diaphragm floor, the vent bracing system, and any other containment components influenced by loads !cnosed at the downcomer ends.

The basic methodology proposed by the Mark II ow'ners" for conservative assessment of the various downcomer group combinations has not changed from the lead plant specification except that an actual dynamic force rather than a static equivalent load is applied at the end of the downcomer.

The method proposed by the Mark II owners in Reference 41 for assessing the structural response of a given element is given as follows:

(1) Establish the governing parameters (force moment, stress, and so forth) for a H ven structural element under consideration.

(2) Study the influence of a single downcomer dynamic load on the parameters.

(3) Establish an influence zone of multiple downcomers on the structural element.

(4) Arrange the total number of downcomers in the influence zone into various groups of downcomers accounting for the direction of the load such that the maximum value of the parameter is obtained.

(5) For a given number of downcomers in a group, obtain a multivent reduction factor M for the desired probability level.

The tip lateral forcin each downcomer in the group is then given by F(t) = MA(t) sin (nt/t)g for for 0

<t<t.

A(t)=(50-20j)kibffor35t56ms (6) Steps (4) and (5) should be repeated for each assumed grouping for that zone of influence.

(7) The maximum value of the parameters obtained from step (6) shall be used in assessment of the structural element.

The downcomer arrangement and the bracing configuration for Mark II plants vary from plant to plant.

Thus, all but item (5) above must be performed on a plant-specific basis.

Item (5) will be assessed in this report.

To quantify the reduction factor M and subsequent load F(t) mentioned in item (5)lowingcrucialassumptions:pedaprobabilisticapproachwhichemployedthe the Mark II owners develo fol (1) The angle of the chugging impulse on a single downcomer is random and is uniformly distributed around the horizontal plane.

2 35

(2) Because the component of the thugging impulse in a particular direction is of interest, the impulse magnitude distribution is multiplied by the distribution of the cosine of the angle of the force.

(3) The impulse magnitude distribution on any downcomer follows the frequency distribution generated from Figure 3-1 of Reference 42.

This distribu-tion is based on correlated test data.

(4) Because each downcomer vent is identical to all other vents, the magni-tude and direction of the chugging impulse applied to a vent is statistically independent from all other vents.

For this analysis, all vents chug at the same time; the randomness of chugs in time is neglected.

(5) For N downcomers, the resultant impulse distribution for one chug is calculated as the sum of N distributions of the product of the impulse msgnitude distribution and the cosine of the impulse angle.

(6) The number of chags which can be experienced by each downcomer is 265, 2 to 1.5 based on consideration of a range of liquid breaks from 0.1 ft 2

ft,

For plant application, the probability level of step (5) above (probability of exceeding a given impulse component magnitude in 265 chugs) is taken to be 10 4 The corresponding multivent reduction factor M is shown in Figure CDI-1 of Reference 41.

The structural element in question is then assessed assuming an impulse per vent corresponding to the 10 4 probability level for each vent in the influence zone.

2.3.3 Lateral-Load Evaluation 2.3.3.1 Data Base In arriving at their load specifications for both single vent and multiple-vent lateral loads, the Mark II owners relied mainly on the data from the original 4T tests.18'17 However, other tests, both foreign and domestic, are also significant for comparison of single-vent loads as well as to support some of the assumptions used for the multivent methodology.

The staff considered all of these data bases during its review of the loads and the Mark II owners also looked at most of them.

A brief summary of the relevant data bases for dynamic lateral loads along with each one's importance for load specification and evaluation is presented below.

4T18'17:

Data fro.) this single-vent, full-scale facility were used by the Mark II owners to establish the magnitude and duration of their dynamic single-vent specification.

The data were also used by the owners to generate a histogram (frequency distribution, used for the multivent load methodology).

This facility was prototypical in many respects but did not have the drywell located above the wetwell, as is the case in a Mark II containment.

4TC0 0:

The 4T facility mentioned above was made prototypical by putting the 2

drywell directly above the wetwell, and blowdowns with expanded mass flow and break size ranges were conducted.

Lateral-load data from these tests were limited and were used by the staff to confirm statistical trends observed in other tests.

2-36

GKM-II5:

Lateral-load data from this full-scale, single-vent, prototypical foreign ifcensee facility are the most extensive data base for lateral loads from any one prototypical facility.

Lateral-load data were recorded during blowdowns which covered all of the Mark II conditions and even went outside the Mark II range.

The staff used this data base as its primary source for the statistical evaluation of the single-vent lateral loads. While not all data from all tests were available and while the data were for the most part in terms of brace loads and not tip-load resultants, these data still made up the most comprehensive data base for statistical purposes.

G KM-I I-M31:

One of the Mark II plant owners contracted for a test program conducted in this single-vent, full-scale facility by a foreign licensee.

The test facility was designed as a completely prototypical single cell of the

. plant in question and thus was also prototypical of Mark II plants.

Data from this facility was not immediately available to the Mark II Dwners Group but were con?idered by the staff in its review of the lateral-load specification.

KWU-Karl s tei n43: Multivent tests with an unprr surized wetwell were conducteo in this foreign l'censee facility.

Full-scale tests witn one and two vents and half-scale tests with one, three, four, and six vents provided the most pertinent lateral-load data for cultivent configurations.

These tests were important for confirming that lateral loads on individual downcomers are random and statistically independent in magnitude as well as direction.

JAfyl 8'9'27-30:

Tests in this full-scale, seven-vent, protoptyical facility were conducted by JAERI over a wide range of blowdown conditions.

Only limited lateral-load data were obtained from this facility because of instrumentation restrictions.

However, these tests provide direct evidence of the phasing between vents during chugging and, therefore, during lateral loading.

2.3.3.2 Single-Vent Loads 2.3.3.2.1

Background

The single-vent, lateral-load specification for the lead Mark II plants con-sisted of a static-equivalent load.

This load corresponded to the maximum load observed during foriegn licensee tests on a single prototypica.1 downcomer.5 The Mark II owners initiated a task in their Long-Term Program to develop a dynamic lateral load to provide more rigorous criteria for lateral loads.

The load definition was based primarily on the data obtained f rom the three phases of the original 4% test program.

Because the lateral load could not be measured directly, the experimental approach was used to measure the response of the vent system to the load.

The lateral loads were measured by three methods:

(1) strain gages on the vent, (2) strain gages on the bracing, and (3) accelerometers at the vent exit.

A mathematical correlation model was developed by the Mark II owner = to describe the transient ro sonse of the vent expressions for the dynamic load specification.

2-37

The development and solution of the dynamic model is described in Reference 42.

The dynamic model is capable of simulating the transverse response and reaction forces occurring in a single-cell Mark II main-vent downconer system subjected to a dynamic lateral load. The vent system is modeled as a Timoshenko beam of 44 differential equations are those for a vibrating.

finite length.

The governing uniform, prismatic bar.

The model consists of a fourth-order, partial differential equation with its boundary and initial conditions.

The system of equations was solved on a digital computer using a numerical approach employing the method of

" finite differences."

2.3.3.2.2 Load Assessment The result of the model development and the test data evaluation was a dynamic leteral-load specification.

The lateral load on the downcomers was originally specified by the Owners Group as a dynamic " bounding load" derived from the original 4T test data.

This half sinusoidal load of 30 k1bf amplitude and 3 ms duration

• bounded all r

4T test data,42 and it was the premise of the Owners Group that no lateral loads in excess of this value would occur in a Mark II plant.

The NRC asked the Ownerr Group to verify its " bounding load" against appli-cable foreign data (see Section 2.3.3.1) and also check it against whatever data were available from the more recent 4TC0' tests.

This request resulted from observations of some high valuer, in the earlier foreign test data for lateral loads and shortcomings of the original 4T tests with respect to the i

Mark II simulation.

The owners responded by submitting a bounding-load analysis of the Karlstein data and a selected portion of the GKM-II data48 and the 4TCO data.

Simul-taneously, data from the PP&L GKM-II-M tests also became available to the NRC at Data from each of these new tests showed maximum dynamic loads of roughly similar magnitudes to the owners' 30 k1bf/3 ms load which was intended to represent a bounding value for Mark II.

a Another difficulty arose from the size of the data base.

The number of chugs available in each of these tests--or even in all of them together--was sub-stantially less than the number expected to occur in a typical LOCA.

This suggests that either the proposed bounding load was too low, and/or a true physical bounding load had not been measured in any of the available data bases, which reflected only a statistical distribution of chug strengths characteristic of the randomness in the chugging process.

The NRC asked thrs J

owners to examine their lateral-load data from the viewpoint of a probability

• The owners' bounding load is'actually presented as one where the ampl{~tude decreases linearly with duration from 30 to 10 k1bf as the duration increases from 3 to 6 ms.

It is, however, near the 30 k1bf/3 ms end of this specification where the peak impulse and, almost invariably, the most critical loading condition for the downcomer occurs.

2-38

distribution; at the same time, the NRC initiated its own statistical inves-tigation of the available data.

The GKM-II, 4TCO, and the PP&L GKM-II-M tests are representative of Mark II conditions and, therefore, comprise the most appropriate data.

Of these, the 4TCO data base is sparse with respect to lateral loads. The most extensive data, and the most useful for a statistical analysis, come from the GKM-II tests.

The results of nine (KM-II tests were utilized in the study by the NRC staff consultants (Lehnet).

These tests covered various conditions of mass-flow and pool temperature.

The data comprise some 2000 usable chugs, divided into nine different pool-temperature / mass-flow conditions.

The tests cover much of the chugging map applicable to Mark II conditions.

Exceedance probabilities

• were determined for each of the tests and for the whole body of data.

The exceedance probabilities for the two tests covering low-r. ass-flow and cold pool conditions had a higher probability of high lateral loads than for the total bod of data in the nine GKM-II tests provided in Reference 43.

The-data from thes No tests comprised about one quarter of the data base provided in Reference 44.

The NRC staff and its consultants made certain deductions from the statistical data.

First, there is no indication that the proposed load s)ecification, 30 k1bf, represents some " bounding value" of the load The GK4-II-M data stretch to 30 kibf, and slightly beyond, without any sign of a downturn, such as might be exaected when some absolute physfcal bound of a statistical quantity is approacled.

Secondly, a calculation was made of the probable number of exceedances of the proposed 30 klbf lateral-load specifica-tion for a typical Mark II plant with 100 downcomers and for a long-duration 265 pool-chug LOCA. The data from the low-mass-flow cold-3001 tests were used. These calculation > indicate that the proposed 30 klaf will be exceeded over a 100 times

• in an extended cold pool LOCA.

Calculations for a more likely 100 pool-chug LOCA indicate that the 30-k1bf load will still be exceeded about 40 times.*

In addition to the GKM-II data, the staff and its consultants evaluated the proposed load against data from the 4TC0 tests and the PP&L GKM-II-M tests.

I For the 4TCO tests, data from a total of five relatively cold pool tests were utilized.

Two of the five runs were liquid breaks, in contrast to the GKM-II data which were from steam breaks.

The lateral-load data from the 4T00 tests was too ifmited to allow any deductions about the very low probability high load occurrences.

ilowever, the data are not inconsistent with the GKM-II data.

a Exceedance probabilities corresponding to the PP&L GKM-II-M data were also calculated.

Two tests with a derate mass flow and cold pool temperatures were studied, because the highest lateral loads were found to occur most frequently under these conditions.

There are about 300 chugs in this~two-test data base.

The a PP&L GKM-II-M data fall in line with the earlier GKM-II data, although they are too limited to st7tch to the lower probability high-load region.

Insofar as the data go, however, they would suggest a somewhat lower cold pool probability than the GKM-II data, in line with the exceedance probabilities for the general GKM-II data for the body of GKM-11 data in the nine tests provided in Reference 43.

"Exceedance probability of the load F is defined as a ratio of the number of chugs with a force larger than F to the total number of chugs in the data base.

2. ;9 l

Based on the above study (Lehner), the NRC staff has concluded that the proposed l

30-kibf dynamic lateral load could be exceeded an unacceptably large number of.

l times in LOCAs correspondi'g to small-or moderate-line breaks.

The lateral leads corresponding to low probabilities were determined by the l

staff and its consultants.

The GXM-II data discussed above were used in these calculations.

The cold pool / low-mass-flow data were fitted with the exponential

' form l

-F P(F) = e F-where:

P(F) = exceedance probability of the load F F = lateral load and a = an empirical constant that fits the cold pool data This exponential fit of the data was used to determine a load value that would ensure an exceedance probability that is small compared with one time per LOCA under all LOCA conditions.

The NRC has specified a revision to the single-vent laad specification 3rovided in Section 2.3.2.1.

This load criterion consists l

of a dynamic load t1e same es that discussed in Section 2.3.2.1 with an ampli-tude of 65 k1bf and a 3-ms period.

With this load, the probable number of exceedances would be below 10 2 per LOCA based on all the CKM-II data and of the order of 0.1 based on the cold pool subset of the data.

2.3.3.2.3 Conservatisms in the Methodology The NRC acceptance criteria for single-vent dynamic-load evaluation consist of the impulsive load specification provided in Section 2.3.2.1 of this report, along with an additional case wit 1 a load amplitude of 65 k1bf and a 3-ms period.

This specification ~ includes a number of conservatisms not discussed in the previous section.

An accurate quantification of these conservatisms was not developed; however, the nature of these conservatisms is discussed below.

High-Amplitude, Short-Period Load Experimental observations indicate that the lateral load is an impulsive load.

The low-intensity chugs are best characterized by a long duration (F = 10 kibf, T = 6 ms).

The infrequent high-intensity chugs are characterized by a short duration (F = 30 k1bf, t = 3 ms).

No credit was taken for this data trend in establishing the very high-intensity, low probability load specifi-cation (F = 65 kibf, T = 3 ms).

A large data base for the limiting cold pool /

low-mass flow case would be needed to confirm a significant reduction in the load duration for very high-intensity vent loads.

"Inese estimates of the numt'r of exceedances were obtained directly from the experimentally determined exceedance probability curves.

There was no need to extrapolate the exceedance probability curves.

2 40

Load-Combination Considerations The 65-kibf lateral-load specification corresponds to an exceedance prob-l ability of 0.1 per LOCA based on the cold pool subset

• GKM-II data and 10 2 per LOCA based on all the GKM-II data in nine tests.

T.is load is used directly l

in combination with other maximum loads, such as the S'..y and the seismic dynamic load, for the evaluation of structures, pipirj, and equipment.

The l

probability that the 65-k1bf vent load would occur together with other maximum events is much lower than that for a LOCA taken by itself.

I Low-Probability Extrapolation the exceedance probabilities corresponding to the 65-k1bf load criteria were obtained by fitting the GKM-II data to a negative exponential distritution and extrapolating the exceedance probability curve to low probabilities.

The negative ex)onential distribution provided a good fit for the GKH-II data.

However, otler data uts appear to be better fit with a normal or log-rmmal distribution.

The extrapolation of these distributions would indicate Icwer exceedance probabilities for the 65-kibf vent load.

2.3.3.2.4 Load Assessment for 28-Inch Vents Most of the downcomers in domestic Mark II facilities have a diameter of 24 l

inches or smaller.

However, there are a few icolated cases where a larger 28-inch downcomer is utilized.

As discussed in Mction 2.3.2.1 the basic load specification for a 24-inch downcomer is multiplied by a factor of 1.34 to account for the observed increase in lateral loads with vent diameter.40 The lateral load specification for 24-inch downcomers was based on experimental observations in a aumber of test facilities with downcomers having diameters I

in the range of 20 to 24 inches.

Foreign data were also available for 12-inch diameter downcomers.

The dependency of lateral-load amplitude on diameter was obtained by correlating these experimental data.

Correlation of both the i

maximum values and the average values of laterai load demonstrates th-Nad l

factor of f = 1.34 is conservative for the limited extrapolation of the a

l based on experiments with 24-to 28-inch downcomers.

i 2.3.3.3 Lateral Loads on Multiple Vents l

?.3.3.3.1 Assessment of Underlying Assumptions To properly evaluate the multivent, lateral-load methodology propcsed by the Mark II owners, it is necessary to assess the merit of each one of the assump-tions--(1) through (6), stated in Section 2.3.2.2--on which the method is based.

The staff agrees with assumption (1) that the angle of the chugging imp'ulse is random and uniformly distributed.

All available single-vent data show that no preferential lateral 1cading direction exists with prototypical Mark II downcomers.

More importantly, the multivent Karlstein data obtained from tests conducted in a highly asymmetric pool with asymmetric bracing show that the randomness of direction is not affected by pool geometry or the proximity of other vents or pool boundaries.

2-41

In the staff's judgment, assumption (2), the multiplication of the impulse magnitude by the cosine of the angle (that is, selecting a force component rather than a resultant), is not always justified.

The validity nf the assumption depends on the particular loading situation:

If the structure under consideration is loaded significantly enly when the load imposed has a particular direction with respect to a fixed reference frame, then multiplication by the cosine is justified.

If the zone of influence causes the same stresses in the structure irrespective of the direction of applied load then the resultant force, not the component, should be used and cosine multiplication is not justifie:i.

By using tne component rather than the resultant, the probability of exceeding a given force in an arbitrary direction will be underestimated.

However, calcula-tions by the staff's consultants have shown that at the chosen probability level of 10 4 the errors are limited to small values.

For instance, the impulse per downcomer for two downcomers is estimated too low by 6%.

This percentage error increases slightly as the number of downcomers increases.

The staff concludes that conservatisms (to be discussed below), some of which increase substantially with the number of downcomers, more than offset the small underestimation here.

The staff agrees only partially with am.ption (3) that the frequency distribu-tion generated from Figure 3-1 of Reference 42 is appropriate for use with any j

downcomer for the multivent method.

Certainly, it is expedient and reasonable to pick a single distribution from.which to select the impulse magnitude.

The distribution chosen by the owners Is conservative with respect to the lateral loads recorded during *he 4T tests.

The mean value and overall shape of the distribution is conservative for lateral-load data in general.

However, the 4T data ends abruptly at tip loads of 3 kips and 3 ms duration, and therefore does not contain impulses above 60 kibf-ms.

The likelihood of low probability, high-load (65 kips) values implied by the extrapolation procedures outlined in Section 2.3.3.2.1 was the reason for the staff's modification of the single-vent lateral load discussed in Section 2.3.3.2.2.

If a distribution having a

" tail" reflecting the low probability but high-value end of the impulse spectrum i

and the selected 4T distribution are used to calculate the pisbabilities of exceeding the difference 7 tween the two is much less for mul.ivent a) plication than for single-vent appl,c, tion.

Even for only two vents, the proba)ility of two high-impulse, low probatility events occurring at the same time and oriented in similar directions is extremely small. With more vents in a cluster, the difference in the calculations performed with the two different frequency

);

distributions quickly becomes negligible, i

Calculations were performed by the staff's consultants with the selected 4T 3

distribution and one identical to the owners except for an ex)onential (infinite) tail for high impulses.

These calculations show that, with t1e owners' specified distribution at the 10 4 probability level, the impulse is underestimated by 16% for two vents, 4% for four vents, and 3% for eight vents.

The staff concludes that conservative as?ects of the method, to no discussed below, offset the underestimates whic1 may occur by using the owners' specified frequency distribution in the multivent calculations.

i The staff agrees with assumption (4) of the owners that for multivent lateral-load calculations each vent is statistically independent from all other vents and can be treated identically with all other vents regardless of location in 2-42 4

l

the wetwell. This assumption agrees with what is known about the re.ndonness of the chugging process in general and is specifically borne out by the large number of lateral-load observations made in the Karlstein test facility using full-scale and half-scale vents in a highly asymmatric wetwelles.

The other part of assumption (4)--that for purposes of calculation all chugs are assumed to occur at exactly the same time--is a very conservative assumption. - This conservatism will be discussed more fully below.

Assumption (5) in the staff's judgment is a reasonable way to calculate the resultant impulse distribution for N downcomers.

]

The use of 265 as the largest number of chugs which a downcomer can experience, as assumed in (6), is a reasonable basis in the staff's opinion.

In summary, the staff finds assumptions (1), (4), (5), and (6) to be acceptable.

Assumptions (2) and (3) contain some nonconservatisms, but the staff beltsves l

these are offset by larger 1,nherent conservatisms which are discussed below.

}

2.3.3.3.2 00nservatisms in the Proposed Methodology 4

There are three principal conssrvatirms in the lateral-load multivent methodology j

proposed by the Mark II owners.

(1) The first conservatism arises from the way load combinations are treated.

J The multivent lateral loads for which each structural element is assessed are assumed to occur in combination with other loads such as eartnquake i

loads or SRV actuations. The methodology presumes that these combinations

)

exist throughout the 265 chugs or a time span on the order of 500 seconds.

i In reality, seismic and SRV loading would be confined to much shorter i

periods.

Therefore, assessing the structural element with impulses corresponding to probabilities for 265 chugging events introduces c.nsider-i able conservatism into the method.

i j

(2) As stated in assumption (4) of Section 2.3.2.2 phasing between vents is neglected and all vents are assumed to chug at exactly the same time.

This is a very conservative assumption, even for two vents, and it becomes j

even more conservative as the number of vents in a cluster is increased.

All available multivent data (with those of Reference 26 being the most i

typical) show that vents do not chug simultaneously, but rather that l

chug-start times between vents differ from a few to several tens of milliseconds. According to Neference 26, a conservatively small time l

window from the first to the last vent chug in the seven-vent cluster of i

the JAERI facility can be shown to be 50 ms.

Because the lateral load on a single vent for high impulses has a duration of only 3 or 4 ms, even a slight difference in start times will significantly reduce the resultant load from a cluster of vents. Therefore, assuming all vents in a zone of influence to be impulsively loaded at exactly the same time is-a very conservative assumption.

i (3) Choosing the probability of exceeding a particular impulse component in 265 chugs as 10 4 to obtain impulses per vent for structural element assessment is a conservative assumption.

Such a choice provides an additional margin for the specification of multivent loads beyond the already conservative probabilities for the single-vent lateral loads.

2-43 4

n.

-a

-mvn,

r

In comparing the possibly nonconservative aspects of assumptions (2) and (3) of Section 2.3.2.2 with the conservatisms just stated above, the staff finds the following:

For the very worst case, which occurs with a two-vent cluster, j

the combined effect of (2) and (3) is an underestimate of the impulse per vent of 22% at the 10 4 probability level.

This is easily more than balanced by J

conservatisms (1) and (2) above.

For a greater number of vents per cluster, the percentage underestimation is less and the conservatisms greater, so that the overall load specification becomes even more conservative.

Therefore, the staf f finds the multivent lateral-load methodology as specified by the Mark II owners acceptable.

2.3.4 Evaluation Summary The NRC has reviewed the single-vent dynamic-load specification proposed by the Mark II owners for downcomers up to 24 inches in diamster. With the tddition of a high-intensity load case of amplitude 65 kibf and period 3 ms, the staff finds the proposed load specification conservative and acceptable.

For large-diameter, 28-inch downcomers, the staff finds that a value for the load multiplier of 1.34 applied to the basic load specification for 24-inch downcomers is conservative and acceptable.

With respect to the proposed multivent lateral-load specification, the staff has reviewed the appropriate documentation with the corresponding data and finds the hydrodynamic load application aspects (part (5) of the outline in l

Section 2.3.2.2) of the methodology acceptable as specified by the Mark II owners.

The other portions of the methodology (items (1-3) and (6 and 7)) are plant specific. The staff's acceptance of the mnthodology is Losed on the use of a probability of exceeding a specified impulse per vent in 265 chugs equal to 10 4 2.4 Submerged-Structure Drag Loads 2.4.1

Background

l l

The NRC acceptance criteria for loads on submerged structures for Mark II lead plants are listed in NUREG-04871 and modified in Supplement 1 to that report.2 For LOCA-related loads, the acceptance criteria are based primarily upon the methodology presented in GE report NEDE-21730.48 This methodology divites the LOCA event into the water-jet period, LOCA-air-bubble period, pool-swell period, and pool-fallback period.

To remove the extreme conservatism associated with the crude modeling of the water jet, the Herk II owr.ers included the deve'opment of analytical modeling known as the Ring Vortex Model as Task A.5.7 in the Long-Term Program. This model was presented as an alternate to the methods described in References 1 and 2.

The technical basis, experimental verification, and a proposed load specification are provided in Reference 47.

With the exception of this load, the Mark II owners have adopted the conservative lead plant submerged-structure-loads acceptance criteria for plants following the lead plant.

2-44

2.4.2 LOCA-Water-Jet Ring Vortex Model Evaluation The fundamental assumptions and the mathematical formulation of the model in Reference 47 as an inviscid flow with vorticity generated at the vent exit are welljustified up to the point of vent clearing.

Questions associated with boththenumerIcsandboundaryconditionshavebeenaddressedbytheMarkII owners and shown not to have serious consequences.

The method has been further verified by generating a Hill's vortex, a known analytical solution. Within the limitations of the single-cell cylindrically symmetric model, the analysis is capable of adequate representation of the physical phenomena.

The model is verified by comparison to the EPRI (1/13-scale) tests, QSTF (1/4-scale) tests, and 4T (full-scale) tests.

The most detailed comparison is made in the EPRI tests where both the flow visualization of the jet boundary and wall pressure data are compared to the model results.

While the main features of the phenomena are accurately represented by the model, the precise fet penetration and spreading history are strongly dependent on the vent-exit fet-velocity history.

Conservati;m in the load specification will depend strongly on the conservatism in the drywell pressure history, and the model used to deduce vent-exit velocity.

In addition, there appears to be an under-estimation of the jet spreading, especially past vent-clearing time, suggesting thal. the effect of the air bubble may be significant to the transverse motion of the water near the vent exit.

Pressure comparisons for both EPRI and 4T tests suggest that the model contains the essential features of the physical process.

The comparison further suggests that, at least in terms of its gross effect on the pool, the post-vent-clearing model is adequate in terms of predicting the wall pressures up to peak values.

Care should be exercised in concluding that the detailed flow field near the vent exit must, therefore, be similarly well modeled.

The load specification based on the Ring Vortex Model is divided according to the type of structure (cylindrical or not) and location (horizontal or vertical, withinoroutsideofthejet).

Arguments are further presented for the specific drag and inertia coefficients used and for the procedure by which the multivent effect is included.

Each of these specifications is examined below.

Horizontal Cylindrical Structures Below Vent Exit, Outside Jet Boundary The owners have proposed that the model be used directly to compute local velocity and acceleration and that loads be computed on the basis of a drag coefficient of 1.2, an inertia coefficient of 2.0, and negligible lift.

These factors provide adeguate conservatism provided there is assurance that the vent-exit velocity is conservatively computed for the plant.

To ensure this, the staff requires that the vent-clearing model for predicting drywell pressure history must be applied conservatively to yield the highest values of vent-exit velocity and acceleration at vent clearing.

Horizontal Cylindrical Structures Below Vent Exit, Inside Jet Boundary The use of vent-exit velocity and acceleration and the use of jet impact load involving total momentum transfer provide a large degree of conservatism.

2-45

i Vertical Structures Adjacent to the Vent To compensate for the lack of conservatism associated with the lower jet spreading in the model, the report proposes that the most conservative load l

along any vertical line be applied to the structure regardless of the structure location.- Alternatively the LOCA-bubble charging load of the lead plant l

criteria may be used.

Because it is very difficult to determine if the suggested l

use of the model incorporates any conservatism, a comparison to the LOCA l

charging load should be always made and the larger of the loads used for assessment.

Noncylindrical Structures l

The different values of drag, inertia, and lift coefficients for noncylindrical structures are recognized.

The alternative use of the equivalent diameter as l

specifiedinReference48isacceptablesubjecttothelimitationslistedin l

Supplement 1 of NUREG-0487.2 1

Structures Above Vent Exit Pool-swell and fallback loads are considered limiting; therefore, no direct use of the ring vortex model is proposed.

This is reasonable because peak vertical velocity and acceleration will occur once the water slug is moving in i

virtually one dimension.

1 Multivent Flows The expected inphase clearing of vents is used to justify the use of the single-cell results which are insensitive to the location of the cell boundaries within regions of significant loads.

If the LOCA bubble load is computed on the basis of a single vent, that load should provide a bounding asymmetric load for vertical structures between two vents.

The proposed peak load applica-tion of the model may very well be conservative, as well, although that may be more difficult to demonstrate.

Interference Effects No " side-by-side" interference is expected for structures in Mark II plants, l

and closely spaced structures in the flow direction are to be treated by the methodo?agy described in Reference 48.

The use of the criteria in Supplement 1 to NUREG-04872 is sufficient to require computation of interference effects whenever they are significant.

l Boundary Pressure Calculation 1

The only use made of the boundary pressure results is to show that the DFFR method yields conservative values compared to the model.

The staff agrees with that statement.

Because no proposal is made to change the boundary-pressure load specification, this statement has no effect ca any acceptance criteria.

i 2-4S

2,4.3 Evaluation Summary The staff finds that the ring vortex model as described in Reference 47, subject to the restrictions provided in Section 2.4.3 (items (1), (3), and (6)),

provides a conservative methodology for the calculation of LOCA submerged-structure drag loads during the water-jet and air-bubble periods.

The staff finds this method an acceptable alternative to the lead plant load criteria of References 1 and 2.

l 4

f 4

i i

i 2 47

3 RESOLUTI N OF THE ISSUE 3.1 Implementation of Guidelines This report identifies those LOCA-related pool dynamic Inads for Mark II containments that are acceptable to the staff.

These include a combination of generic loads proposed by the Mark II Owners Group and acceptance riteria developed by the NRC.

All applications / licenses for Mark II plants, including those reviewed by the NRC against the criteria in NUREG-0487 and Supplements 1 and 2 to thet reporttenes should be reviewed against the loads referenced in Appendices A and C of this report.

Applicants / licensees should identify deviations from the generically acceptable loads and justify these deviations.

Plant-unique progrees for two Mark II facilities have been identified in support of plant-specific chugging and condensation-oscillation loads that differ from the generic loads referenced in this report.

These plant-specific programs / loads are being reviewed by the staff on a plant-specific basis.

The results of these reviews will be reported in the NRC Safety Evaluation Report fnr each plant.

3.2 _R_ecommended Changes to the Standard Review Plan To incorporate the loads referenced in this report, the following chenge to the Standard Review Plan should be made:

Standard Review Plan Section 6.2.1.1.c II.

Acceptance Criteria Modify item 2. as follows:

The acceptability of pool dynamic loads for plants with Mark II containments is based on conformance with generic loads previously reviewed and found acceptable by the NRC and NRC acceptance criteria. These loads and criteria are in Appendices A and C of NUREG-0808, and are reproduced in Appendix to this section of the Standard Review Plan.

3.3 Recommended Development of a Regulatory Guide It is recommended that a Regulatory Guide be developed referencing the loads lof Appendices A and C of this report hs acceptable loads for the evaluatinn of plants with Mark II containment designs.

3-1

4 REFERENCES References cited in this report are available as follows:

Those items marked with one asterisk (*) are available for inspection in the NRC Public Documen'. Room, 1717 H St., Washington, DC 20555; they may be copied for a fee.

Material marked with two asterisks (**) is not publicly available because it contains proprietary information; however, a nonproprietary version is available for inspection in the NRC Public Document Room and may be copied for a fee.

Those reference items marked with three asterisks (***) are available for purchase from the NRC/GPO Sales Program, U. S. Nuclear Regulatory Commission, Washington, DC 20555, and/or the National Technical Information Service, Springfield, VA 22161.

All other material referenced is in the open literature and is available through public technical libraries.

(1)

U. S. Nuclear Regulatory Commission, " Mark II Containment Lead Plant Program Load Evaluation and Acceptance Criteria," USNRC Report NUREG-0487, October 1978.***

(2)

U. S. Nuclear Regulatory Commission, " Mark II Containment Lead Plant Program Load Evaluation and Acceptance Criteria," USNRC Report NUREG-0487, Supplement 1, September 1980.***

(3)

U. S. Nuclear Regulatory Commission, " Mark II Containment Lead Plant Program Load Evaluation and Acceptance Criteria," USNRC Report NUREG-0487, Supplement 2, February 1981.***

(4)

L. L. Myers, T. R. McIntyre, and R. J. Ernst, " Mark III Confirmatory Test Program Phase I - Large-Scale Demonstration Test, Test Series 5701 through 5703," General Electric Proprietary Report NEDM-13377,,L tober 1974.**

(5)

T. Y. Fukushima and others, " Test Results Employed by GE for BWR Containment and Vertical Vent Loads," General Electric Proprietary Report, NEDE-21078-P, October 1975.**

(6)

K. Namatame and other, " Full-scale Mark II CRT Program, Facility Description Report," JAERI-M 8780, March 1980.*

I (7) General Electric Company, " Mark II Containment Dynamic Forcing Function Information Report," General Electric Topical Report NEDO-21061, Revision 0, November 1975.*

(8)

Y. Kukita, K. Namatame, N. Yamamoto, and M. Shiba, " Full-Scale Mark II CRT Program, Data Report 4 (Test I?>1)," JAERI Report M-8763, March 1980.*

4-1

f (9)

I. Takeshita, N. Yamamoto, Y. Kukita, K. Namatame, and M. Shiba, " Full-Scale Mark II CRT Data Report, Data Report 8 (Test 1201)," JAERI Report M-8887, 1

June 1980.*

(10) I. Takeshita, N. Yamamoto, Y. Kukita, K. Namatame, and M. Shiba, " Full-Scale Mark II CRT Data Report, Data Report 9 (Test 1202)," JAERI Report M-8961, July 1980.*

(11) I. Takeshita, Y. Kukita, N. Yamamoto, K. Namatame, and M. Shiba, " Full-Scale Mark II CRT Data Report, Data Report ID (Test 1203)," JAERI Report M-9403, March 1981.*

(12) Y. Kukita, I. Takeshita, N. Yamamoto, K. Namatame, and M. Shiba, " Full-Scale Mark II CRT Data Report, Data Report 11 (Test 1204)," JAERI Report M-9404, March 1961.*

(13) I. Takeshita, Y. Kukita, N. Yamamoto, K. Namatame, and M. Shiba, " Full-Scale Mark II CRT Data Report, Data Re, ort 12 (Test 1205)," JAERI Report M-9405, March 1981.*

(14) A. J. James, "The General Electric Pressure Suppression Containment Analytical Model," General Electric Topical Report NEDM-10320, March 1971.*

(15) Y. Kukita, " Analysis of Internal Pressure flesponse of the Mark II Containment During a Coolant Loss Accident," JAERI Report M-7504, January 1978.*

(16) T. R. McIntyre, M. A. Ross, and L. L. Myers, " Mark II Pre sure Suppression Test Program -- Phase I Tests," General Electric Proprietary Report NEDE-13442P-01, May 1976.**

{

(17) T. R. McIntyre and others, " Mark II Conta!nment Pressure Suppression Test Program, Phase II and II Tests," General Electric Proprietary Report NEDE-13468-P, December 1976.**

l (18) Letter from S. J. Stark, General Electric, to K. Kniel, NRC, Subject-

" Mark II Containment Program Differential Pressure Upload on Diaphragm l

Floor," Letter Number 125-81, June 1981.**

(19) General Electric Company, " Mark II Containment Dynamic Forcing Function Information Report," General Electric Topical Report NEDE-21061, Revision 3 June 1978.*

(20) P. F. Bird and others, "4T Condensation Oscillation Test Program Final Tr.st Report," General Electric Proprietary Report NEDE-24811-P, May 1980.**

(21) General Electric Company, " General Condensation Oscillation Load Definition Report," General Electric Proprietary Report NEDE-24288-P, November 1980.*

(22) Y. Kukita and others, "The Satistical Evaluation of Steam Condensation Loads in Pressure Supression Paol (1)," JAERI-M 9665, August 1981.

4-2

(23) General Electric Company and S. Levy, Inc., " Condensation Cscillas ion (CO) Load Data for LaSalle," July 1980.** (Also published in the LeSalle County Station Design Assessment Report.)

(24) General Electric Company, "4TCO Chugging Data Report - Six Key Runs,"

General Electric Proprietary Report NEDE-24285-P, January 1981.**

(25) General Electric Company, " Mark II Improved Chugging Methodology," General Electric Proprietary Report NEDE-24822-P, May 1980.**

(26) General Electric Company, " Generic Chuggirg Load Definition Report,"

General Electric Proprietary Report NEDE-24302-P, April 1981.**

(27) Y. Kukita and others, " Full-Scale Mark II CRT Program Data Report No. 1 (Test 0002)," JAER1-M 8598, February 1979.*

(28) Y. Kukita and others, " Full-Scale Mark II CRT Program Data Report No. 2 (Test 00d3)," JAERI-M 8761, February 1979.*

(29) Y. Kukita and others, " Full-Scale Mark II CRT Program Data Report No. 3 (Test 0004," JAERI-M 8762, February 1979.*

(30) Y. Kukita and others, " Full-Scale Mark II CRT Program Data Report No. 5 (Test 2101)," JAERI-M 8764, April 1979.*

(31) Pennsylvania Power and Light Company, " Design Assessment Report,"

l Susquehanna Steam Electric Station Units 1 and 2, Docket Nos. 50-387 and 50-388.*

(32) GKSS Trip Report from E. W. McCauley, LLL, to T. Lee, NRC,

Subject:

" Preliminary Data from Tests M2, M3, M4, MS, and MS-2," October 1980.**

(33) J. A. Block, F. X. Dolan, and B. R. Patal, " Comparison of Single and Multivent Chugging-Phase 2," General Electric Proprietary Report NEDE-25289-1-P, August 1980.**

(34) General Electric Company, " Scaled Multivent Chugging Test Program,"

General Electric Proprietary Report NEDE-24300-P, April 1961.**

l (35) P. M. Morse and F. U. Ingard, Theoretical Acoustics, McGraw Hill Book Company, 1968.

(36) Letter from H. Chau, Long Island Lighting Company, to K. Kniel, USNRC,

Subject:

" Alternate Improved Generic Load Definition," July 1981.

~

(37) Long Island Lighting Company, " Design Assessment Report," Shoreham Nuclear Power Station, Docket No. 50-322.

(38) Letter from H. Chau, Long Island Lighting Company, to W. Butler, NRC,

Subject:

" Additional Information on Shoreham Condensation / Oscillation and Chugging Loads," Jur.e 15, 1981.

(39) W. M. Davis, " Mark II Main Vent Lateral Loads," General Electric Proprietary Report NEDE-23806 P, October 1978.

4-3

(40) Letter from R. Buckholz, General Electric, to K. Kniel, NRC,

Subject:

" Mark II Containment Program Response to NRC Questions on March II Single Vent Lateral Loads," January 16, 1981.*

l L

(41) Letter from R. H. Buckholz, General Electric, to J. F. Stolz, NRC,

Subject:

" Mark II Containment Program Method of Applying Mark II Single Vent Dynamic Lateral Load to Mark II Plant With Multiple Vents," Aprii 9, 1980.**

(42) General Electric Company, " Dynamic Lateral Loads on a Main Vent Downcomer-Mark II Containment," General Electric Proprietary Report, NEDE-24106-P, March 1978.**

(43) J. E. Borhaug, " Dynamic Lateral Loads on Mark II Main Vent Downcomer--

Correlation of Independent Reference Data," General Electric Proprietary Report NEDE-24794-P, March 1980.**

(44) S. P. limoshenko, " Vibration Problems in Engineering," Second Edition.

D. Van Nostrand Company, Inc., New York, 1937.

(45) S. T. Nomanbhoy, " Load on the Vent Struts Due to Condensation of Steam in a Water Pool " General Electric Proprietary Report NEDE-23627, June 1977.**

(46) R. J. Ernst, T. G. Peterson, and G. H. Salas, " Mark II Pressure Suppression Systems Loads on Submerged Structures--An Application Memorandum," General Electric Proprietary Report, NEDE-21720, December 1977.

(47) Burns & Roe, Inc., "An Analytical Model for LOCA Water Jet in the Mark II Containments (The Ring Vortex Model)," Burns & Roe Propriotary Report, October 1980.

(49) Cincinnati Gas and Electric Co., " Final Safety Analysis Report", Amendment 13, Appendix G, Zimmer Nuclear Power Station, Docket No. 50-358.**

4-4

APPENDIX A NRC ACCEPTANCE CRITERIA

• MARK II LOCA-RELATED P00L DYNAMIC LOADS A.

Pool-Swell Load 1.

Pool-Swell Llavation the maximum pool-swell height shall be taken as the greater of (a) or (b) as follows:

(a) 1.5 times the vent submergence; (b) the elevation corresponding to a drywell floor uplift AP = 2.5 psid.

The pool surface elevation cor:'espcading to the maximum wetwell air compression will be calculated assuming a polytropic process with an exponent of 1.2.

2.

Pool-Swell Velocity The pool-swell velocity used to deteraine impact and drag loads on wetwell components shall consist of the velocity predicted by the pool swell analytical model described in NEDE-21544-P1 multiplied by i

a factor of 3.1.

3.

Impact / Draft Loads on Grating The static drag load on grating in the pool-swell zone of the wetwell shall be calculated for grating with open area greater than or equal to 60% by forming the product of the pressure differential as given in Figure 4-40 of NED0-21061, Revision 2, and the total area of the grating.

To account for the dynamic nature of the initial loading, 2

the load shall be increased by a multiplier given by F /D = 1 + e + (0.0 m W for SE Wf < 2000 in/sec, where:

FSE = static equivalent load W

= width of grating bars, in, f

= natural frequency of lowest mode, Hz D

= static drag load "These LOCA-related criteria replace those presented by the staff in NUREG-0487 and Supplements 1 and 2.

SRV pool dynamic load criteria for Mark II containment are addressed by the staff's A-39 program.

A-1

i i

4.

Asymmetric Bubble Load A load equal to 20% of the maximum LOCA vent-clearing bubble pressure is to be applied to 1/2 of the submerged boundary.

This load is to be applied statically together with normal hydrostatic pressure to l

the submerged portion of the containment.

5.

Impact loads ori Small Structur,y Thehydrodynamicloadingfunctionthatcharacterizespoolimpacton small horizontal structures shall have the versed sine shape P(t) = P (1-cos2nf) max where:

p

=pressureactingontheprojectedareaofthestructure, psi thetemporalmaximumofpressureactingontheprojected pmax = area of the structure, psi t

= time, sec 1

= duration of impact, sec For both cylindrical and flat structures, the maximum pressure pmax and pulse duration I will be determined as follows:

(a) The hydrodynamic mass per unit area for impact loading will be obtained from the appropriate correlation for a cylindrical or flat target in Figure 6-8 of Reference 3.

(b) The impulse will be calculated using the equation MH 1

IP*F V x (32.2) (144) where:

I

= impulse per unit area, psi-sec p

M

[=above 2

hydrodynamic mass per unit area, 1bm/ft, from (a)

V = impact velocity, ft/sec, rietermined according to Section A.2

^5 mall structures, in the present context, are defined as pipes, I-beams, and other similar structures having one dimension less than or equal to 20 inches.

The acceptance criteria are not applicable tn the determination of ovaling stresses in cylindrical pipes.

A-2

(c) The pulsa duration will be obtained from the equation a

t T = 0.0 630 (cylindrical target) t = 0.011 W Wec (flat target) t = 0.0016 W V < 7 ft/sec where:

I = pulse duration, sec D = diameter of cylindrical pipe, feet W = width of the flat structure, feet j

V = impact velocity, ft/sec (d) The value of P will be obtained using the following equation:

max P,,x = 2I /t p

For both cylindrical and flat structures, a margin of 35% will be added to the P values (as specified above) to obtain j

max conservative design loads.

The load acceptance criteria, as specified above, corresponds to impact on rigid structures.

The effect of finite flexibility of real structures will be accounted for in the following manner: When structural dynamic analysis is performed, the

" rigid body" impact loads will be applied; however the masses oftheimpactedstructureswillbeadjustedbyaddIngonthe hydrodynamic masses of impact.

The numerical values of hydro-dynamic masses will be obtained from the appropriate correlations for cylindrical and flat structures in Figure 6-8 of Reference 3.

B.

Steam Condensation and Chugging Loads 1.

Single-Vent Lateral Loads i

The following dynamic single-vent load specification will be used:

(a) A tip lateral force given by:

F(t)=A(t) sin (f)0it1I whereA(t)=(50-20j)k1bffor3ms1ti6ms shall be applied to each downcomer with I varied between 3 and 6 ms as indicated.

A-3

i i

In addition, a separate assessment shall be made for a load with a tip lateral force of F(t)=65 sin ($1)kibf0iti3as for each downcomer.

(b) For 28-inch downcomers the single-vent and multiple-vent lateral load shall be scaled up by a factor of 1.34 above the acceptable loads for a 24-inch downcomer.

C.

References A nonproprietary versions of the reports marked with an asterisk is available for inspection and copying for a fee in the NRC Public Document Room (PDR). 1717 H St., N.W., Washington, DC 20555.

The document marked with two asterisks also is available for inspection and copying for a fee in the PDR.

1.

R. J. Ernst and M. G. Ward, " Mark II Pressure Suppression Containment Systems:

An Analytical Model of the Pool Swell Phenomenon," General Electric Proprietary Report NEDE-21544-P, December 1976.*

2.

General Electric Company, " Mark II containment Dynamic Forcing Function Information Report," General Electric Topical Report NEDO-21061 Revision 2, September 1976.

3.

T. R. McIntyre, " Mark III Confirmatory Test Program, One-Third Scale Pool Swell Impact Tests Test ?;.;es 5805," Gtneral Electric Proprietary Report NEDE-13426P, Augu.s 1975.*

A-4

APPENDIX B MARK II OWNERS GROUP SUPPORT PROGRAM Pool dynamic loads for the evaluation of the Mark II ccatainment design were developed by the Mark II Owners Group.

These loads were a result of the Mark II Containment Support Program. This program consisted of a number of experi-mental and analytical tasks that were funded by the Mark II Owners Group.

Program management for this effort was provided by the General Electric Company.

Documentation of this work was provided in the form of reports submitted to the NRC over the course of the program.

Table B-i identifies these documents.

Most of the documents listed in Table B-1 were reviewed by the staff under the TAP A-8 program to establish acceptable LOCA related pool dynamic loads for the Mark II containment design evaluation.

Documentation of other Mark II-related tasks falling outside of the A-8 program; are included in this table.

These other tasks include those associated with the development of safety /

relief valve (SRV) pool dynamic loads and those associated with combining loads. The SRV tasks were reviewed under the NRC A-39 program while load combination reports were reviewed by the NRC as topical reports.

t (h

B-1

TABLE 8-1 1%RE II Pak * * - REPORTS TRANSMITTED TO 4RC TITLE 00CtstENT NO.

SENT BY SEhT 10 LETTER DATE LETTER No.

Mark 11 Containment Forcing Functions NEDE/NED0-21061

1. F. Stuart

8. 5. Boyd 10/24n5 Information Report, Rev. O, 9n5 E,elmation of Effects of Pool Swell in the NEDM-20725-01 I. F. Stuart R. 5. Boyd 1/15n6 8-76 Mark I and II Containennts 11n 4 Mark II Contalrument Forcing Functions MEDE/NEDO-21061 I. F. $tuart

8. S. Boyd 4/9n6 102-76 Information Report, Rev. 1, 9/75 Mark II Pressure S w ression Test Program -

NEDE-13442-P-01 J. F. Quirk W. R. Butler 5/:S/76 172-76.

Phase I Tests, Sn6 m rt I'

• vssure Swpression Test Program -

NEDO-13442 J. F. Quirk W. R. Butler 6Hn6 1t4)-76 Phase 1

~ta 6n 6 Mark II Containment Supporting P; gram NEDO-21297 J. F. Quirk W. R. Butler 6/9#6 195-76 j

Report, Sn 6 f

Mark 11 Phase ! 4T Tests Application J. F. Quirk W. R. Butler 6/14n 6 196-76 Memorandum, 6H 6 Errata (1), Mark II Pressure Suppression Test NECE/NEDO-13442 J. F. Quirk W. R. Butler 6/28/76 227-76 Program - Phase I Tests, 6 H6 Mark II Containment Dynamic Forcing Functions NEDE/NE00-21061 L. J. Sobon R. S. Boyd 9/In6 301-76 cu Information Report, Rev. 2, 9/76 Errata (1), Mark 11 Containment Dynamic Forcing I C I/NEDO-21061 L. J. Sobon R. 5. Boyd 10/14/76 347-76 Functions Information Report,10n6 i

Mark :: Pressure Suppression Test Program -

NEDE-13468-P L. J. Sobon

0. D. Parr 1/4n7 3-77 Phase a and III Tests, 12n 6 N

(Cover Date is 10/76)

Amendment 1. Mark II Containment Dynamic NEDO-21061 L. J. Sobon C. D. Parr 2/2/77 41-77 Forcing Functions Information Report, 12/76 Mark II Pressure Suppression Cortainment System -

NEDE-21544-P L. J. Soboe D. D. Perr 2/10n?

58-77 An Analytical Model of the Pool Swell Phenomenon, 12/76 L. J. Schon O. D. Part 2/24/77 79-U Mark 11 Pressure Supp ession Test Program -

Phase I, II & III 4T Tesss - Appifcation Memo, 1/77

• Caorso Reiter Valve Loads Tests - Test Plan, NEDM-20988 L. J. Soboa S. Miner 3/10/77 91-77 Rev. 2, 12/76

• • Technical Basis for Use of SRSS Method far L. J. Sobon

0. D, Parr 3/30n 7 116-77 Combining Dynamic toads for Mark II Plants.

Interim Report, e n?

Most of the reports in this table were reviewed under the NRC TAP A-8 Program except:

Reports reviewed uader the A-39 Program

• • Reports reviewed outside the A-8 and A-39 Procrams

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TITLE DOCUMENT NO.

SENT BY SENT TO LETTElt DATL LETTER NO.

Camparison of the 1/13 Scale k rk II Containment NED0-21667 L. J. Sobon D. D. Parr 9/23/77 365-77 Multivent Pool Swell Data with Analytical Methods, SH7 "Teciwitcal Basis for Use of SR" Method for NED0-24010 L. J. Sobon

0. D. Parr 9/26n7 364-77 Commining Dynamic Loads for *~.rk II Plants, 7n?

' Mark 11 BWR-5RV Ramshead Buuole Dynamics -

CEM-0394 L. J. Sobon J. F. Stolz 10/19/77 399-77 Analytical Model Comparison with Test Data, 907 Mark II Lead Plant Topical Report: Pool NEDO-23617 L. J. Sobon J. F. Stolz 10/254 7 413-77 Boundary and hin Vent Chugging Loads Justification, 7/77 Results of Pressure Suppression System Tests NEDD-24013 L. J. Sobon J. F. Stolz 10/25/77 413-77 Employed by CE, Series No. 2, 7# 7 Calculation of Spatial Pressure Distribution NED0-24014 L. J. Sobon J. F. Stolz 10/25n7 a13-77

)

l When Several Condensation Events are Superf eposed, 7U7 Load Prediction for tfe Pressure Suppression NEDO-24015 L. J. Sobon J. F. Stolz 10/25/77 413-77 l

l System in Two BWR P1 arts, 7/77 Pressure Oscillation Due to Condensation of NE00-24016 L. J. Sobon J. F. Stolz 10/25/77 413-77 l

Steam in a Water Pool, 7/77 cp j,,

Influence of rack Pressure During Chugging in NEDO-24017 L. J. Sobon J. F. Stolz 10/25/77 413-77 Simulated LOC 4*; at the CE Test Facility, 7/77 Loads on Vent Struts Due to Low m ss Flux Steam NED0-23627 L. J. Sobon J. F. Stolz 10/25/77 413-77 Condensation in a Water Pool, 7n ?

Amenhnt 2, Supplement 2 Nrk 11 Containment NEDO-21061 L. J. Sobon J. F. Stolz 11/3/77 419-77 Dynamic Forcing Functions Information Report.

9H7

(

Single Vent Chugging Model 9/77 NEDE-23703-P L. J. Sobon J. F. Stolz 11/15/77 442-77

• Ramshead SRV Loads - h thodology Suemary 10/77 NEDO-24070 L. J. Sobon J. F. Stolz 11/21U7 445-77

" Preliminary Scaled Multivent Test Program NED0-23697 L. J. Soben J. F. Stolz 1/3/78 1-78 Plan." 12/77

• • Analytical Model for Estimating Drag Forces on NEDO-21471 L. J. Sobon J. F. Stolz 1/27/78 40-78 Rigid Submerged Structures Caused by LOCA and Safety Rollef Valve Ramshead Air Discharges."

9/77 (MK I, MK II)

• " Analytical 14) del for Liquid Jet Properties LEEDE-21472 L. J. Sobon J. F. Stolz 1/26/78 45-78 for Predicting Forces on Rfgid Submerged Structures." 9/77 (MK I, MK II}

l TITLE DOCUNfMT NO.

SENT SY SENT T0 LETTER DATE LETTER 181.

• • • 11 Pressure Suppression Containment Systems -

NEDE-21730-P L. J. Sobon J. F. Stolz 1/31/78 51-78 Loads on Sutserged Structures - Appilcatiocs Memorane s."

12n 7

• • Evaluation of Ef fects on Pool Swell in the MEDM-20725-01 L. J. Sobon J. F. Stolz 2/23/78 81-70 m-I and T'-II Contatnments." 11n4

• Proprietary Appendla to MEDO-21471 S n 7 NEDE-21471-P L. J. Sobon J. F. Stolz 2/24n8 83-78 "The Multivent Hydro # namic Model for Calculating NEDC-21669-P L. J. Sobon J. F. Stolz 3/8n8 99-78 Pool P undary Loads Due to Chugging - NK II Conta..ments." 2/78

'Addenem 1 to Caorso Reitef Valve. cads Tests -

NEDM-20988 L. J. Sobon J. F. Stolz 3/17/78 113-78 Test Plan, NEDM-20988, Revision 2, 10n 1

• Suppression Pool Temperature Response to a NEDC-23689-P L. J. Sobon J. F. Stolz 3/31n8 138-78 Safety / Relief Valve Discharge Through a Ramshead in NK I and II Contalrments." 3n8

• Merk II Containment Swporting Program - Summary NEDE-23710-P L. J. Sobon J. F. Stolz 4/3n 8 130-78 of 4T Fluid Structure Interaction Studies." 4H8 "NK II containment Supporting Program Report."

MEDD-21297 L. J. Sobon J. F. Stolz 4/4n8 137-78 i

i Rev. 1, 3n 8 "MK II Contaf rment Supporting Program Report."

NEDO-21297 L. J. Sobon J. F, Stolz 4/18/78 159-78 7

Rev. 1 3n 8 un

• • Emecutive Sammary Report and 5tumery " Technical L. J. Sobon J. F. Stolz 5/In8 180-78 84 sis for the SRS$ Method af Coe ining Dynaale Response", 4 H8 MK II Containment Program Report, " Single tent NED0-23703 L. J. Sobon J. F. Stolz 5/2n 8 186-78 Model." 9H7

• • Analytical Model for Liquid Jet Properties for NED0-21472 L. J. Sobon J. F. Stolz 5/2R8 187-78 l

Predicting Forces on Algid Structured Structures,"

9/77 L. J. Sobon J. P. Knight 5/5/78 193-78

• • MK II containment Program - Responses to ND*

Request for Additional Information - Load Combinations L. J. Sobon J. F. Stolz 6/19#8 244-78 MK II Containment Request for Additional Information, 6n 8

• • Square Root $4m of the Squares (5R55)

L. J. Sobon J. P. Knight 6/.12n$

246-78 m II Containment Dynamir,Forting Functions NEDO-21061 L. J. Sobon J. F. Stolz 6/3008 262-78 Information Report, Rev. 3, E n8 Proprietary Supplement to MK II Dynamic Forsing NEDE-21061-P L. J. Sobon J. F. Stolz 6/30/78 263-78 Functions Information Report, Rev. 3,Sn8

TITtt DOCUMEmi No_

SENT SY SENT TO LETTER DATE LETTER NO.

s sterk II Pressure 5+pression Test Program.

MEDO-23678 L. J. Sobon J. F. Stolz 6/30n8 272-78 Pt.ases I. II, and III of the 4T lasts -

Appifcation Memo. 6/78 Responses to NRC Request for Additional L. J. Sobon J. F. Stolz 6/30n8 275-78 Information (Round 3 Questions)

MK II r.ontainment Program tetter Report, " Lead L. J. Sobon J. F. Stolz 7n#8 280-78 Plant Containment Response to Improved Chugging Load Definition". 6/78

• " fluid-Structure Interaction Analysis of Rasshead NEDO-23834 L. J. Sobon J. F. Stolz 7B/78 281-78 Safety Relief Valve Discharge in MK I Steel Containment Torus." 6/78

" Chugging Parametric Test Report - Small NEDE-21851-P L. J. Sobon J. F. Stolz 7/12/78 273-78 Scale.* 6H8 "The Multivent Hydredynamic Model for Calculating NEDO-21669 L. J. Sobon J. F. Stolz 7/17#8 293-78 Pool Soundary Loads Due to Chuggi g - NK II Containneats." 2n 8

• • CAOR50 Relief valve Loads Test - Test Plan, NEDM-20988 L. J. Sobon J. F. Stolz 7/1848 296-78 Revision 2 Addendues 1 and 2", 4/78

• 0ynamic Lateral Loads on a Mair Vset Downcuser -

NEDE-24106-P L. J. Sobon J. F. Stolz 7/21#8 296-78 MK II Containment". 3n8 1

4

• • Fluid-Structure Interaction Analysis of Rasshead NEDO-23834 L. J. Sobon J. F. Stolz 7/2dn8 315-78 1

ca Safety Relief Valve Discharge in s4C I Steel Containment Torus" 6n8

• Evaluation of Fluid-Structure Interaction NEDE-21936-P L. J. Soben J.'F. Stolz 7/28/73 316-78 Effects on 8WR MK II Containment Structures". 7R 8

• • 5uppression Pool Teeperature Response to a NEDO-23689 L. J. Sobon J. F. Stolz 8/4/'8 322-78 Safety /Re) ef Valve Discharge Through a Ramshead in MK I and MK II Containments". 3/78

• • MK II Pressure Suppression Containment Systems - NE00-21730 L

J. Sobon J. F. Stolz 8/11/78 339-78 Loads on Submerges Structures - An Appilcation Memorandum", 7# 8

" Chugging Parametric Test Report - Small-Scale "

NEDO-21851 6/78

" Dynamic Lateral Loads on a Main Vent Downcc<ar - NED0-24106 L. J. Sobon J. F. Stolz 9/13n8 363-78 MK II Containment," 3/78

• Evaluation of Fluid-Structure Interaction Effects NEDO-21936 on 8WR rel II Containment " 8/78

""$R55 Application Criteria as Applied to MK II NEDO-24010-1 L.

Sebon J. F. Stolz 11/in8 40J-78 Load Combinations Cases," Supplement 1, 10n 8 (Advarice Coplas)

TITit 00CtMENT NO.

SENT BY SENT TO LETTER DATE LETTER W.

"let II Main Vent Lateral Loads," 10#8 MEDE-23806-P L. J. Sobon J. F. Stolz 11/20n8 417-78

• • • 5855 AppIlcation Lriteria as Apptled to MK II NEDO-24010-1 J. Sobon J. F. Stolz 12/5/78 425-78 Lead Combination Cases," Supplement 1,10/73

• • " Functional Capability Criteria for Essential NEDO-21985 L. J. Sobon J. F. Stolt 12/11/78 436-78 MK II Piping," 9/78

• MK II Containnsnt Supporting Program - Sunnary NEDO-23710 L. J. m J. F. Stolz 12/14n8 440-78 of 4T Fluid Structure Interaction Studies," 9n8

="1/4-Scale Test Report, Loads on Submerged NEDO-23817-P L. J. Sobon J. F. Stolz 12/15n8 441-78 I

Structures Dee to LOCA Air Bubbles and Water i

Jets," 9B S

• • • 8 asis for Criteria for Combination of Earthquake NED0-24010-2 L. J. Sobon J. F. Stolz 1/9n9 8-79 and other Treasier.t Responses by the Square Root Sus of the $@ ares Met.W " Supplement 2,12HS (Advance C Nies)

• MK II M,5G Went Lateral Leeds Summary Report

• MEDO-23806 L. J. Sobon J. F. Stolz 1/2a*-'

t0-79 12n8

• • • Sasis for Criteria for Combination of Earthquake MEDD-24010-2 L. J. Sobon J. F. Stolz 2/$n9 32-79 and Other Transient Responses by the Square Root se of the 4-~ s Method," Supplement 2, 11n8 1

7 Sr. J. E. Borhaug's Presentation, " Responses to L. J. Soboa C. J. And6rson 2/16n9 46-79 N

84L Comments on Dynamic Lateral Loads," 11/17 L. J. Sobon R. J. P.-ttson 2/27n9 55-79

• • LOCA, & 1 SRV Load Combinations p

L. J. Sobon J. F. Stolz 3/16n9 76-79

" Asymmetric LOCA Pool Son.ndary Load for MK II" l

L. J. Sobon J. F. Stolz 3/20n9 80-79

" Vent Clearing Pool Soundary Loads for MK II Plants

• L. J. Sobon J. F. Stolz 3/28/79 90-79

• Caorso SRV Discharge - Summary of Result *, and Preliminary Observations, 3/7S let II Contanment Supporting Program Report, NEDO-23697A L. J. Sobon J. F. Stolz 4/4H9 92-19

" scaled Multivent Test Program Plan - Phase 1 "

1/79

• • 5quare Root Sum of the Squares (5R55), Digitized L. J. Sobon J. P. Kn1 ht 4/30n9 11t-79 2

Time History Data MK II Ieproved Chugging Load Definition.

L. J. Sobon J. F. Stolz 4/30n9 120-79 Task A.16. R<.sponses to Informal Questions

• • 5quare Root Sua cf the Squares (5255) References L. J. Sobon J. P. Knight 5/15/79 136-79

' Technical Description of the Ring Vortex Model L..I Sobon J. F. Stolz 5/22R9 144-79 3/79

TITLE DOCLMENT No.

SENT Sf SENT TO LETTER CATE LETTER Ho.

• • Square Root %m of the Squares (5R55). Digittred L. J. Sobon J. P. Knight 5/29/79 150-79 Tlee History Data L. J. Sobon J. F. Stolz 5/29/79 151-79 NRC Request for A gitional Information, DFFR Rownd 1 Questions NRC Requests 'for Additionai Information DFFR L. J. Sobon J. F. Stolz 5/30/79 152-79 Round 3 Questions L. J. Sobon J. F. Stolz 6/12/79 164-79

• • Square Root Sam of the Squares (5R55) References L. J. Sobon J. P. Knight 6/12/79 165-79

• • 5quare Root Seas of the Square (5255), Table of Contents and Microfiche Copy of Olgttired flee History Data

• *Mt II containment Supporting Program CAOR50 MEDE-25100-P L. J. Sobon J. P. Knight 6/28/79 U6-79 Safety Re11ef Valve Tests Phase 1 Test Report "

5n9 "A Comparison of the MK II Nltivent Model with MEDE-25116-P L. J. Soben J. P. Stolz 7/30n 9 194-79 Prelimis.ary Results free a Scaled Nittwent Test Program," 5/79

• • Preliminary Draft Copies i ' Supplewnt 3 NEDO-24010-0 L. J. Sobon R. Mattu 8/In9 j

i to 5455 Report (Draft) 1 Errate No. 2. "MK II Suppression Test Program -

NEDE-13468-P L. J. Schon J. F. Stolz 8/6/79 203-79 Phase II and III Tests," 10n6 co

• *MK II Contalement Supporting Program CAOR50 NEDO-25100 L. J. Sobon J. F. Stolz 8/13/79 207-79 Safety Relief Valve Tests, Phase I Test Report."

Sn9 "5quare Root sum of the Squares (SR55), Time L. J. Sobon J. P. Knight 8/13n9 208-79 History Response Plots Creare Report TN-297 *Connep Tests, Experfeental TM-297 L. J. Sobon J. F. Stolz 8/15n3 210-79 Data Report *, 6/79

• Mark Il poet Compared with Preliminary Results NEDO-251%

L. J. Sobon J. F. Stolz 10/8/79 253-79 free SM Test Program *, 8n 9

• 5ealed N1tivent Test Program Plan (Phase 2 NED0-23697A L. J. Sobon J. F. Stolz 10/9n9 255-79 Test)*, Rev. 1, Supplement 1, 8n 9 L. J. Sobon

5. H. Hanauer 10/19n9 262-79

" LOCA, + SRY, Load Combination (Load Case 10) p

" Additional Demonstrations of Statistical Basis NED0-24010-3 L. J. Sobon J. F. Sto;z 11/6/79 244-79 for the SRSS Method, Supplement 3, SM S

• • Square Root 54m of the Squares (SRSS)

L. J. Sobon J. P. Knight 1/29/80 025-80 l

l

,y

TITLE DOCUMENT N0.

i'NT BY SENT TO LETTER DATE LETTER NO.

Dr. J. E. Borhoug's Presentation: " Responses R. H. Suchholz C. J. Arulerson 2/29/00 850-80 to SeIL Comments on Dynamic Lateral Loads",

11/17/78, Lebanon, New Hampshire

-" Comparison of $1ngle and Multivent Chugging LEEDE-24!81-1-P R. H. Buchholz J. F. Stolz 3/3/00 051-80 at Two Scales" 1/00

• • *An Analytical Method for Determining the R. H. Suchholz J. F. Stolz 3/21/80 064-80 Probabillcy Distribution of the Maximum of Combined Respt,ases."

Mark II Dyc nic LaterAal tr. ads on Mark II NEDE-24794 P R. H. Buchholz J. F. Stolz 3/31/80 070-80 kin Vert Downcomer--Correlation of Independent Reference Dais."

Method of Applying Mark II Single Vent R. H. Buchholz J. F. Stolz 4/9/80 077-80 Dycanic Laterial Load to Mark II Plant with Multiple Vents

• Assumptions for Usa in Analyzing Mark II R. H. Buchholz J. T. Stolz 4/18/80 071-80 SWR Suppression Pool Temperature Response to Plant Transients Involving Safety /Reiter Valve Discharge R. H. Buchholz J. F. Stolz 4/21/80 081-80

• Response to NUREG-0487 Criteria fr,c Computing Loads on Submerged Structures I

NRC Requests for Additfor.a1 Information, R. H. Suchhe12 K. Knfel 5/29/80 103-80 OFFR Round 2 questions - Proprietary Information NRC Request for Additional Infonnetton, R. H. Buchholz K. En s1 5/29/80 104-80 Offa Round 2 Questions Mark Il containment Program - General Electric NEDE-24822-P R. H. Buchholz K. Kniel 6/6/80 109-00 Report LEEDE-24882-P, " Mark II Isproved Chug 81ng Methodology" -

Mark II Containment Program NEDE-24811-P R. H. Buchholz K. Enlel 6/9/80 110-80 General Electric Report NEDE-24811-P, "4T Condensation Oscillation (4TC0)

. Test Program Final Test Report"

• Mark II Containment Program NEDE-24757-P R. H. Buchholz K. Knfel 7/3/00 118-80 General Electric Report NEDE-24757-P.

"Caerso Safety Rollef Valve Discharge Tests - Phase II Test Report" Mark II Containment Program LEEDE-24811-P R. H. Buchholz K. Kniel 7/8/80 121-80 General Electric Report NEDE-24811-P, "4T Condensation Oscillation (4TCO) 7est Program Final Test Report"

.___.m_-

' ' I

'- - *I

00CtmE"T m0.

SE-T Bf SE~T T0 LETTER DATE LETTER N0.

TITLE Mark II Cutainment Program NEDO-24781-1 R. H. Suchholz K. Kafel 7/8/80 122-80 General Electric Report M D0-24781-1

' Comparison of Single and Multivent Chugging at Two Scales *

• Merk !! Containment frogram NEDE-24798-P R. H Buchholz K. Kafel 7/28/80 130-80 General Electric Report MC c-24798-P.

'Caorso Entended Disch4;ge Test Report"

• Mark II Containment Program NEDO-24757 R. H. Buchholz K. Entel 8/25/80 145-80 General Elect *1c Report NED0-24757

'Caorso Safety 1elief Valve Discharge Tests - Phase !! Test Report" Mark II Containment Program NEDO-24811 R. H. Suchholz K. Kafel 9/3/80 153-80 General Electric poport NEDO-24811 "4T Condensation Osct11ation (4TCO)

Test Program Final Test Report" eMark II Containment Program NEDO-24798 P. H. Suchholz K. Kniel 9/12/80 158-80 General Electric Report NEDO-24798 "Caorso Estended Discharge Test Report" Mark II Containment Program NE00-24822 R. H. Suchholz K. Kniel 9/12/80 160-80 General Electric Report NE00-24822

" Mark 11 Improved Chugging Methodology" 1

Mark II Containment Program NEDO-24794 R. H. Buchholz K. Kniel 9/12/80 159-80 to 0

General Electric Report NED0-24794

" Dynamic Lateral Loads on Mark II Main Vent Downconer - Correlations of Independent Reference Data" Mark II Containment Program NEDE-24822-P R. H. Suchholz K. Kniel 9/12/80 161-80 Errata to General Electric Report NEDE-24822-P. "% rk II Improved Chugging Methoda.ogy" Mark 11 Containment Program NEDE-24794-P R. H. Buchholz K. Knfel 9/17/80 162-80 i

Errata to General Electric Report NEDE-24794-P. " Dynamic Lateral Loads on Mark Il Main Vent Down-comer - Correlation of Inde-pendent Reference Deta" Mark II Containment Program NEDE-25289-1-P R. H. Buchholz K. Knfel 9/17/80 163-80 l

General Electric Report NEDE-25289-1-P. " Comparison of l

Single and N1tivent Chugging -

J Phase 2" Mark 11 Costainment Program NEDO-24822 R. H. Suchholz K. Kafel 10/1/80 165-80 Errata to General Electric Report NEDO-24822. " Mark II Improved Chugging Methodology" l

7 l

l TITtt 00CtmENT mo.

SEuf Of SEmT To TETTER C.TE LETTER 402

' Mark II Contairment Program R. H. Suchholz K. Kniel 10/27/80 144-80 Burns and Roe, Inc. Proprietary Report 'An Analytical Model for LOCA Water Jet in Mart II Conta'rments (The Ring Vorten Model)*

rk II Contaltment Program NEDO-24111

k. H. 'Aachhola K. Knlel 11/13/80 193-80

. rata to Geceral Electric Report MD0-24811,

• 4T Condensation Oscillation (4TCO)

Test Program Final Test Report

• Mrk II Contairment Progras NEDE-24811-P R. H. Buchholz K. Kafel 11/13/80 194-80 Errata to General Electric Report M DE-24811-P. "4T Condensation oscillation (4TC0)

Test Program Final Test Report" i

Mark Il contairment Program NEDE-24288-P R. H. Buchholz K. Katel 11/26/80 205-80 General Electric Repcrt NEDE-24288-P.

" Generic Condernat'on Oscillation Load Deffnttion Report" R. H. Buchholz K. Kafel 12/11/80 188-80

• Mark II Containment Program Surns and Roe, Inc.

Non-Proprietary Report

'An Analytical Model for LOCA Water Jet in Mark !! Containments 7

(The Ring Vorten Model)"

w*

Mark II Containment Program NEDD-25289-1 R. H. Buchholz K. Kniel 12/11/80 213-80 General Electric Report NEDD-25289-1, ' Comparison of Single and Multivent Chugging

• Phasa 2*

Mark II Containment Program NEDE-24288-P R. H. Buchholz K. Kafel 12/11/80 214-80 Proprietary Affidavit for General Electric Report NEDE-24288-P.

' Generic Condensation Oscillation Load Definttfon Report" (Mark II Contairment Progrrg Assimptions for Use in Analyzing Mark !!

R. H. Suchholz Karl Kafel 1/9/81 1-81 BWR Suppresst m Pool Temperature Response to Plant Transtonts Involving Safety / Relief Valve Discharge - Revfston 1 Mark !! Contairment Program Responses to NRC Questions on Mark II Single R. H. Buchholz Karl Kniel 1/16/81 12-81 Vent tateral Loads Mark Il Containment Progree NEDE-24285-P R. H. Buchholz Karl Kniel 2/3/81 17-81 General E!actr i Report NEDE-24285-P,

'4TCO Chugging Data P port - Six Key Runs"

TITLE DOCUhcNT NO, 5,ENT 8Y SENT To LETTER DATE

_ LETTER NO.

eMark II Contairment Program NEDE-25100-P R. H. %chho12 Karl Kniel 2/18/81 20-81 Errata to General Electric Report NEDE-25100-P, "Caorso SRV Discharge Tests, Phase I Test Report"

• Mark 11 Con 6ainment Prcgram NEDO-25100 R. H. Suchho12 Karl Knlel 2/18/81 21-81 Errata to General Electric Report NEDO-25100, "Caorso SRV Discharge Tests, Phase I Test Repert" Mark II Containment Program NED3-24288 R. H. Buchho12 Karl Kniel 3/6/81 36-81 General Electric Report NEDO-242no,

" Generic Condensation Oscillation to e Definition Report

• s h rk 11 Containment Program R. H. Buchholz Karl Kniel 3/6/81 38-81 Quencher Condensation Performance krk II Containment Program NEDO-24285 R. H. Buchholz Karl Kniel 3/25/81 59-81 General Electric Report NEDO-24285 "4TCO Chugging Data Report - Six Key Runs" eMark II Cont 41nnent Program NEDE-24835-P R. H. 8uchho12 Karl Kniel 3/31/81 59-81 General Elec tric Report NEDE-24835-P, "Caorso 3afe'y Rollef Valve Tests -

REVAL0lt L. L Comparison" cp

,L Mark II Containment Program NEDE-24302-P R. H. Buchho12 Karl Knlel 4/27/81 81-81 N

General Electric Report NEDE-24302-P.

"Gensric Chugging Load Definition Report" (Advance Copy) b rk II containment Program NEDE-24300-P R. H. Buchholz Karl Kniel 5/1/81 75-81 General Electric Report NEDE-24300-P, " Scaled Multivent Chugging Test Progra.a"

• Mark II Containment Program R. H. Buchholz Karl Kniel 5/12/81 84-81 Nthods for Calculating Mass and Energy Release for Suppression Pool Temperature Response to Safety Relief Valve Discharges k rk II Containment Program NEDE-24302-P R. H. Buchholz Karl Kniel 6/2/81 105-81 General Electric Report NEDE-24302-P.

" Generic Chu9ging Load Deftaltion Report" Mark II Containment Program R. H. Suchho12 Karl Knlel 6/5/81 109-81 Main Vent lateral Loads S. J. Stark Karl Knlel 6/30/81 125-81 Mark II Containment Program Differential Pressure IJpload on Diaphragm Floor I

APPENDIX C SUPMARY OF MARK II LOCA-RELATED P0OL DYNAMIC LOADS The Mark II program to establish LOCA-related pool dynamic loads has been in existence since April 1975.

Since that time, a number of different load specifications have been developed.

The purpose of this appendix is to identify, in one location, those generic load specifications that the staff finds acceptable.

A summary of generic loads acceptaSle to the NRC is provided in Table C-1.

This table includes the following information:

load identification, a summary of the load specification, reference to the Mark II Owners Group document providing(where applicable)ption of the load specification, NRC acceptance a detailed descri criteria

, and reference to the NRC NUREG section that describes the NRC specific load evaluation.

The staff finds most of the generic i.0CA-related ?ool dynamic load specifications proposed by the Mark II owners acceptable.

For t1e few cases where the staff was unable to conclude that a proposed load was acceptable the staff developed acceptance criteria.

Thecriteriaprovideloadspecificatlonsthatareaccept-able to the staff.

The staff finds that the detailed loads specifications referenced in Table C-1, along with the acceptance criteria provided in Appendix A of this report, constitute a complete set of acceptable LOCA-related pool dynamic loads.

These load specifications complete the TAP A-8 program.

l l

l C-1

TABLE C-1 MARK II LOCA-RELATED HYDRODYNAMIC LOADS SIM ARY TABLE

~'

MARK II 0.G.

DETAILED LOAD NRC ACCEPTANCE NRC LOAD OR PHENOMENON LOAD SPECIFICATION DESCRIPTION CRITERIA EVALUATION (2)*II.A.1 A.

Submerged Boundary Loads 24 psi overpressure added Letter No. 80-79 During Vent Clearing to local hydrostatic below from GE to NRC vent exit (walls and base-dated 3/20/79 mat) - linear attenuation to pool surface.

B.

Pool-Swell Loads 1.

Pool-Swell Analyti-cal Model 7

a) Air-Bubble Calculated by the pool-NEDE/NEDO-21061 (1)III.B.3.a.1 Pressure swell analytical model NEDE-21544-P (PSAM) used in calculation of submerged boundary loads.

b) Pool-Sweil Use PSAM with polytropic Letter No. 360 A.1 (2)II.A.2 Elevation exponent of 1.2 to a max-from LILCO to imum swell height which is NRC dated 2/16/79 the greater of 1.5 vent submergence or the eleva-tion corresponding to the drywell floor uplift AP=2.5 psid.

1) - Reference NUREG-0487 1-ReferenceNUREG-0487 Supplement 1

- Reference NUREG-0808

MARK II 0.G.

DETAILED LOAD NRC ACCEPTANCE NRC LOAD OR PHENOMENON LOAD SPECIFICATION DESCRIPTION CRITERIA EVALUATION c) Pool-Swell Velocity histon> vs. pool NECE/NE00-21061 A.2 (1)III.B.3.a.3 Velocity elevation predicted by the NEDE-21544-P PSAM used to ccmpute impact loading on small structures and drag on gratings between initial pool surface and maximum pool elevation and steady-state drag between vent exit and maximum pool elevation.

Analytical velocity variation is used up to maximum velocity.

Maximum velocity applies thereafter up to maximum pool swell.

PSAM predicted n

4, velocities multiplied by a factor of 1.1.

d) Pool-Swell Acceleration predicted by NEDE/NEDO-21061 (1)III.B.3.a.4 Acceleration the PSAM.

Pool accelera-NEDE-21544-P tion is utilized in the calculation of accelera-tion loads on submerged components during pool swell.

e) Wetwell Air Wetwell air compression Letter No. 360 (2)II.A.2 Compression is calculated by PSAM con-from LILCO to sistent with maximum pool NRC dated 2/16/79 swell elevation in B.1.b above.

MARK II 0.G.

DETAILED LOAD NRC ACCEPTANCE NRC LOAD OR PHENOMENON LOAD SPECIFICATION DESCRIPTION CRITERIA EVALUATION f) Drywell Metleds of NEDM-10320 and Letter No. 082 (1)lII.B.3.a.6 Pressure HEDO-20533 Aapendix B.

from LILC0 to Utilized in

) SAM to cal-NRC dated 6/11/81 culate pool swell loads.

NEDM-10320 NED0-20533 App. B 2.

Loads on Submerged Maximum bubble pressure NEDE/NE00-21061 (1)III.B.3.b Boundaries predicted by the PSAM NEDE-21544-P added uniformly to local hydrostatic below vent exit (walls and basemat) linear attenuation to pool surface.

Applied to walls up to maximum pool swell elevation, n

A 3.

Impact Loads a) Small 1.35 x Pressure-Velocity A.5 (1)III.B.3.c.1 Structures correlation for pipes and I beams based on PSTF impulse data and flat pool assumption.

Variable pulse duration.

b) Large None - Plant unique load (1)III.B.3.c.6 Structures where applicable.

c) Grating P drag vs. grating area NEDE/NEDO-21061 A.3 (1)III.B.3.c.3 correlation and pool velocity vs. elevation.

Pool velocity from the PSAM.

P drag multiplied by dynamic load factor.

MARK II c.G.

DETAILED LOAD NRC ACCEPTANCE NRC LOAD OR PHENOMENON LOAD SPECIFICATION DESCRIPTION CRITERIA EVALUATION 4.

Wetwell Air Compression a) Wall Loads Direct application of the NEDE/NEDO-21061 (1)III.B.3.d.1 PSAM calculated pressure due to wetwell compression.

(3)2.1.2.7 b) Diaphregm 5.5 psid for diaphragm Letter No. MFN-125-81 Upward Loads loadings only.

dated 6/30/81 5.

Asymmetric LOCA Use 20 percent of maximum Letter No. 76 A.4 (2)II.A.3 Pool bubble pressure statis-from GE to NRC tically applied to 1/2 of dated 3/16/7f the submerged boundary.

F C.

Steam Condensation and Chugging Loads 1.

Downcomer Lateral Loads a) Single-Vent Dynamic load to end of NEDE-23806-P B.1.a (3)2.3.3.2 Loads (24 in.)

vent.

Half sine wave with a duration of 3 to 6 ms and corresponding maximum amplitudes of 65 to 10 K1bf.

Prescribed variation of Letter No. 077-80 b)

Multip(le-Vent (3)2.3.3.3 Loads 24 in.)

load per vent vs. number of from GE to NRC vents.

Determined from dated 4/9/80 single vent dynamic load specification and multivent reduction factor.

c) Single / Multiple Multiply basic vent Letter No. MFN-012-81 B.1.b (3) 2.3.2.1 vent loads loads by factor f=1.34 from GE to NRC (28 in.)

dated 1/16/81

l l

l l

i D

LED bAb NRC ACCEPTANCE NRC l

LOAD OR PHENOMENON LOAD SPECIFICATION DESCRIPTION CRITERIA EVALUATION 2.

Suonerged Boundary Loads (3)2.2.1.3 a) High/ Medium Bounding C0 pressure NEDE-24286 P Steam Flux Con-histories observed in densation 4TCO tests.

Inphase Oscillation application.

Load (3)2.2.2.3

<)

Low Steam Flux Conservative set of NEDE-24302-P Chugging Load 10 sources derived from NEDE-24822-P 4TCO tests.

Applied to Letter No. MKII-085-LIC plants 7 using the from LILCO to NRC IWEGS/ MARS acoustic model.

dated 7/14/81 source desynchronization c3 f,

of 50 as or alternate load using 7 sources derived from the 4TCO key chugs without averaging.

- Symmetric All vents utilize source Load of equal strength for each of the sources.

- Asymmetric Source strengths Si =

Load Case S (lia) applied to all vents on + and - side of containment.

Sources based on the symmetric sources.

Asymmetric parameter a based on ras moment method of interpreting experimental 4TCO single-vent and JAERI multivent data.

MARK II 0.G.

DETAILED LOAD NRC ACCEPTANCE NRC LOAD OR PHENOMENON LOAD SPECIFICATION DESCRIPTION CRITERIA EVALUATION.

D.

Secondary Loads (1)III.E.1 1.

Sonic Wave Load Negligible Load NEDE/NEDO-21061 (1)III.E.2 2.

Compressive Wave Negligible Load NEDE/NEDO-21061 Load 3.

Fallback Load on Negligible Load NEDE/NEDO-21061 (1)III.E.5 Submerged Boundary

~

4.

Thrust Loads Momentum balance NEDE/NED0-:"51 (1)I I.E.6 5.

Friction Drag Standard friction drag NEDE/NED0-21061 (1)lII.E.7 Loads on Vents calculations 6.

Vent Clearing Negligible Load NEDE/NEDU-21061 (1)III.E.8 nL Loads

U.S. NUCLEAR REGULATORY COMMISSIOd 7

BIBLIOGRAPHIC DATA SHEET NUREG-0808

4. TITLE AND SU8 TITLE (Add Vobine Na. //sparsenew/

2. (Leave 6/mik/

" Mark II Containment Program Evaluation and Acceptance Criteria

7. AUTHORIS)

6. DATE REPORT COMPLETED Clifford J. Anderson and Members of the A-8 Task Force My'"

D'1981

^"

9. PERFORMING ORGANIZATION NAME AND MAILING ADDRESS (Incase I, Com/

DATE REPORT ISSUED U. S. Nuclear Regulatory Commission IT981 Au ust Division of Safety Technolog'y Washington, D. C.

20555

e. a.eae umk1

12. SPONSORING ORGANIZATION NAME AND MAILING ADDRESS (incke le Co*/

0. PROJECT / TASK / WORK UNIT NO.

U. S. Nuclear Regulatory Commission H. CONTRACT NO.

Division of Safety Technology Washington, D. C.

20555

13. TYPE OF REPORT PE RIOD COVE RED (Inclusive damsf Regulatory Report

15. SUPPLEMENTARY NOTES

14. (Leave um>&f

16. ABSTRACT 000 words or less)

This report, prepared by the staff of the Office of Nuclear Reactor Regulation, provides a discussion of LOCA-related suppression pool hydrodynamic load; in boiling water redctor (BWR) facilities with the Mark II pressure-suppression containment desigr This report concludes NRC Generic Technical Activity A-S, " Mark II Containment Pool Dynamic Loads," which has been designated an " Unresolved Safety Issue" pursuant to Section 210 of the Energy Reorganization Act of 1974.

On the basis of large-scale tests conducted in 1979, the Mark II Owners develor,ed improved condensation-oscillation and chugging loads for the suppression-pool boundary and lateral loads for the containment downcomers. The staff has reviewed these pro-posed loads and ccncluded that, with a few specified changas, the:e loads provide con-servative loading conditions.

In addition, the staff has conducted a study which confirms that the lead plant pool swell loads, adopted by the Mark II Owners as the final load specifications, are conservative. This study used the results of full-scale multivent Mark II tests conducted in Japan. The staff acceptance criteria for pool swell loads from the Lead Plant Program and new criteria for steam loads developed in the Long Term Program have been consolidated and. constitute Appendix A of this report.

17. KEY WORDS AND DOCUMENT ANALYSIS 17st DESCRIPTORS Mark II long Tenn Program Mark II Acceptance Criteria Pool Dvnanic loads Pressure Suopression Containment 17tk IDENTIFtE RS/OPENf NDE D TERMS

18. AVAILABILITY STATEMENT 9

i C SS (7% report /

21. NO. OF PAGES Unlimited

20. SECURITY CLASS (T4 pese)

22. PRICE linrl m i fi orf S

wnc comu ass tv.77)

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