ML19322A061

(^^ TABLE OF CONTENTS Page ABSTRACT 111 LIST OF ACRONYMS xv

1.0 INTRODUCTION

1-1 1.1 Description of the Mark I Containment 1-1 1.2 Mark I Containment Program 1-2 1.3 Load Definition Report - Objective 1-5 2.0 REVIEW OF PHENOMENA 2-1 2.1 Design Basis Accident 2-1 2.2 Intermediate Break Accident 2-4 2.3 Small Break Accident 2-6 2.4 Safety / Relief Valve Actuation '-7

(, 2.5 other Considerations 2-8 3.0 LOAD COMBINATIONS 3-1 4.0 LOCA RELATED LOADS 4.0-1 4.1 Containment System Temperature and Pressure Response 4.1.1-1 4.1.1 Design Basis Accident 4.1.1-1 4.1.2 Intermediate Break Accident 4.1.2-1

.4.1.3 Small Break Accident 4.1.3-1 4.2 Vent System Thrust Loads 4.2-1 4.2.1 Analytical Procedure 4.2.1-1 4.2.2 Assumptions 4.2.2-1 4.2.3 Analysis Results 4.2.3-1 4.2.4 Application 4.2.4-1 4.3 Pool Swell Loads 4.3-1 4.3.1 Torus Net Vertical Load Histories 4.3.1-1 4.3.2 Torus Shell Pressure Histories 4.3.2-1 A

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NEDO-21888 TABLE OF CONTENTS (Continued)

Page 4.3.3 Vent System Impact and Drag 4.3.3-1 4.3.4 Impact and Drag on Other Structures Above the Pool 4.3.4-1 4.3.5 Froth 1mpingement Loads 4.3.5-1 4.3.6 Pool Fallback Loads 4.3.6-1 4.3.7 LOCA Jet Load 4.3.7-1 4.3.8 LOCA Bubble - Induced Drag Loads on Submerged Structures .

4.3.8-1 4.3.9 Vent Header Deflector Loads (Later) 4.3.9 4.4 Condensat. ion oscillaton Loads 4.4-1 4.4.1 Torus Shell Loads (Later) 4.4-1 4.4.2 Loads on Submerged Structures (Later) 4.4-1 4.4.3 Lateral Loads on Downcomers (Later) 4.4-1 4.4.4 Vent System Loads (Later) 4.4-1 4.5 Chuggir.g Loads 4.5-1 4.5.1 Torus Shell Loads (Later) 4.5-1 4.5.2 Loads on Submerged Structures (Later) 4.5-1 4.5.3 Lateral Loads on Dovncomers (Later) 4.5-1 4.5.4 Vent System Loads (Later) 4.5-1 5.0 SAFETY RELIEF VALVE DIfCHARGE LOADS 5.0-1 5.1 Introduction 5.1-1 5.2 T-Quencher Loads 5.2.1-1 5.2.1 S/RV Discharge Line Clearing Transient Loads 5.2.1-1 5.2.2 Torus Shell Pressure 5.2.2-1 5.2.3 S/RVDL Reflood Transient 5.2.3-1 5.2.4 T-Quencher Water Jet Loads on Submerged Structures 5.2.4-1 5.2.5 T-Quencher Bubble-Induced Drag Loads on Submerged Structures 5.7 5-1 5.2.6 Thrust Loads on T-Quencher Arms 5.2.6-1 5.2.7 Maximum S/RVDL and Discharge Device Pira Wall Temperature 5.2.7-1 5.2.8 Fatigue Cycles 5.2.8-1 O

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' 'T TABLE CF CONTENTS (Continued)

Page 5.3 Ramshead Loads 5.3-1 5.3.1 S/RV Discharge Line Clearing Transient Loads 5.3.1-1 5.3.2 Ramshead Torus Shell Pressures 5.3.2-1 5.3.3 S/RVDL Reflood Transient (Ramshead) 5.3.3-1 5.3.4 Ramshead Water Jet Loads on Submerged Structures 5.3.4-1 5.3.5 Ramshead Bubble-Induced Drag Load on Submerged Structures 5.3.5-1 5.3.6 Maximum S/RVDL and Discharge Device Wall Temperature 5.3.6-1 5.3.7 S/RVDL Fatigue 5.3.7-1 6.0 OTHER CONSIDERATIONS 6.0-1 6.1 Seismic Slosh Loads 6.1-1 ,

6.2 Post Pool Swell Waves 6.2-1 6.3 Asymmetric Vent Performance 6.3-1 rx 6.4 Downcomer Air Clearing Lateral Load 6.4-1 f

\- 6.5 Sonic Wave 6.5-1 6.6 Compressive Wave 6.6-1 6.7 Safety / Relief Valve Steam Condensation Loads 6.7.1-1 6.7.1 T-Quencher Discharge 6.7.1-1 6.7.2 Ramshead Discharge 6.7.2-1

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<w LIST OF ILLUSTRATIONS

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Figure Title Page 1.1-1 Typical Mark 1 Containment 1-8 1.1-2 Typical Composite Section Through Suppression Chamber 1-9 1.1-3 Typical Mark I Plant S/RV Discharge Line Configuration 1-10 1.1-4 Discharge Devices Employed in Mark 1 Plants 1-11 3.01 Loading Condition Combinations. Structures Affected:

Vent lleader, Main Vents. Downcomers and Torus Shell, Accident condition: DBA 3-6 3.0-2 Loading Condition Combinations, Structures Affected:

Vent Ileader, Main Vents, Downcomers Torus Shell and Submergei Structures, Accident Condition: IBA 3-7 3.0-3 Loading Condition Combinations. Structures Affected:

Vent lleader, Main Vents, Downcomers, Torus Shell and Submerged Structures, Accident Condition: SBA 3-8 1.0-4 Loading Condition Combinations, Structures Affected:

Small Structures Above Pool, Accident Condition: DBA 39 3.0-5 Loading Condition Combinations, Structures Affected:

Submerged Structures, Accident Conditions: DBA 3-10 4.1.1-1 Typical Containment Temperature Response to Design

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Basis Accident (Recirculation Line Break) 4.1.1-11

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\_ ,/ 4.1.1-2 Typical Containment Pressure Response to Design Basis Accident (Recirculation Line Break) 4.1.1-12

'. 1.1-3 Drywell Pressure Response to Main Steam Line Break and Recirculation Line Break 4.1.1-13 4.1.2-1 Typical Containment Temperature Response to an Intermediate Break Accident 4 .1. 2-5 4.1.2-2 Typical Containment Pressure Response to an Intermediate Break Accident 4.1.2-6 4.1.3-1 Typical Containment Temperature Response to a Small Break Accident 4.1.3-5 4.1.3-2 Typical Containment Pressure Response to a Small Break Accident 4.1.3-6 4.2-1 Definition of Positive Thrust Loads 4.2.4-5 4.2-2 Single Main Vent Forces (0-5 see) 4.2.4-6 4.2-3 Vent lleader Forces Per Mitre Bend (0-5 see) 4.2.4-7 4.2-4 Single Downcomer Forces (0-5 sec) 4.2.4-8 4.2-5 Total Vertical Forces. Net Vertical Force (0-5 sec) 4.2.4-9 4.2-6 Single Main Vent Forces (0-30 see) 4.2.4-10 4.2-7 Vent IIcader Forces Per Mitre Bend (0-30 see) 4.2.4-11 O

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NEDO-21888 LIST OF ILLUSTRATIONS (Continued)

Figure Title _Page 4.2-8 Single Downcomer Forces (0-30 sec) 4.2.4-12

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Total Vertical Forces, Net Vertical Force (0-30 sec) 4.2.4-13 4.2-10 Pressure Time Histories (0-5 sec) 4.2.4-14 4.2-11 Pressure Time Histories (0-30 sec) 4.2.4-15 4.2-12 Application of Thrust Force on Main Vent End Cap and Main Vent Mitre Bend 4.2.4-16 4.2-13 Application of Vent Header Forces 4.2.4-17 4.2-14 Application of Downcomer Forces 4.2.4-18 4.3.1-1 Typical Net Torus Vertical Loading History 4.3.1-9 4.3.1-2 Comparison of Typical Analytical and Test Drywell Pressure Histories 4.3.1-10 4.3.1-3 Enthalpy Flow Comparison 4.3.1-11 4.3.2-1 Typical Average Submerged Torus Pressure History 4.3.2-6 4.3.2-2 Typical Torus Airspace Pressure History 4.3.2-7 4.3.2-3 Submerged Location on Typical Torus Shell 4.3.2-8 4.3.3-1 Vent System Coordinates 4.3.3-8 g

4.3.3-2 Downcomer Impact and Drag Pressure Transient 4.3.3-9 4.3.3-3 Application of Impact / Drag Pressure Transient to Downcomer 4.3.3-10 4.1.3-4 Vent Header Local Impact Pressure Transient 4.3.3-11 4.3.4-1 Typical Mark I Torus Sector 4.3.4-12 4.3.4-2 Typical Pool Surface Displacement Longit ud inal Distribution 4.3.4-13 4.3.4-3 Typical Plant Pool Surface Displacement in XY Plane 4.3.4-14 4 . 3 . 4 -/- Typical Pool Surface Velocity Longitudinal Distribution 4.3.4-15 4.3.4-5 Typical Plant Velocity Transient in XY Plane 4.3.4-16 4.3.4-6 Calculation of Distance Over Which Impulse Acts 4.3.4-17 4.3.5-1 Definiti.on of Froth Impingement Region I 4.3.5-7

4. 3.5- 2 7efinition of Froth Impingement Region II 4.3.5-8 4.3.5-3 rroth Loading History - Region I 4.3.5-9 4.3.5-4 Froth Loading History - Region II 4.3.5-10 4.3.5-5 Froth Impingement Region II - Possible Directions of Load Application 4.3.5-11 4.3.5-6 Possible Directions of Froth Fallback Load Application 4.3.5-12 O

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Figure Title Page

. 4'.3.6-1 Possible Directions of Fallback Load Application 4.3.6-3 .

i i 4.3.8-1 Torus and Typical Submerged Structure Geometry 4.3.8-5 ;

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' Sample History of Total X-Forces on Vent Header Support 4.3.8-2 Column 4.3.8-6 4

4.3.8-3 Sample Timo History of Total Z-Forces on Vent Header Support Column 4.3.8-7 j 5.1-1 T-Qencher' Load Definition Scheme 5.1-3 5.1-2 Ramshead Load Definition Scheme 5.1-4 5.2.1-7 5.2.1-1 Sample Prediction of S/RVDL Internal Pressure Transient 5.2.1-2 Sample Prediction of T-Quencher Internal Pressure Transient 5.2.1-8 5.2.1-3 Sample Prediction of Thrust Loading on an S/RV Pipe Segment Initially Filled with Gas 5.2.1-9 I' 5.2.1-4 Sample Prediction of Thrust en S/RV Pipe Run Between the !

Discharge Device and the First Upstream Elbow (Pipe Run Initially Filled with Water) 5.2.1-10 l

5.2.1-5 Sense of Thrust Loading 5.2.1-11

$- 5.2.1-6 Sample Prediction of Mass Flow Rate of Water Exiting '

T-Quencher 5.2.1-12 5.2.1-7 Sample Prediction of Water Mass Acceleration 5.2.1-13

,' 5.2.2-1 Sample Prediction of Torus Shell Pressure Loading Transient 5.2.2-5

~5.2.2-2 Sample Prediction of Torus Shell Longitudinal Pressure Distribution 5.2.2-6

, 5.2.2-3 Sample Prediction of Torus Shell Radial Pressure Distribu-

tion at Section A-A in Figure 5.2.2-2 5.2.2-7

5.2.3-1 Sample Prediction of S/RVDL Reflood Transient 5.2.3-5 5.2.4-1 . Phases of Quencher Jet Formation and Decay 5.2.4-5 ,

5.2.4-2 Jet Sections Along the Quencher Arm on Torus Plan View 5.2.4-6

5.2.5-1 _ Sample Predicted Time History of. Total X-Forces on Downcomer 5.2.5-5 5.2.5-2 Sample Predicted Time-History of Total Z-Forces on Downcemer 5.2.5-6 5.2.6-1 . Thrust Loads on Arm Endcaps 5.2.6-7 5.2.6-2 . Generalized Shape of Thrust Loading Transient and Application Point with Time 5.2.6-8 t

5.2.7-1 Example of Predicted S/RVDL and Discharge Device '

Temperature Distribution 5.2.7-3 5.3.1-1 Sample Prediction of the S/RVDL Internal Pressure Transient 5.3.1-5 5.3.1-2 Sample Prediction of Thrust Load Transient at a Typical S/RVDL Segment 5.3.1-6

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NEDO-21888 LIST OF ILLl'STRATIONS (Continued)

O Figure Title Page 5.3.1-4 Sample Prediction of Water Clearing Thrust Lead on the Last Pipe Segment of the S/RVDL 5.3.1-8 5.3.1-5 Sample Prediction of Mass Flow Rate of Water Exiting Ramshead 5.3.1-9 5.3.1-6 Sample Preiiction of Water Mass Acceleration During S/RV Dischuge 5.3.1.10 5.3.2-1 Normal Shell Pressure Transient for Any Location on the Pool Wall Due to an S/RV Discharge for a Cold Pipe (Monticello Test Data) 5.3.2-7 5.3.2-2 Normalized Positive Pressure Distribution at the Torus Cross-Section Coincident with the Ramshead CL 5.3.2-8 5.3.2-3 Sample Prediction of Torus Shell Longitudinal Pressure Distribution Resulting from an S/RV Discharge through a Ramshead Nischarge Device 5.3.2-9 5.3.2-4 Procedural Flowchart 5.3.2-10 5.3.2-5 Normalized Shell Pressure Transient for any Location on the Pool Shell Due to an S/RV Discharge - Leaky Valve (Monticello Test Data 5.3.2-11 5.3.2-6 Normalized Shell Pressure for any Location on the Pool Shell Due to a Subsequent Valve Actuation - NWL (Monticello Test Data) 5.3.2-12 5.3.4-1 Sample Prediction of a Typical Structure Loading History 5.3.4-5 5.3.5-1 Sample Predicted Time History of Total X-Forces on Downcomer and Vent Header Support Column 5.3.5-5 5.3.5-2 Sample Predicted Time History of Total Z-Forces on Downcomer and Vent Header Support Column 5.3.5-6 6.1-1 Application of Seismic Slosh Drag Load 6.1-4 O

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(} LIST OF TABLES V

Table Title PaFe 1.2-1 Mark 1 Utilities and Plants in the Mark I Owners Group 1-7 3.0-1 S/RV Discharge Load Cases foe Mark I Structural Analysis 3-3 3.0-2 Event Timing Nomenclature 3-4 3.0-3 Structural Loading 3-5 4.1.1-1 Fuel Decay Heat and Sensible Energy for All Mark I P1' ants 4.1.1-7 4.1.1-2 Plant Vent System Loss Coefficients 4.1.1-8 4.1.1-3 Typical Plant Conditions at Instant of DBA Pipe Break 4.1.1-9 4.1.1-4 Summary of !! ark I Plant Break Areas 4.1.1-10 4.2-1 Nomenclature for Section 4.2 4.2.4-3 4.3.1-1 Mean Peak Net Torus Vertical Loads Operating Drywell/

Wetwell Pressure Differential (aP) 4.3.1-7 4.3.1-2 Single Test Peak Net Torus Vertical Peak Loads Zero Drywell/Wetwell Pressure Differential 4.3.1-8 4.3.2-1 Torus Shell Pressure History Multipliers M3 and M9 4.3.2-5 4.3.3-1 Internal Vent Header / Torus Airspace Pressure Differential

,f s at Time of Vent Header Impact 4.3.3-7

\- / 4.3.4-1 Hydrodynamic Mass and Acceleration Drag Volumes for Two-Dimensional Structural Components 4.3.4-9 6.1-1 }& I Containment Slosh Results 6.1-3 l

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! LIST OF ACRONYMS i=

A/E- Architect / Engineer l-ABS Absolute Summation Method ADS Automatic Depressurization I BWR Boiling Water Reactor DBA _ Design Basis Accident ECCS Emergency Core Cooling Systems EPRI Electric Power Research Institute FSAR Final Safety Analysis Report FSTF Full Scale Test Facility FUCI j

HEM Homogeneous Equilibrium Model 1 - HPCI High Pressure Coolant Injection j

f F IBA Intermediate Break Accident j LDR Load Definition Report

[ LOCA Loss of Coolant Accident LPCI Low Pressure Coolant Injection LTP Long Term Program

' MSIV- Main Steam Isolation Valve MSLB Main Steam Line Break MWT Megawatt Thermal i

I NOC Normal Operating Conditions NRC Nuclear Regulatory Commission 4

NSSS'- ' Nuclear-Steam Supply System NWL . Normal Water Level xv Revision 0-

NEDO-21888 LIST OF ACRONYMS (Continued)

PULD Plant Unique Load Definitions g QSTF Quarter Scale Test Facility RCIC Reactor Core Isolation Cooling RHR Residual Heat Removal RLB Recirculation Line Break RPC Reactor Pressurs Vessel SBA Small Break Accident S/RV Safety / Relief Valve S/RVDL Safety / Relief Valve Discharge Line STP Short Term Program SRSS Square Root Sum of the Squares O

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1.0 INTRODUCTION

- In June of 1976 at the request of the Mark I owners Group, General Electric Company initiated a program to provide the domestic Mark I utilities with detailed containment load definitions for specific hydrodynamic events appli-cable to their plants. More precisely, Mark I containment pressure suppression component load definitions were to be provided for the postulated Loss-of-Coolant' Accident (LOCA) and Safety / Relief Valve (S/RV) actuation events. This report represents the culmination of these activities and provides the hydro-dynamic load definition procedures to be utilized by each utility in a plant unique structural evaluation. Complete load definitions for the components

! discussed herein are obtained through consideration of these loads in conjunc-tion with the seismic, deadweight and buoyancy effects.

Tae loads and methodologies given in Sections 4, 5 and 6 of this report are defined in sufficient detail that a plant unique analysis can be performed I for each Mark I containment. The loads are presented on a generic basis except where geometric or system differences require that the loads be J O' define'd on a plant uni'que basis. In cases where it has been more efficient to do so, the method to calculate a plant unique load, rather than the load itself, is presented. For each loading condition, the analytical and/or experimental bases for the stated definition are also presented. A complete list of acronyms utilized in this report is given in the preface.

1.1 DESCRIPTION

OF THE MARK I CONTAINMENT The Mark I containment is a pressure suppression system which houses the Boiling Water Reactor (BWR) pressure vessel, the reactor coolant recirculating loops and other branch connections of the Nuclear Steam Supply System (NSSS).

It consists of a drywell, a pressure suppression chamber (or wetwell) which is approximately half filled with water, and a vent system which connects the

, drywell to the wetwell suppression pool. The suppression chamber is toroidal in shape andLis located below and encircling the drywell. The drywell-to-wetwell vents are connected to a vent header contained _within the airspace of the wetwell. Downcomer pipes project downward from the vent header and termi-

-(f nate below the water surface of the suppression pool. The pressure suppression' 4

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NED0-21888 chamber is shown in relation to the steel drywell in Figure 1.1-1. Figure 1.1-2 shows a typical cross section of the suppression chamber.

In the highly unlikely event of a high energy NSSS piping failure within the drywell, reactor water and/or steam would be released into the drywell atmos-phere. This postulated event is referred to as a Loss of Coolant Accident (LOCA). As a result of increasing drywell pressure, a mixture of drywell atmosphere, steam, and water would be forced through the vent system into the pool of water maintained in the suppression chamber. The steam vapor would condense in the suppression pool, thereby limiting internal containment pressure. The noncondensible drywell atmosphere would be transferred to the suppression chamber and contained therein. Section 2 of this report presents a complete description of these events.

B'4R's utilize saf ety/ relief valves (S/RV) attached to the main steam line as a means of primary system overpressure protection. The outlets of these valves are connected to discharge pipes which are routed to the suppression pool as shown in Figures 1.1-2 and -3. The discharge lines end in either Ramshead or T-Quencher type discharge devices. Figure 1.1-4 presents the configuration lll of each type of discharge device.

Follewing the actuation of an S/RV, steam enters the safety / relief valve discharge line (S/RVDL), pressurizing the line and forcing the air and water initially in the line to be expelled to the wetvell and contained therein.

Subsequent to water and air clearing of the S/RVDL, the steam released by the S/RV actuation flows through the S/RVDL and discharge device into the suppression pool and is condensed therein. Section 2 of this report presents a complete description of these events.

1.2 MARr; I CO:;TAINMENT PROGRAM For the Mark I containment design, the pressure and temperature loads associated with a LOCA, were based on experimental technology obtained from testing of the 9

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Bodega Bay and Humboldt Bay Power Plant pressure suppression concepts.

The tests were performed to demonstrate the viability of the pressure sup-pression. concept for reactor containment design by simulating the LOCA with various equivalent piping break sizes up to twice the cross-sectional break area of the largest reactor system pipe. The test data provided quantitative information for establishing containment us,ign pressures.

In performing large scale testing of an advanced design pressure-suppression containment (Mark III), ard during in-plant testing of Mark I containments, suppression pool hydrodynamic loads not explicitly included in the original Mark I containment design basis were identified. These additional loads could result from dynamic effects of drywell air and steam being rapidly forced into the suppression pool during a postulated LOCA, and from suppres-sion pool response to S/RV operation generally associated with plant transient operating conditions. Since these hydrodynamic loads were not explicitly considered in the original design of the Mark I containment, the NRC staff in early 1975 requested a detailed reevaluation of the containment system from

() each domestic utility with a Mark I containment.

Recognizing the joint need to respond to the NRC requests for additional information and the essential similarity of all the Mark I plants, the domestic Mark I utilities formed an Owners Group on April 23, 1975. The Owners Group provided a strong, unified, and consistent approach to resolution of the open issues through the pooling of individual resources. The member utilities of the Mark I Owners Group and their respective plants are listed in Table 1.2-1.

The Mark I- Owners Group retained the General Electric Company to develop and manage a program which would address and resolve the stated concerns.

A two-phase program'vas established; it was described to the NRC in letters submitted during the week of May 5,1975. The Phase I effort, called the Short Term Program (STP), provided a rapid confirmation of the adequacy of 1-3 Revision 0 L

NEDO-21888 the containment to maintain its integrity under the most probable course of the postulated LOCA considering the latest available information on the important h suppression pool dynamic loads. The first phase demonstrated the acceptability of continued operation during the performance of Phase II, called the Long Term Program (LTP). The LTP included detailed testing and analytical work to define precisely the specific hydrodynamic loads for which each containment would be assessed to establish conformance to the original intended design safety margins.

The STP was completed la late 1976 following the docketed submittal by each utility of the documentation listed in References 1.2-1 through 1.2-11 and the associated plant unique analysis of the external torus support system and the externally attached piping. Review of this documentation by the NRC led to issuance of the Mark I Containment Short Term Program Safety Evaluation Report in December 1977 (Reference 1.2-12). This report concluded that licensed domestic BWR Mark I facilities could continue to operate safely, without undue risk to health and safety of the public, during an interim period while the Long Term Program was conducted.

The Long Term Program activities were initiated in June 1976. A detailed description of this program and plans for its implementation are give in References 1.0-13 and 1.0-14. These references define the objectives of the Program, provide Program cask descriptions, and describe the integrated activities leading to a definition of hydrodynamic loads for reevaluation of the containment structure by the individual utilitias. In all, seven Program work areas were established:

1.0 Program Action Plan 2.0 Preliminary Load Evaluation Activities 3.0 Structural Acceptance Criteria 4.0 Generic Structural Evaluation 3.0 Load Evaluation 6.0 Load Mitigation Development Tcsting 7.0 Load Definition Report The extensive testing and analytical efforts performed under the eighteen separate tasks of Area 5.0 and the pool swell and S/RVDL discharge mitigation tasks of Area 6.0 are the primary bases for the load definitions of Area 7.0. h These load definitions are presented in this report.

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, 1.3 LOAD DEFINITION REPORT - OBJECTIVE 1

! The objective of this Load Definition Report (LDR) is to provide load definition i

i procedures for the postulated LOCA and S/RV actuation events for use in structural l reevaluation of the pressure suppression chamber, vent system, S/RV discharge

! lines and other Mark I containment components.

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(~') Table 1.2-1 V

MARK I UTILITIES AND PLANTS IN THE MARK I' 0WNERS GROUP I

f Utility Name Plant Name 4 Boston Edison Company Pilgrim l Boston, Massachusetts !

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Carolina Power & Light Company Brunswick 1,2 Raleigh, North Carolina Commonwealth Edison Company Dresden 2,3

j. Chicago, Illinois' Quad Cities, 1,2 1

j Detroit' Edison Company Fermi 2 4

Detroit, Michigan d

Georgia Power Company Hatch 1,2 Atlanta, Georgia Iowa Electric Light & Power Company Duane Arnold Cedar. Rapids, Iowa Jersey Central Power & Light Company Oyster Creek

, - Morristown, New Jersey 4 Nebraska Public Power District Cooper Columbus, Nebraska j- Niagara Mohawk Power Corporation Nine Mile Point 1 Syracuse, New York t

Northeast Utilities Service Company Millstone Berlin, Connecticut Northern States Power Company Monticello Minneapolis,-Minnesota

- Philadelphia Electric Company Peach Bottom 2,3

, _ Philadelphia, Pennsylvania Power Authority of the' State of New York Fitzpatrick New York, NewLYork Public Service Electric and Gas Hope Creek 1,2

. Newarki New Jersey-Tennessee Valley Authority Browns' Ferry 1,2,3 Knoxville, Tennessee Yankee Atomic Electric Company Vermont Yankee f-~i Westboro, Massachusetts V

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PRESSUR E SUPPR ESSION POOL Figure 1.1-1. Typical Mark I Containment 1-8 Revision 0

VENT MAIN HEADER VENT TORUS OOWNCOMER WATr RFACE O _ - __g 5/ \ - -

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- SUPPORT COLUMNS

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S/RV OISCHARGE PIPE DEVICE Figure 1.1-2. Typical Composite Section Through Suppression Chamber Revision 0 1-9

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T-QUENCHER ENO CAP HOLES ' '/ DISCHARGE (OPTIONAL) DEVICE HOLES SYMMETRICAL ABOUT HORIZONTAL Q ,

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Figure 1.1-4. Discharge Devices Employed in Mark I Plants i.

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NEDO-21888 REFERENCES FOR SECTION 1 1.2-1 J. M. Humphrey, et al. , " Mark I Containment Evaluation Short Term Program Final Report, Volume I, Program Description and Summary of Conclusions", General Electric Company, Report No. NEDE-20989-P, September 1975.

1.2-2 R. H. Buchholz, et al. , " Mark I Containment Evaluation Short Term Program Final Report, Volume II, LOCA Related Hydrodynamic Loads",

General Electric Company, Report No. NEDC-20989-2P, September 1975.

1.2-3 J. M. Humphrey, et al., " Mark I Containment Evaluation Short Term

, Program Final Report, Volume III, Load Application and Screening of 1 titructural Elements", Report No. NEDC-20989-3P, September 1975.

1.2-4 " Mark I Containment Evaluation Short Term Program Final Report, Volume IV, Structural Evaluation", Prepared by the Bechtel Power Corporation for the General Electric Company, Report No.

NEDC-20989-4P, September 1975.

1,2-5 " Mark I Containment Evaluation Short Term Program Final Report, Volume V, Independent Assessment of the Mark 1 Short Term Program",

Prepared by Teledyne Materials Research for the Ceneral Electric Company, Report Uo. NEDC-20989-5, September 1975.

1.2-6 " Mark I Containment Evaluation Short Term Program Final Report, Addendum 1 to Volume IV, Structural Evaluation", Prepared by the Bechtel Power Corporation for the General Electric Company, Report No. NEDC-20989-4PA, November 1975.

1.2-7 " Mark I Containment Evaluation Short Term Program Final Report, Addendum 2, Loads and Their Application for Torus Support System Evaluation", General Electric Company, Report No. NEDC-20989-P, Addendum 2, June 1976.

1.2-8 S. Kayhan, et al. , " Mark I Containment Evaluation Short Term Program Final Report, Addendum 3, Vent Header and Vent Pipe Impact Loads",

General Electric Company, Report No. NEDC-20989-P, Addendum 3, August 1976. !

1.2-9 " Mark I Containment Evaluation Short Term Program - Final Report, Addendum 4, Revisions to Volume IV, Structural Evaluation", 1 Prepared by the Bechtel Power Corporation for the General Electric - j Company, Report No. NEDC-20989-4A, November 1976. ;

1.2-10 " Description of Short Term Program Plant Unique Torus Support l Systems and Attached Piping Analysis", NUTECH, Report No.

MKl-02-012, Revision 2, June 1976.

La 1-13 Revision 0

NEDO-21888 1,2-11 Letter from L. J. Sobon (GE) to V. Stello (NRC), Subject-Mark I Short Term Program Report Questions, October 29, 1976.

1.2-12 " Mark I Containment Short Term Program Safety Evaluation Report",

NUREG-0408, Nuclear Regulatory Commission, December 1977.

1.2-13 Letter from L. J. Sobon (CE) to V. Stello (NRC), Subject-Mark I Containment Program - Program Action Plan (Revision 3), February 1978.

1.2-14 " Mark I Containment Program - Program Action Plan", Revision 3, dated February 15, 1978, General Electric Company, Nuclear Energy Division, February 1978.

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SECTION 2 i

REVIEW OF PHENOMENA 2

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NED0-21888 2.0 REVIEW OF pHE!IOMENA This section describes the sequence of events of the Mark I containment related phenomena for the postulated LOCA and S/RV actuation conditions. A basis is provided for understanding the loading conditions which result from these events.

For a postulated pipe break inside the drywell, three LOCA categories are con-sidered. These three categories, selected on the basis of postulated break size, are referred to as the Design Basis Accident (DBA), Intermediate Break Accident (IBA) and Small Break Accident (SBA). S/RV actuation can occur as a result of a number of system conditions. Although the load magnitudes depend on the initial system conditions prior to S/RV discharge, the sequence of the loading phenomena is the same for all conditions.

2.1 DESIGN BASIS ACCIDENT The DBA for Mark I plants employing jet pumps within the BWR is the instantaneous

() double-ended guillotine breit of the BWR recirculation pump suction line at the reactor vessel nozzle safe-euj : 3 pipe weld. For Mark I plants that do not employ jet pumps within the BWR, the same DBA break definition applies except that the break location is the BWR recirculation pump discharge line at the reactor vessel nozzle safe-end to pipe weld. As discussed in Section 4.1.1, these postulated break conditions result in the maximum flow rate of primary system _ fluid and energy into the drywell, through_the vent system and into the wetvell. The DBA results in the maximum pressurization rate and peak pressure in the drywell; it therefore produces the most limiting pool swell and vent system thrust loads. The DBA event is evaluated up to the time that the low pressure Emergency Core Cooling System (ECCS) starts to flood. the reactor vessel, which occurs at approximately 30 seconds after the pipe break as pre-

~

sented in Section 4.1.1.3. The use of this break to determine the DBA loading for all Mark I plants is verified by a direct comparison of the drywell and wetwell pressure ~ responses resulting from a recirculation line break with that from a main steam line break.

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NEDO-21888 The sequence of events within the wetwell which follow the postulated break, is divided into two phases:

a. Fool S2 ell - This phase covers the dynamic effects of drywell and vent system air being forced through the vent system into the suppression pool to the wetwell airspace.

b. Stean Condensation - This phase covers the dynamic events during the period following initial air clearing when the flow into the suppression pool is a steam - air mixture. The steam is con-densed at the downcomer exit while the air rises through the pool to the wetwell airspace.

The reactor will automatically scram due to high drywell pressure. :tain steam line isolation will occur due to low reactor water level. No mechanical S/RV actuation will occur because of the rapid reactor vessel depressurization and large rate of reactor fluid and energy inventory loss through the break. It is assumed that spurious actuation of a single S/RV can occur at any time during the DBA.

O Uith the postulated' instantaneous rupture of a recirculation line, a pressure wave traveling at sonic velocity would expand from the break location into the drywell atmosphere and through the vent system. The wave amplitude attenuates rapidly as it expands into the larger drywell volume. The wave-front enters the vent system with nearly uniform amplitude, but is greatly attenuated from its initial value at the break location.

The rapid bulk pressurization of the drywell immediately following a postulated DBA and prior to vent clearing theoretically generates a weak compressive wave in the downcomer water legs. This wave could propagate through the suppression pool and induce a much-attenuated loading on the torus shell.

Immediately following the postulated DBA pipe rupture, the pressure and temperature of the drywell atmosphere and vent system increase rapidly. These combined pressure and temperature transients induce mechanical and thermal loadings on the vents, vent header and downcomers. With the drywell pressure increase, the water initially standing in the downcomers accelerates into the 2-2 Revision 0

NED0-21888 pool'until the downcomers clear of water. During this water clearing process, the suppression pool fluid is accelerated causing drag loads on structures within the suppression pool located immediately below the downcomers. Following downcomer water clearing, :he downcomer air, which is at essentially drywell pressure, is exposed to the relatively low pressure of the wetwell, producing a downward reaction force on the torus. The consequent bubble expansion causes 4 the pool water to swell in the torus, compressing the airspace above the pool.

During Gm early stages of this process, the pool swells in the bulk mode i

(i.e., the water is accelerated upward by-the rising air bubble motion) and structures close to the pool surface experience impact and drag loads as the pool impacts-and flows past the structure. Following the initial air bubble expansion and pool surface rise, the bubble pressure decreases as the bubble overexpands and the pool liquid mass decelerates. The net effect of the pool deceleration is an upward lifting force on the torus. Eventually, the bubbles

" break through" to the torus airspace and an air / water " froth" mixture continues i

upward due to the momentum previously imparted to the water slug. The upward motion of this froth mixture causes impingement loads on the torus and other structures encountered, but the loaus are lower in magnitude than the impact loads associated with bulk pool swell.

As the air / water mixture in the suppression pool experiences gravity-induced phase. separation, pool upward movement stops and the liquid falls back. Struc-tures within the path of the fluid motion experience a downward loading; the submerged portion of the torus experiences a small pressure increase. Follow-ing pool fallback, there are waves on the suppression pool surface which' induce low magnitude loads on the downcomers, torus and any other structures close to the water surface.

1 The. transient associated with drywell air venting to the pool typically lasts j for 3 to 5 seconds. Since the air originally contained within the drywell and vent system is transferred to the wetwell airsptce, the wetwell will experience

.a rise in stat' g ru sure. Following air carrytver, there will be a period )

. of high stW U _ a through the vent system. T.e discharge of steam into the

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pool an, i b ...3quent condensation causes poct pressure oscillations which

. will be aansmitt d'to submerged structures anu the torus shell. This pheno- I J.

D)

( '. -menon is referred to as " condensation oscillation". As the reactor vessel C-3 Revision 0

NED0-21883 depressurizes, the steam flow rate to the vent system decreases. The reduced steam ficw rate leads to a reduction in the drywell/wetwell pressure differen-tial. Steam condensation during this period of reduced steam flow is charac-terized by movement of the water / steam interface up and down within the downcomer as the steam volumes are condensed and replaced by surrounding pool water. This phenomenon is referred to as " chugging". During the steam condensation period the downcomers experience a lateral loading due to the asymmetric collapse of steam bubbles at the downcomer exit. Also, the submerged structures and containment walls experience pressure oscillations due to steam bubble forma-tion and collapse.

Shortly after the postulated pipe rupture, the Emergency Core Cooling System (ECCS) automatically begins to pur.p water from the plant condensate storage tank and/or the suppression pool into the reactor pressure vessel (RPV) to flood the reactor core. Eventually water cascades into the drywell from the break, causing steam condensation and drywell depressurization. As the drywell pressure falls below the pressure in the wetwell airspace, the drywell vacuum relief system is actuated and air from the wetwell enters the drywell, equaliz-ing the drywell and wetwell pressures slightly above their initial values.

Following vessel flooding and dry all/wetwell airspace pressure equalization, suppression pool water ir continually recirculated from the pool to the reactor vessel by the ECCS pumps. The core decay power results in a slow heat-up of the suppression pool. The suppression pool cooling mode of the Residual Heat Removal (RHR) system is manually actuated to remove energy from the suppression pool to return the containment to normal temperature conditions.

2.2 INTERMEDIATE BREAK ACCIPENT The IBA for a FMR Mark I plant is a postulated pipe rupture small enough that rapid reactor depressurization will not occur, but large enough that the High-Pressure Coolant Injection (HPCI) system cannot maintain reactor vessel water level. The IBA is defined as a liquid line break of 0.1 ft .

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] The high drywell pressure resulting from the postulated accident conditions will scram the reactor. The sequence of events following this scram could eventually lead to closure of the main steamline isolation valves (MSIV) due to low reactor sater level. The close ' of the MSIV's would result in an increase in RPV pressure that is relieveo by opening the S/RV's.

Following the postulated break, steam fills the drywell causing the drywell pcessure slowly to increase and displace the water initially in the submerged portion of the vent system into the suppression pool. The drywell pressure transient is sufficiently slow that the dynamic effect of the water clearing in the vents is negligible. The subsequent clearing of air from the vent system occurs more slowly than for the DBA and thus imparts lower loadings to the wet-well components. As the flow of air, steam and water continues from the drywell to the wetwell, the wetwell airspace pressure increases. Following the initial purge of air from the drywell, steam begins to flow through the vent system and condenses within the suppression pool. As with the DBA event, the conden-sation oscillation and chugging phenomena occur during the steam condensation process.

y The Automatic Depressurization System (ADS) will actuate due to high drywell pressure and low reactor water level at approximately 300 seconds for a plant with turbine driven feedwater pumps. For plants with motor driven feedwater ,

pumps it is assumed that the feedwater system continues to maintain the vessel level until the pumps are manually tripped at 600 seconds. The ADS will initiate approximately 300 seconds af ter the motor driven pumps are tripped, or 900 seconds after the break occurs. The reactor will be depressurized approximately 200 seconds af ter the ADS is initiated. Thus, the IBA is evaluated to 500 seconds for plants with turbine driven feedwater pumps and 1100 seconds for plants with motor driven feedwater pumps. During operation of the ADS, steam from the RPV is vented directly to the suppression pool via the S/RVDL. As the reactor depressurizes, the Core Spray systems and the Low Pressure Coolant Injection (LPCI) mode of the RHR system will be activated to flood the RPV and cool the core. Eventually, water will cascade into the drywell causing steam condensation and drywell depressurization. As the drywell depressurizes below

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l the pressure within the wetwell airspace, the wetwell to drywell vacuum l breakers open, equalizing the containment pressures and terminating the event. Since the reactor depressurization transient is less rapid for the IEA chan for the DBA more decay heat is discharged into the suppression pool which results in the suppression pool temperature being higher for the IBA than for the DBA at the time of complete reactor depressurization.

2.3 SMALL BREAK ACCIDENT The SBA for a BWR Mark I plant is a postulated pipe rupture in which the fluid loss rate from the reactor system is insufficient to depressurize the reactor and small enough that HPCI operation is sufficient to maintain reactor water level. The SBA is defined as a 0.01 ft steam break. Following the break, the drywell pressure will slowly increase until the high drywell pressure scram set point is reached. Following the reactor scram, MSIV closure may occur due to the water level transient in the reactor. If the main steam lines isolate the reactor system, pressure will increase and safety / relief valves will open intermittently to control system pressure.

O The drywell pressure increase gradually depresses the water level in the vents until the water is expelled and air and steam enter the suppression pool. The rate of airflow will be such that the air will bubble through the pool without causing pool swell. The steam will be condeased and drywell air will pass to the wetwell airspace. A gradual pressurization of the wetwell results at'a rate dependent upon the break size and degree of drywell steam / air mixing.

Eventually, the steam / air flow through the vents results in essentially all the drywell air being transferred into the wetwell. Following the air transfer, wetwell pressurization occurs at a rate dependent on the suppression pool heatup rate. Condensation oscillatian will not be present due to the low mass flux, but ch gging may occur.

Operator actions will determine the subsequent course of events. It is assumed that there will be no operator action for 10 minutes following the pipe break.

After 10 minutes, it is assumed that the operator will rapidly depressurize the reactor via ADS. An individual plant may choose to conduct a plant unique analysis to demonstrate that a less rapid cooldown rate is acceptable.

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~ NEDO-21888

'2.4 SAFETY / RELIEF VALVE ACTUATION

)

Prior to the initial actuation of a safety / relief valve caused by a normal operational transient, the S/RV discharge lines contain air at atmospheric pressure and suppression pool water in the submerged portion of the piping.

Following S/RV actuation, steam enters the S/RVDL compressing the air within the line and expelling'the water slug into the suppression pool. . As the water is cleared, the S/RVDL undergoes a transient pressure loading. The submerged portion of .the'line also experiences water clearing thrust loads caused by directional changes of ' momentum within the line and discharge device. The water jets. entering the pool from the discharge device result in drag loads on submerged structures within the influence of the jets.

Once the water has cleared from the discharge device, the compressed air enters the pool in the form of high pressure bubbles. These bubbles expand result.ing in an outward acceleration of the surrounding pool water. The

- momentum of the accelerated water results in an overexpansion of the bubbles, _

causing the bubble preseure to become negative relative to the ambient pressure

( )- of the surrounding pool. This negative pressure slows and reverses the motion of the _ water, leading to a compres' s ion of the bubbles and a positive pressure relative to that of the pool. The bubbles continue to oscillate in this manner for approximately one second as they rise to .the pool surface. As the bubbles oscillate, the associated local pool motion causes drag loads on nearby sub-merged structures.' The _ positive and negative pressures developed within the bubbles attenuate with distance and result in an oscillatory pressure loading-on the submerged portion of the torus shell.

1 Following air and water _ clearing ~ f rom the line, steam is discharged through :

the- line to the suppression pool and condensed therein. The steam condensation phenomenon produces a high-frequency, low-magnitude oscillatory loading on the torus shell.

Following closure of an S/RV, the steam pressure in the S/RV decreases rapidly as the steam flows out into the pressure suppression pool. At a sufficiently low steam pressure, pool water reenters the S/RVDL. A rapid depressuri-zation of the line then occurs as the steam remaining in'the line is condensed 2-7 Revision 0

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NEDO-21888 by the in-flowing water. This depressurization causes the water to reflood into the S/RVDL and the vacuum breaker valve on the S/RVDL to open, allowing g

drywell gas to enter the line. The reflooding water may rise in the line to a level somewhat above its initial pre-S/RV actuaticzi level before equilibrium is reestablished. The actual reflood water level depends primarily on the size of the S/RVDL vacuum breaker.

2.5 OTHER CONSIDERATIONS In addition to the LOCA and S/RV discharge events, the dynamic behavior of the pool water during and after seismic motion is evaluated. Horizontal and vertical seismic motion is transmitted to the pool tnrough the torus support system. The seismic motion generates a sloshing wave behavior on the pool surface which produces minimal loads on the downcomers, submerged structures and the torus shell.

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4 SECTION 3 LOAD CdMBINATIONS l

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-( 3.0 LOAD COMBINATIONS

. l]s This section identifies the timing sequence of the loading conditions for the structural components in the wetwell due to the hydrodynamic phenomena described in Section 2. In addition, those sections of the report which define the hydro-dynamic structural-loading are identified.

The combinations of loading conditions affecting the various major structural components are presented in Figures 3.0-1 through 3.0-5. These figures identify the hydrodynamic loading conditions resulting from LOCA and from S/RV discharges.

Seismic loadings, and structural and water deadweight loads are not presented in this report. The figures in this section provide the timing sequences of the

~ loading conditions for all three postulated LOCA's, the DBA, IBA, and SBA.

The~ lengths of the bars in the figures indicate the time periods during which a loading condition may exist. A loading condition such as condensation oscillation is assumed to exist continually during the indicated time period. For S/RV d .a -

charge loads, the duration of the loading is short, but the loads may occur a: any

['] time during the indicated time period. Loads are considered to act simultane<.usly V

on a structure at a specific time -if the loading condition bars overlap at that time.

The bar charts for the DBA show the loading condition combinations for postulated breaks in the recirculation line, the main steam line or any other breaks large enough to produce significant pool swell loads. The bar chart for the IBA shows 4

conditions for a break size sufficiently large such that the HPCI system cannot

-prevent ADS actuation on low-water level, but smaller than that break size which would produce significant pool swell loads. A break size of 0.1 ft 2 13- assumed for the 'IBA. The bar charts for the SBA show conditions for a break size equal to 0.01 ft in cross-sectional flow area. For an SBA, the HPCI system would be able to maintain the water level and the reactor would be depressurized by means of operator initiation of'the-ADS.

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Table 3.0-1 identifies the S/RV discharge loading conditions.

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NEDO-21888 Table 3.0-2 identifies the timing nomenclature used in the bar charts.

Table 3.0-3 identifies the S/RV and LOCA loads which potentially affect struc-O' tural components and identifies the appropriate section of the LDR defining the loads. For S/RV piping and other structures within the wetwell, the location of the structural component must be considered to determine if any of the identi-fled conditions affect the structure.

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NED0-21388 d

Table 3.0-1 S/RV DISCHARGE LOAD CASES FOR MARK I STRUCTURAL ANALYSIS Any 1 .GS Multiple

Valve Valves Valves

~! Initial Conditions 1 2 3 A. First Actuation Al A2 A3

3. First Actuation Leaking S/RV** B3 C. Subsequent Actuation C3 e

n :;umcer (ene or cora) and location of 5/RV's assumed to actuate will be determined by plant unique analyses.

- Only one S/RV is assumed to leak for plants equipped with Ramshead dis-

' charge devices for normal operating transients. The loads for T-Quencher discharge devices are not affected by leaking S/RV's. No S/RV's are cen-sidered to laak prior to a '_CCA.

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NED0-21888 l l

Table 3.0-2 EVENT TIMING NOMENCLATURE Time Description t

y Tha onset of condensatiori oscillation (see

'iection 4.4) t The beginning of chugging (see Section 4.5) 2 t The end of chugging (see Section 4.5) 3 t

4 Time of complete reactor depressurization (see Section 4.1) t ADS

,U)S actuation on high drywell pressure and low reactor water level, break size and plan .epend-ent. The ADS may also be actuated by the operator for a small break accident.

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NED0-21888 Table 3.0-3

- STRUCTURAL LOADING f I Other Wetwell!

Interior l

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4.1 Containment .*re w rv a nd Tv epe r t a r. X X X X X' X X 5!.X l i

4.2 Vent Systee Thrust Leads s X X l f root a e!1 ! I 4.3

. 3.1 Torus set 'Jertical Loads X X ,

4. 3. 2 farus R. ell Pressare Histaries X X ; .

4.3.3 Vent Systei tmpact and Drag X X X j

4. 3. 4 Ispact saJ prag o- Other Structures X XlX 4.3.3 Froth I :pingemer X X X! X

4. 3. 6 Pool Falib ad X X X t

' X 4.3.7 LeCA Jet X

4. 3. 8 LOCA Abble Drag X X X i

4.4 Condensatten oscillation ,

i 1 4.4.1 Taras Shell Loads X X 4.4.2 Load on Abner,ted structures ! X X X.

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u.4.3 b reral Loads on Down:omers .

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I 4.4.4 Verit Svetes Loads X X: ;

4.5 Chugging i i i 4.5.1 Torus shall Loads X X t

e 4.5.2 Loads on submerg.J structures X X 4.5.3 Lateral Laads on Downc,mers i X. X 4.5.4 Vent system Loads lX X'

3.2 T

encher Loads 3.2.L Discharge Line Clearing X i ,

3.2.2 Toras Shell *ressures X X t 5.2.4 let Loads on submerged Structures X X X X '

5.2.5 Air Sabble Drag X' X .s 5.2.6 Thrust LoAle on T-eaer.cher Arms X i

e 3.2.7 5/RVDL Lnytronmental Temper cure X l 3.3 eassne44 Lo. ads J 3.3.1 31scharze Line Clea ring X 5.3.2 Terus shell Pressures X X

! 3. ).4 et Lwds on 5.abeerged S ructure . X X X s.3.3 Air hbble Dr24 X X X X 3.3.6 J/RV3L Laviron tental Temper e .

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LOCA PRESSURE AND TEMPERATURE TRANSIENTS SECTION 4.1 VENT SYSTEM AIR, STEAM AND LIQUID FLOW AND PRESSURE TR ANSIENTS SECTION 4.2 z SINGLE S/RV ACTUATION 9 (S/RV EVENT CASE A1) t SECTIONS 5.2 OR 5.3 8

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CONDENSATION OSCILLATION SECTION 4.4 CHUGGING SECTION 4.5 NOTE: CONSIDERATION SHOULD BE GIVEN TO RE ACTION LO ADS FROM ATTACHED STRUCTURES.

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==0.1 ~ 1.5 t1 t2 t3 TIME AFTER LOCA (sec)

Figure 3.0-1. Loading Condition Combinations Structures Affected: Vent Header, Main Vents, Downcomers and Torus Shell i Accident Condition: DBA O

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LOCA PRESSURE AND TEMPER ATURE TR ANSIENTS SECTION 4.1 SINGLE S/RV ACTU ATION *

(S/RV EVENT CASE All 1 SECTIONS 5.2 OR 5.3 9

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8 o S!RV ACTU.\TlON ON 2 SETPOINT (S/i1V EVENT CASE A3, C3) ADS ACTUATION oa (S/RV EVENT CASE A2)

CONDENSATION OSClLLATlON SECTION 4.4 O

CHUGGING SECTION 4.5 NOTE: CONSIDERATIONS SHOULD 8E GIVEN TO REACTION LOADS FROM ATTACHED STRUCTURES.

LOADING DOES NOT COMBINE WITH OTHER S/RV CASES l 1 l l l t1 TADS > 90 t2 t3 t4 TIME AFTER LOCA (sec)

Figure 3.0-2. Loading Condition Combinations Structures Affected: Vent Header, Main Vents, Downcomers, Torus Shell and Submerged Structures Accident Condition: IBA O

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LOCA PRESSURE AND TEMPERATURE TRANSIENTS SECTION 4.1 SINGLE S/RV ACTUATION' (S/RV EVENT CASE A1)

SECTIONS 5.2 OR 5.3 g - - - - - - . - - . - - - - - - - - - - - - - - -

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$ OPERATOR INITIATION OF ADS l

z (S/RV EVENT CASE A2) i k

9 S/RV ACTUATION ON SET POINT (S/RV EVENT CASE A3, C3)

CHUGGING O

SECTION 4.5 NOTE: CONSIDERATION SHOULD BE GIVEN TO REACTION LOADS FROM ATTACHED STRUCTURES.

LOADING 00ES NOT COMBINE WITH OTHER S/RV CASES.

I I I I t2 a 600 13 t4 TIME AFTER LOCA (sec)

Figure 3.0-3. Loading Condition Combinations Structures Affected: Vent Header, Main Vents , Downcomers , Torus Shell and Submerged Structures.

Accident Condition: SBA l

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LOCA PRESSURE AND TEMPERATURE TRANSIENTS SECTION 4.1 5

G 2 5 FROTH IMPINGEMENT j SECTION 4.3.5 0 i s

3 a

1 POOL SWELL F ALLB ACK SECTION 4.3.6 POOL SWELL IMPACT AND OR AG SECTION 4.3.4 NOTES: 1. STRUCTURES BELOW MAXIMUM POOL SWELL HEIGHT

2. STRUCTURES ABOVE MAXIMUM POOL SWELL HEIGHT l l l

~0.1 ~0.7 ~ 1.5 TIME AFTER LOCA (sec) l Figure 3.0-4. Loading Condition Combinations Structures Affected: Small Structures Above Pool

Accident Condition: DBA i

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For S/RV lines consider possibility of loads produced by S/RV actuation Case Al.

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i NEDD-21888 O

LOCA P'RESSURE AND TEMPERATURE TRANSIENTS SECTION 4.1 SINGLE S/RV ACTUATION' (S/RV EVENT CASE A1)

SECTIONS 5.2 OR 5.3 g ______________-_______-

=

0 8

j CONDENSATION OSCILLATION 3 SECTION 4.4 8a CHUGGING SECTION 4.5 POOLSWELL FALLBACK SECTION 4.3.6 LOCA AIR BUBBLE SECTION 4.3.8 l LOCA WATER j JET FORMATION NOTE: CONSIDERATION SHOULO BE GivEN TO SECTION 4.3.7 RE ACTION LOADS FROM ATTACHED STR UCTU RES.

' LOADING DOES NOT COMBINE WITH OTHER S/RV CASES.

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%.1 ~0.7 ~ 1.5 ti t2 t3 TIME AFTER LOCA (sec)

Figure 3.0-5. Loading Condition Combinations Structures Affected: Submerged Structures Accident Condition: DBA O

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t SECTION 4 f

LOCA RELATED LOADS

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G 4.0 LOCA RELATED LOADS

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This section specifies the methodology used to determine the wetwell and-vent system primary . structural loads encountered during a postulated LOCA.

The drywell pressure = and temperature responses which are calculated for a postulated DBA are also presented for a typical Mark I plant.

The primary structural loads generated during a LOCA are a result of the following:

a. The pressures and temperatures within the drywell, vent system and wetwell torus

b. The fluid flow through the vent system

c. The initial LOCA bubble formation in the pool and the resulting displace-ment of water due to pool swell

d. Steam flow into the suppression pool (condensation oscillation and/or chugging)

e. S/RV actuation loads (discussed in Section 5).

The spectrum of break sizes addressed consists of the DBA, IBA and SBA as pre-viously defined in Sections 2.1, 2.2 and 2.3 of this report, respectively.

(vD .

4.0-1/4.0-2 Revision 0

NEDO-21888 4.1 CONTAIhHENT SYSTDI TDfPERATURE AND PRESSURE RESPONSE

./^T

)

The following sections present the drywell and wetwell pressure and temperature responses due to the DBA, IBA and SBA.

4.1.1 Design Basis Accident A. typical set of containment-temperature and pressure responses to a DBA are pre-sented in Figures 4.1.1-1 and 4.1.1-2, respectively. Following the pipe break, the drywell pressure increases and the water initially in the submerged portion of the vent system accelerates into the suppression pool. Until the water is com-pletely cleared from the vent system, the air pressure in the wetwell does not begin to change significantly. Starting near point A (Figure 4.1.1-2) both drywell and wetwell pressures increase as the vents clear. The RPV blow-down continues into the.drywell and flow continues - from the drywell through the vent system and into the wetwell. Flow to the wetwell is initially air, but sub-sequently becomes a mixture of air, water and steam between points B and D. Point B is characterized by the depletion of the fluid inventory initially f-~ occupying the recirculation loop. Point C occurs at the depletion of the sub-

~-

cooled fluid inventory of the recirculation loop and the RPV, and at the transi-tion -to saturated fluid blowdown from the RPV.

The drywell pressure increases more rapidly than the wetwell pressure until the mass flow rate . into the drywell drops sharply. At approximately point D, the liquid level inside the reactor has dropped below the elevation of the pipe break and the blowdown, which was previously all liquid, becomes a steam-liquid

' mixture. The flow through the vent system now begins to decrease and results in a corresponding decrease in the vent system pressure. This permits the drywell and wetwell pressures to converge until they differ only by the hydrostatic pres-sure deteruined by the vent system submergence at point E. This description of the timing of events applies to both temperature and pressure transients in the drywell and wetwell, i

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NED0-21888 4.1.1.1 Analytical Procedure The details of the analytical models used to simulate the short-term transient O

response of the drywell and wetwell to a DBA are presented in References 4.1.1-1 and 4.1.1-2. These models briefly are as follows:

a. A vessel blowdown model which simulates the reactor vessel response to a LOCA and determines the break mass and energy flow rate. (Described in Reference 4.1.1-1)

b. A drywell model which determines drywell thermodynamic conditions as a result of the mass and energy flows into and out of the drywell. (Des-cribed in Reference 4.1.1-1)

c. A model which determines vent clearing time and flow behavior during the LOCA. (Described in Reference 4.1.1-2)

d. A model which simulates suppression pool temperature response by a simple mass and energy balance. (Described in Reference 4.1.1-1)

O

e. A wetwell free airspace model which is used to calculate the airspace pressure and temperature response. (Described in Reference 4.1.1-1)

The models described have been verified by comparison with model tests performed at the Bodega Bay and Humboldt Bay type containment test facilities as described ir References 4.1.1-2 and 4.1.1-3.

4.1.1.2 Assumptions

a. May-Witt fuel decay heat and sensible energy as defined in Table 4.1.1-1 for all plants. This assumption is consistent with the original thrk I plants licensing bases documented in plant FSAR's or PSAR's.

b. The blowdown model is the Homogeneous Equilibrium Model (HEM) which is fully described in Reference 4.1.1-4 This blovdown model replaces the slip ficw model described in Reference 4.1.1-1.

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NED0-21888

c. Steam condensation on drywell structures and internal components is

.()

\_, conservatively neglected.

d. Plant unique calculated vent system loss coefficients, which include the downcomer exit loss,'as identified in Table 4.1.1-2, are used. The method of calculating these loss coefficients has been verified against the 1/12 scale 3-D tests described in Reference 4.1.1-5.

e. The flow of liquid, steam and air in the vent system is assumed to be a homogeneous mixture based on the instantaneous mass fractions in the drywell. This assumption yields increased vent system flow density resulting in higher flow losses and therefore higher drywell pressure.

These conservative drywell pressures were used as the forcing function for the plant unique pool swell tests (see Section 4.3.1).

f. The DBA for plants employing jet pumps within the B'WR is the instan-taneous double-ended guillotine break of the recirculation pump suction line at the reactor vessel nozzle safe-end to pipe weld. For plants that do not employ jet pumps within the BWR, the DBA break location

(~}

is the BWR recirculation pump discharge line at the reactor vessel nozzle safe-end to pipe weld. The fluid inventory initially occupying the recirculation loop includes the subcooled fluid in the recircula-tion system and the reactor vessel region external to the shroud. This maximizes the initial mass and energy release rates to the drywell.

~

g. The bulk wetwell pool and airspace temperatures are assumed equal throughout the transient. This assumption maximizes the wetwell air-space temperatures and pressures.

h. The initial drywell temperature is 135'F at the time of the break. This is the normal operating temperature of the drywell as supplied by the utilities,

i. The~ initial'drywell relative humidity is 20% and the initial wetwell relative-humidity-100%. The 20% drywell humidity represents the mini-mum value of the normal range and maximizes the mass of noncondensibles f-'s .

t ,) in the drywell. The 100%.wetwell humidity maximizes the vapor pressure and thus-results-in the maximum total wetwell pressure.

4.1.1-3 Revision 0

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

NED0-21888

j. The reactor vessel control volume is assumed to include the fluid and structural masses and energy, of the following primary system components, thereby conservatively predicting the blowdown mass and energy and the resulting drywell and wetwell pressures and temperatures:

(1) Reactor vessel (2) Recirculation loops (3) bbin steam lines to the inboard isolation valves (4) RCIC steam line to the first normally closed valve (5) RHR shutdown line to the first normally closed valve (6) LPCI line to the first normally closed valve (7) Core spray line to the first normally closed valve O

(8) HPCI line to the first normally closed valve (9) Feedwater line to the point in the feedwater system where the temperature during normal operation is equal to the saturation temperature at the final calculated reactor vessel pressure.

Feedwater mass below this temperature will not flash during reactor vessel depressurization and therefore will not discharge to the containment.

(10) Isolation condensors for applicable plants.

k. The reactor is operating at 102% of licensed thermal power. This value for a typical plant is identified in Table 4.1.1-3. This assumption maximizes the core decay heat.

1. The initial temperature of the suppression pool is assumed to be equal to the arithmetic mean of the operating range supplied by the Mark I owners. A typical plant suppression pool temperature is shown in Table 4.1.1-3.

4.1.1-4 Revision 0 L

m. The highest torus water level within the normal operating range is

,Q fs,/ assumed, thereby providing the maximum downcomer submergence and mini-mum free airspace. See Table 4.1.1-3 for typical plant conditions.

n. The area of the instantaneous guillotine break is as tabulated in Table 4.1.1.4.

o. No credit is taken for normal auxiliary power. This maximizes the time it takes the ECCS systems to become operable,

p. The single failure of one RHR pool cooling loop is assumed. This has no impact on the short term containment pressure and temperature response,

q. For plants with a drywell-to-wetwell pressure differential sys-tem, the initial drywell and wetvell pressures are selected based on the AP system operating at the nominal value of the normal range specified. The nominal value was set equal to the arithmetic mean of the maximum and minimum values supplied by the Mark I Owners.

For plants without a drywell-to-wetwell pressure differential sys-tem, the initial drywell and wetwell pressures are conservatively assumed to be 0.75 psig. Typical plant conditions are shown on Table 4.1.1-3.

4.1.1.3 Analysis-Results Plant unique containment pressure and temperature responses have been generated l for each plant. These plant unique results are presented in Plant Unique Load Definition documents and submitted individually to the Mark I owners. Typical responses are shown in Figures-4.1.1-1 and 4.1.1-2, for a plant with condi-tions as tabulated in Table 4.1.1-3. ,

l O

%)

4.1.1-5 Revision 0

UEDO-21888 4.1.1.4 Determination of Limiting Break Size and Location In order to verify that the recirculation line break will result in the most O

severe drywell pressure rate and peak pressure loading for all Mark I's, the system response during a postulated main steam line was compared with the response during a recirculation line break. The analytical bases and the conclusions are presented in this section.

The containment system response to a recirculation line break is described in Section 4.1.1. The main steam line break is similar except that the initial blowdown is steam instead of liquid. Drywell pressurization and flow through the vent system into the wetwell is similar to the recirculation line break.

The criteria used to determine the limiting event for containment structure load-ing were the drywell rate of pressurization and the peak drywell pressure. For each plant, the maximum break areas were determined for both main steam and recirculation lines and are listed in Table 4.1.1-4.

During the early portion of the blowdown where the peak drywell pressurization rate and peak drywell pressure are occurring, the plant with the largest ratio of main steam line break area to recirculation line break area was selected as the test case. For other plants, the difference in loading due to a recircula-tion line break from that due to a main steam line break will be larger than the difference observed in the test case plant.

Analyses were performed for both a postulated main steam line break and a recircu-lation line break using the analytical procedures described in Section 4.1.1. The break area and characteristics of the blowdown were adjusted to represent the appropriate break location. The resulting drywell and wetwell pressure time his-tories from the two analyses are presented, superimposed, in Figure 4.1.1-3. The recirculation line break analysis shows both a higher rate of pressurization and a higher peak pressure for the drywell and wetwell and therefore would impose a higher overall containment loading. Because the resulting pressure loading deter-mined for the test case plant indicates that the recirculation line break is the limiting case, it is concluded that for all Mark I's, the recirculation line break generates the most severe DBA containment loading.

O 4.1.1-6 Revision 0

NED0-21888 Table 4.1.1-1 FUEL DECAY HEAT AND SENSIBLE ENERGY FOR ALL MARK I PLANTS Normalized Normalized Decay Heat

Decay Heat

Time and Sensible Time and Sensible (sec) Energy (sec) Energy 0 1.000 1 x 10 _1 1.005 1 x 10 4 4

1.201 x 10 -2

-1 -1 2 x 10 9.657 x 10 2 x 10 1.008 x 10 -2 4x10j 8.251x10f 4 x 10 8.125x10[3 6 x 10 7.093 x 10 6 x 10 7.394 x 10 6.646 x 10 -3

-1 -1 4 8 x 10 6.503 x 10 8 x 10 1 5.310 x 10 -1 1 x 10 55 6.245 x 10 -3 3

2 4.848 x 10 -1

-1 2 x 10 5 5.126 x 10 -3 4 5.459 x 10 -1 4 x 10 4.096 x 10 -3 6 5.663 x 10 -1 6 x 10 55 3.596 x 10 -3 8 5.378 x 10 8 x 10 3.196 x 10 -3 1 x 10 1 4.807 x 10 -1 1 x 10 6 2.985 x 10 -3 2 x 10 1 2.051 x 10 -1 2 x 10 6 2.367 x 10 -3

-2 4 x 10 1 5.501 x 10 -2 4 x 10 60 1.826 x 10 -3 g 6 x 10 1 4.220 x 10 -2 6.x 10 6 1.573 x 10 -3 8 x 10 3.952 x 10 8 x 10 1.430 x 10 -3 3.811 x 10 -2 2

1 x 10 7 1 x 10 2 x 10 2 -2 3.365 x 10-2 2 x 10 1.287 9.790 xx10-10 '3

. 4 x 10 22 2.827 x 10 -2 4 x 10 7.480 x 10 -4

-4

'6 x 10 2.549 x 10 -2 6 x 10 7 6.820 x 10-8 x 10' 2.365 x 10 8 x 10 6.270 x 10 '

1 x 10 33 2.229 x 10 -2 1 x 10 88 -4 2 x 10 6.050 x 10-1.841x10[2 2 x 10 5.335 x 10 '

4 x 10 33 1.512 x 10 -22 6 x 10 1.353 x 10 8 x 10 3 1.257 x 10 -2 i

1

May Witt decay heat normalized to initial power !

l 1

). 4.1.1-7 Revision 0

. _ , . , . , . ~. , . ,,

Table 4.1.1-2 PLANT VENT SYSTEM LOSS COEFFICIENTS Plant Vent System Loss Coefficients Browns Ferry 1, 2, 3 5.32 Brunswick 1, 2 5.17 Cooper Station 5.51 Dresden 2, 3 5.17 Duane Arnold 4.65 Enrico Fermi 2 5.51 Fitzpatrick 5.17 Hatch 1, 2 5.51 Hope Creek 1, 2 5.51 Mills tone 5.17 Monticello 5.17 Nine Mile 1 5.6 Oyster Creek 1 5.6 Peach Bottom 2, 3 5.17 Pilgrin 5.17 Quad Cities 1, 2 5.17 Vermont Yankee 5.17 l

4.1.1-8 Revision 0

,, - . .. . . . - . . . . ~ . - - . - - - -. - - - - . ~ _ . . . - . . ~ - - . . - . - - . . . . . - _ . - . . - . .-. - . . ,

i NEDO-21888 1

k: ' 4 Table 4.1.1.3 L

TYPICAL PLANT CONDITIONS AT INSTANT OF DBA PIPE BREAK l

i i

102% Licensed Power (MWT) '1703 l

! Initial Suppression-Pool Temperature (*F) 77.5 i.

Downcomer Submergence (ft) 4.25 E

3 DC NM Airspace Volume (ft )

,Wetwell 97580 1

Drywell 1.25 :

0.

, Airspace Pressur.' (psig)

,Wetwell 0.10 i

4 1

(-

1-1 J

].

l l

l 01

'4.1.1-9 . Revision 0 -

NEDO-21888 Table 4.1.1-4 SLM!ARY OF !! ARK I PLANT BREAK AREAS Break Area (ft ) Area Ratio Main Steam /

Plant !!ain Steam Line Recirculation Line Recirculation Browns Ferry 1, 2, 3 3.934 4.216 0.933 Brunswick 1, 2 3.185 4.258 0.748 Cooper Station 3.055 4.158 0.735 Dresden 2, 3 2.402 4.261 0.564 Duane Arnold 2.197 2.515 0.874 Enrico Fermi 2 4.091 4.119 0.993*

Fitzpatrick 3.083 4.172 0.739 Hatch 1/2 3.076 4.141/3.988 0.743/0.771 Hope Creek 1, 2 3.960 4.177 0.948 Millstone 2.233 4.353 0.512 Monticello 1.803 4.015 0.449 Nine !!ile 1 3.887 7.053 0.551 Oyster Creek 1 3.308 6.056 0.546 Peach Bottom 2, 3 3.934 4.147 0.949 Pilgrim 2.102 4.359 0.482 Quad Cities 1, 2 2.402 4.261 0.564 Vermont Yankee 1.801 4.145 0.434

Test case plant.

O 4.1.1-10 Revision 0

NED0-21888 O l 1

)

300

-VENT CLE ARING -(A)

-lNVENTORY DEPLETION -(B) l

. -SUBCOOLEO FLUlO DEPLETION -(C) e

' f-BREAK UNCOVERY -(D) )

' (El- PRES $URE EQUALIZATION e l 3 .

Q 150 c:

Y WETWE LL TEMPERATURE f

2

l ,

._. - '7 i

l

' I l o

0 10 20 30 1

TIME AFTER BREAK INITIATION (sec) I 1

I I

Figure 4.1.1-1. Typical Containment Temperature Response to Design Basis Accident (Recirculation lne Break) i O

's 1

1 4.1.1-11 Revision 0 I

NEDO-21888 Ol l

60

- VENT CLE ARING - ( A)

- INVENTORY DEPLETION - (8)

- SUBCOOLED FLUID DEPLETION - (C)

- BR E AK UNCOVERY- ID) 40 - -

l (E)- PRESSURE EQUALIZATION -

s

2 4 ' DRYWELL e a 5

M E :

W ETW E Lf.

20 - -

t e

l 0

O 10 20 30 TIME AFTER BREAK INITIATION (sec) l l Figure 4.1.1-2. Typical Containment Pressure Response to Design Basis Accident (Recirculation Line Break)

O 4.1.1-12 Revision 0

I l

NED0-21888 I

1 1

. O l l

60 I 4

/\ ,s

/ \

\ e#

[ %,j

\

DRYWELL 40 - 7

\

l s i

l N i %

a l N 3 DRYWELL \g ,

a I l E g

~

l O l WETWELL 20 ] ",,.* *" WETWELL I

l

a== == RECIRCULATION LINE BREAK MAIN STEAM LINE BREAK 0

0 10 20 30

! TIME AFTER BREAK INITIATION (sec) i 1

l r

Figure 4.1.1-3. Drywell Pressure Response to Main Steam Line Break and Recirculation Line Break l

lO 4.1.1-13/4.1.1-14 Revision 0

. - . . . - . _ . - . . - . . , . . . - . . - . . - . - - _ . . . _ .- _ . ,~.-. ---.._ -

. REFERENCES FOR SECTION 4.1.1 4.1.1-1 W. J. Bil:nin, "The GE Mark III Pressure Suppression Containment System Analytical Model," General Electric Company, Report No. NED0-20533, June 1974.

4.1.1-2 " General Electric Pressure Suppression Containment Analytical Model,"

General Electric Company, Report No. NEDO-10320, April 1971; Supple-ment 1, May 1971; Supplement 2, June 1973.

1 4.1.1-3 Bodega Bay Atomic Park Unit Number 1, Exhibit C - Preliminary Hazards Summary Report, Appendix I - Pressure Suppression Test Program, Pacific Gas and Electric Company, Docket No. 50-205 December 28, 1962.

4.1.1-4 F. J. Moody, " Maximum Discharge Rate of Liquid-Vapor Mixtures from Vessels," General Electric Company, Report No. NED0-21052, September 1975.

4.1.1-5 Electric Power Research Institute, "Three-Dimensional Pool Swell Modeling of a Mark 1 Suppression System," EPRI NP-906, October 1978.

O 4

O' o

l--

4.1.1-15' Revision 0 i

-w+w e er w.c----g. y - w

.NEDO-21888 4.1.2 Intermediate Break Accident

/~'s V

Typical containment temperature and pressure responses to an IBA are presented in Figures 4.1.2-1 and 4.1.2-2, respectively. Following the break, the dry-well air pressure increases an.d slowly forces the water initially in the sub-merged portion of the vent system into the suppression pool. The drywell pressure transient is sufficiently slow that the dynamic effect of the water in the vents is negligible and the vents clear when the drywell-to-wetwell differential pressure is equal to the submergence pressure at point A in Figure 4.1.2-2. As the flow of air, steam and water continues from the drywell to wetwell, the drywell pressure increases at approximately the same rate as that for the wetwell airspace. After all the drywell air has been purged into the wetwell at point B, the pressure rise for the entire containment slows. The subsequent initiation of the automatic depressurization system at point C, vents primary system fluid into the suppression pool causing its temperature to rise rapidly. The wetwell airspace temperature and pressure then increase. As the ECCS fluid is injected into the reactor, the temperature of the fluid discharged into the

~~

drywell will drop thereby condensing the steam in the drywell which reduces

\- / the drywell pressure. The wetwell to drywell vacuum breakers open and the wetwell pressure is reduced. The containment response is determined for the IBA up to the time when the reactor vessel pressure equalizes with the dryvell pressure, which occurs at approximately 50 psia at point D.

9 During the depressurization of the reactor vessel, the suppression pool temp-erature increases rapidly due to the transfer of stored reactor vessel energy to the pool. The suppression pool temperature transient will level out after the reactor vessel is depressurized. The pool temperature will rise much less rapidly since the heat input to the pool is limited to the core decay heat, and the RHR system is removing heat from the pool. No S/RV, condensation oscillation, or chugging loads occur after complete reactor depressurization, t

/^T

\,,)

4.1.2-1 Revision 0 1

NEDO-21888 4.1.2.1 Analytical Procedure The pressure and temperature response of the containment to the postulated O

IBA was determined by performing a mass / energy balance at the following times:

a. When the vents are half cleared

b. When the vents are fully cleared

c. When half the drywell air has been purged to the wetwell

d. When all the drywell air has been purged *.o the wetwell

e. When the feedwater system is manually shut off at 10 minutes after the accident (for plants with motor driven feedwater pumps only). For plants with turbine driven feedwater pumps, the pumps are assumed to coast down in 7 seconds following reactor isolation; the pressures and temperatures for this point in time are not calculated.

f. Just prior to ADS initiation due to Jow reactor water level and high drywell pressure

g. When the reactor vessel is depressurized to equalize with the drywell pressure (approximately 50 psia) 4.1.2.2 Assumptions

a. The fuel decay heat and sensible energy used for the DBA analysis shown in Table 4.1.1-1 are used for the IBA analysis.

b. A liquid break is assumed with an area equal to 0.1 ft2 The blow-down model is the same as the Homog.aneous Equilibrium model used in the DBA analyses (Reference 4.1.2-1).

O 4.1.2-2 Revision 0

c. Feedwater flow coastdown is modeled for plants with turbine (D

s_si driven feedwater pumps. The feedwater flow is manually stopped-at 10 minutes af ter the break, for plants with motor driven feed-water pumps. -

d. Fluid subcooling is conservatively neglected.

e. The steam is assumed to stratify above the air in the drywell, which gives a higher rate of pressurization of the wetwell.

f. .The bulk wetwell pool and airspace temperatures are equal throughout the transient. This maximizes the wetwell airspace temperature and pressure,

g. The hydrodynamic effects of vent clearing are neglected.

h. A constant liquid flashing rate for the break flow is used.

(~'y 'i. The condensation of steam on drywell walls and structures is con-V servatively neglected.

8

j. A single failure of the HPCI (or FWCI where applicable) system is assumed.

For conservatism, the containment response has been determined assuming l initial conditions which yield the most severe pressure and temperature in the containment. The conditions are as follows:

a. The reactor is operating a 102% of licensed thermal power as identified in Table 4.1.1-3 for a typical plant. This maximizes the core decay heat.

,O.

~

4.1.2-3 Revision 0

b. The temperature of the suppression pool is equal to the maximum technical specification limit for normal operation. This h maximizes the suppression pool temperature and the wetwell airspace temperature and pressure.

c. No credit is taken for normal auxiliary power, which maximizes the time it takes the ECCS system to be operable.

Other plant conditions and assumptions are as described for the DBA in Section 4.1.1.

4.1.2.3 Analysis Results Plant unique containment temperature and pressure responses have been cal-culated and are presented in the Plant Unique Load Definition documents and submitted individually to the Mark I Owners. Typical responses are shcwn in Figures 4.1.2-1 and 4.1.2-2.

O O

4.1.2-4 Revision 0 l l

l

O O O 4

400 -

DRYWE LL AIR PURGED TO WETWELL - 18) VESSE L DEPRESSURIZED - (D)

,it 300 - l l

[ ( Al - VENT CLE ARING l INITIATION OF ADS-(C)

I '

$ l t-3 e

I i

l lll M 8

I Y

$ 2w -l 1

l lI i b

=

WETWELL 00 I

I I i I

I ioo -l i

l lil, l l

1 I

! l I l !

l 8

i io ion two so m g TIME AFTER BREAK INITIATION (sec) 0 D Figure 4.1.2-1. Typical Containment Temperature Response to an o Intermediate Break Accident

50 (D) - VESSE L DLPH ESSUHlZED

,o _

INITIATION OF ADS -(C) l I DRYWELL iI WETWE LL DRYWELL AlH PURGED TO WETWELL- 18) 1 I 4

x -

v i l

s i i

e I I 8

a a.

I Il lI E

8, N g i

h ~

H I

I I

( A) - VENT CLE ARING f 10

I I

I O I !

1 10 100 1000 10.000 g TIME AFTER BREAK INITIATaOp (sec)

E

^

E y Figure 4.1.2-2. Typical Containment Pressure Response to an D Intermediate Break Accident o

4 O O O

NEDO-21888 REFERENCES FOR SECTION 4.1.2 Ch V

4.1.2-1 F. J. Moody, " Maximum Discharge Rate of Liquid-Vapor Mixtures from Vessels," General Electric Company, Report No. NED0-21052, September 1975.

O l

l lO 4.1.2-7/4.1.2-8 Revision 0

%.. 3 Small Break Accident Typical containment temperature and pressure responses to a SBA are shown in Figures 4.1.3-1 and 4.1.3-2, respectively. Following the break, the drywell pressure will increase slowly and the water level in the vents will be slowly depressed until at point .'. in Figure 4.1.3-2, drywell air and steam pass through the suppression pcol where the steam is condensed and the air will rise to the free airspace above the pool. At point B, the air is purged from the drywell to the wetwell, and at point C (10 minutes) the operator manually initiates ADS. At point D, just prior to low pressure ECCS initia-tion, the wetwell airspace volume has been reduced to its minimum value and the maximum wetwell pressure occurs. At approximately 20 minutes after the occurrence of the break, the reactor will be depressurized at point E. Fol-lowing reactor depressurization and ECCS flooding of the vessel, the sequence of events is the same as for the IBA described in Section 4.1.2.

During the depressurization of the reactor vessel, the suppression pool temp-erature increases rapidly due to the transfer of stored reactor vessel energy to the pool. The suppression pool temperature transient will level out after the reactor vessel is depressurize'd. The pool temperature will rise much less rapidly since the heat input to the pool is limited to the core decay heat and the RHR system is removing heat from the pool. No S/RV, condensation oscillation, or chugging loads occur after complete reactor depressurization.

4.1.3.1 Analytical Procedure The temperature and pressure response of the containment to the SBA has been I

determined by performing a mass / energy balance on the reactor vessel to obtain '

the reactor depressurization and temperature histories, and also by performing i

a mass / energy balance to obtain containment pressures and temperatures at the '

following times:

a. When the vents are half cleared.

i

b. When the vents are fully cleared p

l O 4.1.3-1 Revision 0

c. When half the drywell air has been purged to the wetwell

d. When all the drywell air has been purged to the wetwell 9

e. When the reactor vessel is completely depressurized.

4.1.3.2 Asaumptions

a. The fuel decay heat and sensible energy used for the DBA and IBA analyses shown in Table 4.1.1-1 are used for the SBA analysis,

b. A steam break is assumed with an area equal to 0.01 ft2, The blowdown model is the same as the Homogeneous Equilibrium Model used in the DBA and IBA analyses (Reference 4.1.3-1),

c. Feedwater flow coastdown is modeled for plants with turbine driven feedwater pumps. The feedwater flow is manually stopped at 10 minutes after the break, for plants with motor driven feed-water pumps.

d. Stratification of the steam above the air in the drywell is assumed. This maximizes the pressurization rate of the wetwell.

e. The bulk wetwell pool and airspace temperatures are equal through-out the transient. This maximizes the wetwell airspace temperatures and pressures.

f. The hydrodynamic effects of vent clearing are neglected.

g. The condensation of steam on drywell walls and structures is neglected.

h. A single failure of one RHR pool cooling loop is assumed.

O 4.1.3-2 Revision 0

i NEDO-21888 For conservatism, the containment response has been' determined assuming condi-

' ( ,) tions at the time of the break which yield the most severe pressure and temp-erature in the containment. The conditions are as follows:

a. The reactor is operating at 102% of licensed thermal powe as identified in Table 4.1.1-3 for a typical plant. This maxit',es the core decay heat.

b. No credit is taken for normal suxiliary power. This maximizes the time it takes the ECCS system to be operable. Other assumptions 4

and plant conditions at the instant of the break are as described for the DBA except the initial temperature of the suppression pool, i which is as described fir the IBA.

4.1.3.3 Analysis Results Plant unique containment airspace temperature and pressure responses are presented in the Plant Unique Load Definition documents and submitted

() individually to the }brk I Owners. Typical responses are shown in Figures 4.1.3-1 and 4.1.3-2.

O I 4.1.3-3/4.1.3-4 Revision 0 I

', , -. .. . . - , . .- ., . . -. .- . -. . - - - , ,- -- \

O O O

~

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H INITIATION OF ADS (600 seconds) - (C) li

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I h

- l l l 8 y VENT CLEARING - (A) 1 i 4 a

y I II :

300 _

WETWELL l l l

I ll II l

0 1 10 100 1000 10,000 TIME AFTER BHE AK INITI ATION (sec)

?

i li 0 Figure 4.1.3-1. Typical Containment Temperature Response to a o Small Break Accident

SO 40 -

MINIMUM WETWELL AIRSPACE VOLUME -(D)

DE RESSUHlZtD I I tNITI ATION OF ADS (600 sin:onds) - (C) l

~

DRYWELL air PURGED TO WETWELL -(8) 3 l

- l l b z

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WETWELL I

s' l I 0 l 1 1 10 100 1000 10,000 p3 TIME AFTER BREAK INITIATION (sec)

Figure 4.1.3-2. Typical Containment Pressure Response to a l o Small Break Accident I

e G #

NEDO-21888 REFERENCES FOR SECTION 4.1.3 O F. J. Moody, " Maximum Discharge Rate of Liquid-Vapor Mixtures 4.1.3-1 from Vessels," General Electric Company, Report No. NEDO-21052, September 1975.

O o

a 4.1.3-7/4.1.3-8 Revision 0

--- , , , . ,- . -, ,--,y , - ,,

NEDO-21888 4.2 VENT SYSTEM THRUST LOADS

/~T U

The postulated DaA causes the most rapid pressurization of the containment sys-tem, the largest vent system mass flow rate, a..? therefore, the most severe vent system thrust loads. The pressurization of the containment for the IBA and SBA is less rapid than for the DBA; thus, the resulting vent system thrust loads for the SBA and IBA are less severe. Consequently, vent system thrust loads are presented for the DBA only.

Following the pipe break, the drywell pressure increases and the water initially occupying the submerged portion of the vent system is accelerated into the pool, clearing the vents. The structural loading during the vent clearing process is due to pressure imbalances between the increaslag pressure in the vent sys-tem and the surrounding wetwell airspace. Momentum need not be considered since flow velocities during the vent clearing process are low. Following vent clearing, there is a period during which bubble formation and breakthrough occurs. This period is characterized by increasing flow velocities and associated pressure losses in the vent system. Following breakthrough, mass

/"'T flow rates and velocities in the vent system become significant, and momentum V must also be considered in determining thrust loading. The fluid flowing through the vent system following breakthrough is assumed to be a homogeneous mixture of air, steam, and water.

Forces on the vent system are derived from mass and momentum equations.

The sign convention used to define positive forces on a typical vent system is shown in Figure 4.2-1. Vent system thrust loads for this typical plant are shown in Figures 4.2-2 through 4.2-5 for 0 to 5 seconde, and Figures 4.2-6 through 4.2-9 for 0 to 30 seconds. Nomenclature for chis section is given in Table 4.2-1.

Due to variations in vent system configurations, the nomenclature presented in Table 4.2-1 defines additional vent system forces and angles which are not applicable to the typical plant vent system geometry presented in Figure 4.'..-l.

The vent system thrust loads presented in the Plant Unique Load Definition documents and-submitted to the individual Mark I owners, contain figures and 4.2-1 Revision 0

NEDO-21888 legends which define the forces applicable to the plant unique vent system configuration.

O The force transients presented in Figures 4.2-2 through 4.2-9 show distinct changes in slope or inflection points. These are clearly illustrated in Figure 4.2-3 and are due to the large changes in flow conditions during the transient, as well as the pressure changes for the drywell and wetwell as determined for the DBA in Section 4.1.1. Up to the time of vent clearing, the absolute magnitude of the forces increases due to rising drywell pressure (assumed constant throughout the vent system) acting against constant wetwell airspace pressure. Following vent clearing, the absolute magnitude of the force on the main vents continues to increase since drywell pressure has not yet peaked and pressure losses from the drywell to the main vents are fairly small. Forces on the vent header and downcomers, however, are characterized by a decrease in magnitude due to large entrance and exit pressure losses between the components when the vent system flow rate has significantly increased. Figures 4.2-5 and 4.2-9 present the forces resulting from all main vents and downcomers as well as a " net" vertical force on the entire vent system resulting from the sum of all component contributions. These figures do not show any " net" horizontal force on the entire vent system since the sum of all horizontal forces is zero due to the symmetrical configuration of g

the vent system.

The vent system is subjected to internal pressures as shown in Figures 4.2-10 and 4.2-11 for a typical plant. These pressure histories correspond to the force transients shown in Figures 4.2-2 through 4.2-9. As stated earlier, vent system pressures are equal to the drywell pressure during the vent clearing process.

O 4.2-2 Revision 0

- 4.2.1 Analytical Procedure The analytical procedure employs thrust equations which consider the forces due to both pressure distributions and momentum to define horizontal and vertical thrust forces on the following components of the vent system as identified in Figure 4.2-1:

a. Main vents

b. Vent header

c. Downcomers.

i Because both the main vents and the downcomers are located symmetrically about the center of the vent system, the horizontal vent system thrust loads cancel each other resulting in no net effective horizontal vent systes thrust load.

The thrust calculations are based upon plant unique containment pressure

() response and vent system mass flow rates from analyses similar to those described in Section 4.1.1 for the DBA, but are performed at both minimum and zero initial pressure differential between the drywell and wetwell. The thrust equations define loading on the vent system components during four distinct time periods:

a. Before Vent Clearing The entire vent system (excluding the portion of the downcomers containing water) is assumed to be at drywell pressure. Thrust forces result from pressure in the vent system and the surrounding wetwell airspace pressure, as determined for the DBA in Section 4.1.1.

The following equations were used to calculate the horizontal and vertical forces on a single main vent, vent header (per mitre bend)

O 4.2.1-1 Revision 0

NEDO-21888 and a sf ,le downcomer. The nomenclature for terms used below are listed . Table 4.2-1. Depending on the vent system configuration, not all equations listed below are necessarily used to define vent system thrust forces:

FlV1 =

(PW - PDW) (Ayp/n y ) (sin Oy )

FlH1 =

(PW - PDW) (Ayp/ny ) (cos Oy)

FlV2 =

(PW - PDW) (Ayp/n y ) (sin a - sin Oy )

FlH2 =

(PW - PDW) (Ayp/n y ) (cos a - cos Oy )

F2V =

(PDW - PW) (bC"3) (sin 0 2 F2H =

(PDW - PWW) (AVH) (2 - 2 sin S)

F3V =

(PDW - PW) (ADC "2) (sin3 0 - sin 02)

F3H =

(PDW - PW) (ADC "2) (cos30 - cos 02 }

F4V =

(PDW - PW) (ADC "2) (1 - sin 03)

F4H =

(- c s 0 3)

(PDW - PW) (ADC "2)

b. At Vent Clearing At vent clearing all water initially in the downcomer has been cleared. Since the bubbles at the end of U downcomers are just starting to grow, the flow velocities are sufficiently low such that momentum effects can be ignored. Thrust forces are due to imbalances between drywell and wetwell airspace pressures as obtained from the DBA analyses described in Section 4.1.1. The equations used to calculate the forces are the same as those pre-sented in item (a) above.

O 4.2.1-2 Revision 0

NEDO-21888 f3 Af ter Vent Clearing and Before Bubble Breakthrough

() c.

Af ter the vents have cleared, bubble formation occurs at each down-comer exit. Since the analytical methods described in Section 4.1.1 do not account for bubble formation, mass flow rates predicted by this method during bubble formation are not sufficiently descriptive for thrust force calculations. Therefore, plant unique QSTF experi-mental data (Reference 4.2-1) were examined to determine the shape of the force curve. The results of this analysis showed that a straight line approximation of the force curve from the time of vent clearing to the time of bubble breakthrough, when flow rate information obtained from the Section 4.1.1 analyses is applicable, would be conservative since an exponential decay of forces was observed in the QSTF. The time period from vent clearing to bubble breakthrough was conservatively chosen as 200 milliseconds. The straight line approximation is shown between points A and B in Figures 4.2-2 through 4.2-5 and Figure 4.2-10.

The equations used to calculate the horizontal and vertical forces on a single main vent, vent header (per mitre bend) and a single 1

downcomer at the time of bubble breakthrough are presented below.

The mass flow rates plus drywell and wetwell pressures, as deter-mined for the DBA in Section 4.1.1,were used in the calculations at the time of bubble breakthrough. The nomenclature is listed in Table 4.2-1.

FlV1 = ((PW-P1) (Ayp/ny ) - (s Ty V )/(n gye sin e y FlH1 = ((PW-P1) (Ayp/n g ) - (6 T1 V )/("lE c)) e s 07 F1V2 = ((PW-P1) (Ayp/n g ) - (&

T1 V )/("1E c} - (sin a - sin Oy )

FlH2 = ((PW-P1) (Ayp/ng ) - (AT1 V )/("l Ec)) (e s a - cos Oy )

F2V =

((P3-PW) (ADC! "3 + 3 ("3Ec 8 "

2 4.2.1-3 Revision 0

NED0-21888 F2H = ((P2-PW) AVH + ( T V 2)/("38 c)) (2 - 2 sin S) g F3V =

((P3-PW) (bC "2 +( 3 ("2c))(sin 8

03 - sin 0 2)

/

F3H = ((P3-PW) (ADC"2)+(bV3 ("28c (cos 3 0 -cs0) 2 F4V = ((P3-PW) (A /

DC"2)+(bY) 3 ("28 c)) (1 - sin 03)

F4H =

((P3-PW) (ADC"2)*(kY3 ("28c (- c s 03)

d. After_ Breakthrough Following bubble formation and breakthrough, full flow through the vent system is established. During this period, fluid romentum becomes a significant term in the thrust force calculation. Mass flow rates and drywell and wetwell pressures, as detendned fer the DBA in Section 4.1.1, were used in the calculations ror the time period following bubble breakthrough. The equations used to h

calculate the forces are the same as those presented in item (c) above.

O 4.2.1-4 Revision 0

NEDO-21888 g

( 4.2.2 Assumptions

a. Flow losses in the vent system following breakthrough have been selected to yield conservatively high pressures in the main vent, vent header and downcomers, thus resulting in conservatively high thrust forces. The following flow loss distribution is typical for all Mark I plants and is representative of calculated local loss coefficients:

(1) Fressure drop from the drywell to the main vent is 25% of the total calculated pressure drop in the vent system.

(2) Pressure drop from the main vent to the vent header is 37.5%

of the total calculated pressure drop in the vent system.

(3) Pressure drop from the vent header to the downcomer is 27.5%

of the total pressure drop in the vent system.

m u-) In other words, at any point in time the main vent is at a pressure equal to the drywell pressure minus 25% of the total pressure drop in the system, the vent header is at drywell pres-sure minus (25% + 37.5%) of the pressure drop, and so on. This assumes that losses are distributed at entrances and exits of the components. This is reasonable since form and friction losses in the vent system components themselves are small.

It should be noted that the total pressure drop from the drywell to the downcomer is assumed to be 90% [ sum of items (1), (2), and (3) above] of the total pressure drop from the drywell to the wetwell.

The remaining 10% is assigned as exit loss from the downcomer to the suppression pool.

b. Flow in the vent system is completely turbulent and loss coefficients are independent of Reynolds number.

O)

\_

4.2.2-1 Revision 0

-- - y

NED0-21888

c. The flow of liquid, steam, and air following vent clearing is g conservatively assumed to be homogeneous,

d. Momentum effects are negligible prior to vent clearing because fluid velocities are very low during this period.

e. Unsteady fluid flow considerations are negligible following vent clearing because they have a second order effect on the calculated thrust loads.

f. Fluid flow from the drywell to the vent system is assumed uniform.

The effect of nonuniform flow in the downcomers observed in the EPRI 1/12 scale pool swell tests (References 4.2-2 and 4.2-3) is discussed in Section 4.2.4 I

l O

4.2.2-2 Revision 0

NED0-21888 q\' '>

4.2.3 Analysis Results Plant unique force transients and containment pressure responses have been generated for the following two sets of initial containment conditions:

a. Maximum submergence and minimum normal operating pressure differential between the drywell and the wetwell

b. Maximum submergence ano e';o pressure differential between the drywell and the wetwell.

The results are presented in the Plant Unique Load Definition documents

~

and submitted to the individual Mark I Owners. Results for a typical plant are presented in Figures 4.2-2 through 4.2-10.

The vent system thrust loads are calculated and presented for the initial 30 second period following the DBA. Thrust loads after 30 seconds stay virtually

() constant at the value calculated at 30 seconds.

Some tbrk I plants normally operate with zero pressure differential between the drywell and the wetwell and therefore, only one set of results are presented for these plants.

b) v 4.2.3-1/4.2.3-2 Revision 0

'l

f NEDO-21888

(} 4.2.4 Application The horizontal and vertical main vent thrust transients shown in Figures 4.2-2 and 4.2-6 represent the resolution of the thrust load which is acting on the end cap of the main vent. This loading is actually distributed over the end cap area, and is shown in Figure 4.2-12. This figure also shows main vent mitre bend forces for the two plants (Browns Ferry and Oyster Creek) which have this configuration. It should be noted that the Browns Ferry I

main vent mitre bend lies between the drywell and the bellows on the main vent. The Oyster Creek main vent mitre bend lies on a continuous run of piping (i.e., downstream of the bellows).

4

'The vertical and horizontal vent header thrust transients shown in Figures 4

4.2-3 and 4.2-7 represent the vent header loading per mitre bend. Vertical loading is due to the contributions of the individual downcomer pairs, which were assumed equal for each pair in the analysis. As discussed in Section 4.2.2 the flow in the downcomers is not uniform. Figure 4.2-13 defines the effect of this nonuniform flow.

[)

The horizontal and vertical downcomer thrust transients shown in Figures 4.2-4 l- and 4.2-8 are loads for a single downcomer, and are the resultant forces due I to turning the flow through the mitre bend. The resolution of this load into i vertical and horizontal point loads is shown in Figure 4.2-14.

Figure 4.2-14 also shows the mitre bend loading for those plants which have two downcomer mitre bends. Duane Arnold, which is the only plant with straight vertical down-comers, has no downcomer forces. The discussion pertaining to nonuniform downcomer flow on vent header forces also pertains to the . forces generated for downcomers. Figure 4.2-14 defines the effect of this distributed flow.

In addition, total vertical thrust loads and net vertical thrust loads, shown in Figures 4.2-5 and 4.2-9, are defined as follows:

F1VIT = Main vent end cap vertical force multiplied by the number of main vents f%

U 4.2.4-1 Revision 0

NEDO-21888 FlV2T = Main vent mitre bend vertical force multiplied by the number of main vents (Browns Ferry and Oyster Creek only)

F2VT = Vent header vertical force per mitre bend multiplied by the number of vent header mitre bends F3VT = Downcomer mitre bend vertical force multiplied by the number of downcomers F4VT = Second downcomer mitre bend vertical force multiplied by the number of downcomers FNETV = FlVIT + F1V2T + F2VT + F3VT + F4VT For plants which do not have main vent mitre bends, or less than two downcomer mitre bends, forces which apply to these bends will be zero.

O O

4.2.4-2 Revision 0

NEDO-21888 Table 4.2-1 NOMENCLATURE M R SECTION 4.2 PDW Drywell pressure PWR Wetwell airspace pressure Pl Main vent pressure P2 Vent header pressure P3 Downcomer pressure FlV1 Vertical force on a single main vent end cap FlH1 Horizontal force on e single main vent end cap FlV2 Vertical force on a single main vent mitre bend (applicable to Browns Ferry and Oyster Creek only)

FlH2 Horizontal force on a single main vent mitre bend (applicable to Browns Ferry and Oyster Creek only)

F2V Vertical force on vent header (per mitre bend)

F2H Horizontal force on vent header (per mitre bend)

F3V Vertical force on a single downcomer mitre bend F3H Horizontal force on a single downcomer mitre bend F4V Vertical force on second mitre bend of downcomer (if applicable)

F4H Horizontal force on second mitre bend of downcomer (if applicable)

FlVIT Total main vent end cap vertical force = FlV1 x ,

number of main vents l FlV2T Total main vent mitre bend vertical force = FlV2 x number of main vents F2VT Total vent header vertical force = F2V x number of vent header mitre bends F3VT Total vertical force (first downcomer .ntre bend) =

F3V x number of downcomers 4.2.4-3 Revision 0

NEDO-21888 Table 4.2-1 (Continued) g NOMENCLATURE FOR SECTION 4.2 (Continued)

F4VT Total vertical force (second downcomer mitre bend) =

F4V x number of downcomers FNETV FNETV = FlVlT + F1V2T + F2VT + F3VT + F4VT g Vent header flow area Ag Total main vent flow area A Total downcomer flow area DC ny Number of main vents n Number of doutcomers 2

n Number of vent header mitre bends 3

Total mass flow rate V Fluid velocity in main vent V Fluid velocity in vent header 2

V Fluid velocity in downcomer 3

6 Angle of main vent with horizontal 7

0 Angle of first downcomer mitre bend with horizontal 2

0 Angle f sec nd downcomer mitre bend with horizontal 3

a Angle of main vent mitre bend with horizontal S 90* - (vent header mitre bend angle)

O 4.2.4-4 Revision 0

NEDO-21888 1

(

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PLAN F1V1 h ,

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% .2 F3H SECTION A-A F1VI

VERTICAL FORCE ON MAIN VENT END CAP F1H1 =

HORIZONTAL FORCE ON MAIN VENT END CAP F2V =

VERTICAL FORCE ON VENT HEADER (PER MITRE BEND)

F2H =

HORIZONTAL FORCE ON VENT HEADER (PER MITRE BEND)

F3V =

VERTICAL FORCE ON DOWNCOMER MITRE BEND F3H =

HORIZONTAL FORCE ON DOWNCOMER MITRE BEND Figure 4.2-1. Definition of Positive Thrust Loads sJ 4.2.4-5 Revision 0 l

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NEDO-21888 9

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i f

s y F1H1 DISTRIBUTED PRESSURE LOAD ON MAIN VENT END CAP IS RESOLVED INTO VERTICAL AND HORIZONTAL

POINT

LOADS.

FOR PLANTS WITH MAIN VENT MITRE BENDS (BROWNS FERRY AND OYSTER CREEK)

O F1V1 .F1V2 4k

%F1H2

' ; E1H1 F1V2 AND F1H2 ARE ACTUALLY DISTRIBUTED AROUND THE BEND AND ARE THE RESULTANTS OF INTERNAL PRESSURE AND SHEAR FORCES ON THE PIPING Figure 4.2-12. Application of Thrust Force on Main Vent End Cap and Main Vent Mitre Bend I

l 4.2.4-16 Revision 0

NEDO-21888 F2Vj (F2 V)1,2 (F2V)3,4 (F2V)5,6 n , d d 4 1 %

n+'

~' f, f h_4 N 1

I V

\ -

~

s

' i (l l

/

A

(/ (/

A l

s ' td 1 2

F2H - - -

PLAN SECTION F2V IS THE VENT HEADER VERTICAL FORCE PER MITRE BEND AND REPRESENTS THE CONTR'IBUTIONS OF DOWNCOMERS 1 THROUGH 6

)

LET (F2V)1 BE THE CONTRIBUTION TO F2V FROM DOWNCOMER 1 (F2V)2BE THE CONTRIBUTION TO F2V FROM DOWNCOMER 2 (F V)6 BE THE CONTRIBUTION TO F2V FROM DOWNCOMER 6 FROM THE ANALYSIS, (F2V)1 = (F2V)2 = (F2V)3 = (F2V)4 = (F2V)$ = (F2V)6 = F2V/6 HOWEVER, FROM EPRI TESTING (P.FERENCE 4.2 3), I (F2V11 = (F2V/6) (0.79); (F2V)4 = (F2Vi6) (0.95)

(F2V)2 = (F2V/6) (0.9); (F2V)5 = (F2V/6) 11.19)

(F2V)3 = (F2Vi6) (1.1); (F2V)6 = (F2V/6) (1,17)

THESE FACTORS REPRESENT THE MAXIMUM DEVfATION FROM THE UNIFORM FLOW ASSUMPTION AND APPLY ONLY AFTER VENT CLEARING. AS FLOW RATE DECREASES THE ABOVE FACTORS APPROACH 1.0 l Figure 4.2-13. Application of Vent Header Forces

O(/

4.2.4-17 Revision 0

I NEIto-21888 l O

VENT HE ADER F3V j F3V 16 g : F3H :.F3H F4V n

DOWNCOMER SINGLE MITRE BEND DOUBLE MlTRE BEND VENT HE ADER

~

5' 6

2

I

/, h a 2' TOP ,-

VIEW f i

I 5

( 1 3 DOWNCOMER IN THE ANALYSIS F3V, F3H, F4V, AND F4H ARE ASSUMED EQUAL FOR DOWNCOMERS 1 THROUGH 6.

HOWEVER, FROM EPRI TESTING (REFERENCE 4 2 3).THE FOLLOWING FACTORS APPLY AFTER VENT CLEARING:

(F3V)g = F3V X 0.79; (F3V)4 = F3V X 0 95 (F3V)2 - F3V X 0.90; (F3V)$ = F3V X 1.18 (F3V)3 = F3V X 1.01; IF3V)6 = F3V X l.17 THE SAME F ACTORS WOULD APPLY FOR F3H, F4V, AND F4H ll.E. (F4H)1 = F4H X 0 7941 Figure 4.2-14. Application of Downcomer Forces 4.2.4-18 Revision 0

l NEDO-21888 REFERENCES FOR SECTION 4.2 2 4.2-1; J..M. Humphrey, " Mark I Containment Program 1/4 Scale Two-Dimensional Plant Unique Pool Swell Test Report," General Electric Company, Report No. NEDE-21944-P, to be published.

.4.2-2 J. E. Torbeck, et al., " Mark I Containment Program 1/12 Scale Pressure

] Suppression Pool Swell. Tests," General Electric Company, Report No. NEDE-13456-P, March 1976.

4

-4.2-3 The Electric Power Research Institute, "Three Dimf.nsional Pool Swell

'Modeling of a Mark I Suppression System," EPPI, NP-906, October 1978.

O 1

4 5

i i

t

?

i

r. _s 4.2.4-19/4.2.4-20 Revision O.

NEDO-21888 4

_ f ,) 4.3 POOL SWELL LOADS

. s_

In the event of a postulated design basis LOCA due to a break in a large pressurized line, the drywell and vent system are pressurized and the water initially in the downcomers is accelerated downward into the suppression

. pool. After the downcomers are cleared of water, air is discharged into the wetwell below the pool surface and the pool swell transient tarts.

The vent system air is initially at dryvell pressure. When the design basis LOCA occurs, the pressure rapidly increases in the drywell and vent system, which results in the water leg in the downcomers being pushed out into the suppression pool. When this clearing process is completed, the air behind the downcomer water slug produces a bubble at the end of the downcomer. The water above the bubble is accelerated upward as the bubble expands. As the bubble expansion continues, the pool water rises in the torus and compresses the airspace above the pool surface. Eventually, the bubble " breaks through" to the torus airspace and the displaced poal liquid settles back to its fs original level.

U Associated with pool swell are the following eight primary loads:

a. Torus net vertical loadu

b. Torus shell pressu es

c. Vent system impact and drag loads

d. Impact and drag loads on other structures above the pool

e. Froth impingement loads

f. Pool fallback load

g. Downcomer water clearing jet load on submerged structures

h. LOCA bubble induced drag loads on submerged structures G

N.,] '

4.3-1/4.3-2 Revision 0

NEDO-21888 l'M

.( j 4.3.1 Torus Net Vertical Load Histories In the postulated LOCA, the downcomer air, which is at essentially drywell pressure, is injected into-the suppression pool, producing a downward reaction force on the torus. The consequent bubble expansion causes the pool water to swell in the torus, compressing the airspace above the pool and producing an upward reaction force on the torus. The bubble pressure decreases as the bubble overexpands and the pool liquid mass decelerates. Eventually, the bubbles " break through" to the torus airspace (the bubble and torus airspace com-municate) and the displaced pool liquid settles back toward its original level.

These vertical loads create a dynamic imbalance of ferces on the torus, which acts in addition to the weight of the water applied to the torus.

This dynamic force history lasts for only a few seconds and typically includes a downforce shortly af ter vent clearing, followed by an upforce as the airspace is compressed.

/^s

\J A typical force history is shown in Figure 4.3.1-1.

4. 3.1.1 Bases and Assumptions The torus net vertical load transients are based on 1/4 scale 2D plant unique test results performed in the QSTF (Ref. 4.3.1-4). The geometric characteristics of the torus and vent system and he LOCA driving conditions were represented in this facility for each plant csted. The test facility design and operation is based on a scaling law 41ysis which is presented and verified in Reference 4.3.1-1.

For each plant unique series of tests, the plant short term DBA drywell pres-surization history was calculated using an analytical model of the drywell response as described in Section 4.1.1. The resulting pressure history (scaled for the test) was used as a lower bound for the test drywell pressuri-zation history, assuring a conservative load definition. Figure 4.3.1-2 v

4.3.1-1 Revision 0 t

NED0-21888 presents a comparison of an analytically determined drywell response to the measured response for a representative QSTF test. The QSTF tests were per-formed at the minimum plant operating AP and maximum downcomer submergence for the plant. These conditions will result in a conservative prediction of the net torus vertical loads. The plant unique test conditions are included in Reference 4.3.1-4.

As indicated in References 4.3.1-1 and 4.3.1-4, orifices were placed in the 1/4 scale vent system in order to accurately model the vent system resistance (Section 4.1.1 specifies the plant unique flow resistances used). Reference 4.3.1-1 describes the scaling laws used to obtain flow resistances (fL/D) which produced correctly scaled enthalpy flow out of the downcomers. Scaling was based on the following considerations:

a. As identified in Reference 4.3.1-3 the flow resistance of the 1/4 scale vent system was obtained by scaling the plant unique flow resistances (fL/D) by the quantity 1/(length ratio of the model) to satisfy the enthalpy flow scaling law.

b. Additionally, the vent flow resistance was decreased to account for the fact that the initial air temperature in the drywell for the 1/4 scale tests was 70*F, while the plant drywell operating temperature is higher. In order to obtain the correctly scaled enthalpy flow, the QSTF flow resistance had to be decreased (thereby increasing the QSTF enthalpy flow) by the inverse ratio of the actual plant and QSTF drywell temperatures. For example, for a drywell operating temperature of 135'F, the adjustment factor to 460 + 70 the vent flow resistance is = .8% , resulting in the 460 + 135 correct enthalpy flow to the bubbles.

c. As discussed in Reference 4.3.1-4 the vent system flow resistance was evenly split by placing orifices at the top of the vent pipe and in the downconers. This configuration is a reasonable approxi-mation of the continuous flow loss distribution in the actual Mark I vent s:? tem.

O 4.3.1-2 Revision 0

/\s-) Figure 4.3.1-3 comptres the enthalpy flow measured in a typical 1/4 scale test and a calculated en:halpy flow once quasi steady flow has been obtained. The calculated enthalpy flow was found using the drywell pressure history calculated as described in Section 4.1.1 and assuming steady state Fanno flow of 100% air through the vent system. Figure 4.3.1-3 demonstrates that the quasi steady state enthalpy flow achieved in the test exceeds the calculated enthalpy flow and is therefore conservative. The enthalpy flow produced in the test facility is also censervative prior to reaching its quasi steady state flow condition due to the conservatism of the drywell pressure history (see Figure 4.3.1-2) and the selection of the vent system loss coefficient for the QSTF tests.

For all QSTF tests the test conditions were set to correspond to the plant operat-ing conditions which produce the most severe loading. For example, increased dry-well to wetwell aP decreases the torus vertical loads and vent header impact loads. The QSTF tests were performed with the minimum plant operating AP as the maximum test AP. Other parameters treated in a conservative manner include downcomer submergence, torus water level, drywell pressurization, and vent gs system flow resistance. In addition, the geometric tolerances of the facility

\-- were established such that any effect on the measured loads would be con-s e rvative. The load definitions aff'ected by these conservatisms include those discussed in this section and Sections 4.3.2 and 4.3.3.

The net torus load histories are based on a spatial integration of the pres-sure transducers located on the QSTF torus shell. Reference 4.3.1-2 presents net torus load data from the first series of QSTF tests as measured by both the pressure transducer integration and the torus load cell. This data con-firms that the pressure transducer integration accurately measures the net l torus vertical load.

j

(~

~

i k)).

4.3.1-3 Revision 0

NEDO-21888 The net torus loads are based on results from the QSTF, which is a 2D 9, facility with average vent spacing. Reference 4.3.1-2 addresses the possible difference between loads in a three-dimensional test f acility versus those in a two-dimensional facility. That report concludes that loads derived from two-dimensional test results do not need to be adjusted to account for three dimensional effects. Thus, the plant unique net torus load transients shown in the Plant Unique Load Definitions are appropriate for a three-dimensional load definition.

4.3.1.2 Load Definition The torus net vertical dynamic loads are given on a plant unique basis as load histories. Static loads (i.e., water and structural weights) are not included in these load his tories. The net dynamic load is given as an equivalent pressure acting on the projected plan area of the torus. In the Plant Unique Load Definitions, load transients are given for each Mark I plant. For those plants whose normal operating condition includes a drywell/

wetwell AP, load transien es are given for both the zero and actual AP conditions. The peak net torus vertical loads are listed in Tables 4.3.1-1 and 4.3.1-2.

The vertical load transients given in the Plant Unique Load Definitions are mean (average) load magnitades for those cases where the testing basis included multiple tests (i.e. , normal operating conditions) . The load transient for the zero AP condition, for those plants that have an operating aP, was based upon a single test.

The use of mean values is appropriate since there are a number of con-servatisms inherent in the load definition. The calculation of the plant drywell pressurization rate includes the conservatisms outlined in Section 4.1.1.2. Further= ore, the test drywell pressurization transient always bounded the calculated transient, as illustrated in Figure 4.3.1-3. Finally, the QSTF tests were performed to match each plant's unique geometry and operating conditions and facility geometric tolerances were set to produce the most severe loading condition.

O 4.3.1-4 Revision 0

, The weight of the. prml water can be included in the net torus vertical load

-definition by-subtracting the effective water pressure from the load transient.

The effective water prestare is equal to the weight of the pool water divided by the torus horizontal projected area, i.e.,

water net, tot.( ) ~ net ( } ~ A torus i

where 4

P , (t) = mean (average) torus load definition Pnet, total (t) = t tal pool swell load on torus, including dynamic load and static water pressure, psi I

1 W = weight of pool water, lbf wa er Ag ,= torus horizontal projected area, in. .

For determination of the water mass available for resisting torus uplif t, refer to Reference 4.3.1-4.

l d

}

() I 4.3.1-5/4.3.1-6 Revision 0

NEDO-21888 Table 4.3.1-1 MEAN PEAK NET TORUS VERTICAL LOADS

OPERATING DRYWELL/WETWELL PRESSURE DIFFERENTIAL (SP)

Up Down AP Submergence

! Plant (psi) (psi) (psid) (ft) Deflector Browns Ferry 1,2,3 2.99 6.54 1.1 3.50 No Brunswick 1,2 6.14 9.75 0.0 4.33 Yes

, Cooper Station 5.50 9.12 1.0 4.38 No Dresden 2,3 4.93 7.66 1.0 4.00 Yes Duane Arnold 1.99 3.61 1.1 3.38 No Fermi 2 2.66 4.89 1.0 3.33 Yes Fitzpatrick 3.96 5.96 1.7 4.42 Yes Hatch 1 3.23 5.62 1.5 4.00 Yes Hatch 2 5.42 10.74 0.0 4.33 No O

k-) Hope Creek 1,2 3.46 7.60 0.0 3.33 No Millstone 3.54 6.12 1.2 3.33 Yes Monticello 6.44 9.91 1.0 4.25 Yes Nine Mile Point 5.06 8.17 1.0 4.25 Yes Oyster Creek 4.71 8.35 1.0 4.06 Yes Peach Bottom 2,3 4.84 8.82 1.1 4.40 Yes Pilgrim 5.57 7.61 1.5 4.80 Yes I Quad Cities 1,2 4.93 7.66 1.0 4.00 Yes Vermont Yankee 4.33 7.20 1.7 4.54 Yes

The values given are based upon initial QSTF tests and may change due to i additional tests with different operating conditions and plant modifications. '

P>

V 4.3.1-7 Revision 0

NFDO-21888 TABLE 4.3.1-2 SINGLE TEST PEAK NET TORUS VERTICAL LOADS

ZERO DRT4 ELL /'4ETJELL PRESSURE DIFFERENTIAL Up Down Submergence Plant (psi) (psi) (ft) Deflector Browns Ferry 1,2,3 3.13 10.50 3.50 No Brunswick 1,2 ** ** 4.33 Yes Cooper Station 6.70 13.18 4.38 No Dresden 2,3 5.89 11.91 4.00 Yes Duane Arnold 2.82 5.86 3.33 No Fermi 2 3.28 8.01 3.33 Yes Fitzpatrick 6.08 13.00 4.42 Yes Hatch 1 5.33 10.68 4.00 Yes Hatch 2 ** ** 4.33 No Hope Creek 1,2 ** ** 3.33 No Millstone 5.80 10.07 3.33 Yes Fbaticello 7.91 12.58 4. f7 Yes Nine Mile Point 5.86 11.75 4.25 Yes Oyster Creek 5.57 13.15 4.06 Yes Peach Bottom 2,3 6.16 10.27 4.40 Yes Pilgrim 7.44 11.94 4.80 Yes Quad Cities 5.89 11.91 4.00 Yes Vermont Yankee 7.20 15.93 4.54 Yes

The values given are based upon initial QSTF tests and =ay change due to additional tests with different operating conditions and plant modifications.

- These plants have a normal operating condition of zero drywell/wetwell pressure differential, so no single test was performed at this condition.

Their mean peak loads from a series of four tests at this condition are reported in Table 4.3.1-1.

4.3.1-3 Revision 0

NEDE-21888 O

12 8 -

PE AK UPLOAD 4 -

REDUCED UPLOAD 3

S ua O i

,$ ZERO DOWNLOAD

-8 -

PEAK DOWN LOAD 12 l I I l I I I O 200 400 600 800 1000 1200 M00 1600 TIME AFTER BREAK INITIATION (msec)

Figure 4.3.1-1.

(-}

\.s Typical Net Torus Vertical Loading History 4.3.1-9 Revision 0 i

l l . . - . - . - . - . . ..

NEDO-21888 36 0

32 -

l 3 -

TEST 24 - CAL W TED

] 20 -

U 5

5 E

l W j 16 -

i a

a

$ PEAK DOWNLOAD W

12 -

8 -

4 -

0 0 0.2 0.4 0.6 0E 1.0 1.2 TIME AFTER BREAK INITIATION (sec) l Figure 4.3.1-2. Comparison of Typical Analytical and Test Drywell Pressure Histories 4.3.1-10 Revision O

NED0-21888 36 32 -

TEST 2.8 -

CALCULATED 2.4 -

_ TIME OF PEAK UPLOAD E

c 9

$ 1.6 -

Os !

5 w

1.2 -

0.8 -

0.4 -

0 0 100 200 300 400 500 600 700 TIME AFTER VENT CLEARING (msec)

Figure 4.3.1-3. Enthalpy Flow Comparison 4.3.1-11/4.3.1-12 Revision 0

NEDO-21888 O REFERENCES FOR SECTION 4.3.1

.4.3.1-1 D. I.. Galyardt, et al., " Mark I Containment Program 1/4 Scale Pressure Suppression Pool Swell Test Program: Scaling Evaluation," General Electric Company, Report No. NEDE-21627-P, January 1978.

4.3.1-2

" Mark I Containcent Program, Comparison of Subscale Pool Swell Test Results," Prepared for the General Electric Company by Nuclear Services Corporation, Report No. NEDE-24588-P, to be published.

4.3.1-3 J. E. Torbeck, et al., " Mark I Containment Program 1/12 Scale Pressure Suppression Pool Swell Tests," General E]ectric Company, Report No.

NEDE-13456-P, March 1976.

4.3.1-4 J. M. Humphrey. " Mark I Containment Program. 1/4 Scale Two-Dimensional Plant Unique Pool Swell Test Report," General Electric Company , Report No. NEDE-21944-P, to be published.

O O 4.3.1 13/4,3,1_14 V ; Revision 0

NED0-21888 4.3.2 Torus Shell Pressure Histories The net vertical load transient on the torus during pool swell was described in Section 4.3.1. In this section the same torus load is described in terms of the local torus shell pressure history. When integrated over the torus shell inside surface, these local pressure histories result in the net torus load transient due to pool swell.

4.3.2.1 Bases and Assumptions The local torus shell pressure histories are based on test data from the QSTF and the 1/12 Scale 3D Test Facility (References 4.3.2-1 and 4.3.2-2). The loads are defined by previding an average submerged pressure history which is applicable to the submerged portion of the torus shell; a table of multipliers which specifies the variation of the average submerged transient pressure at df.fferent positions on the shell; and a pressure transient applicable to the entire torus airspace, which applies to all points in that region.

O V

The average submerged pressure transients are given on a plant unique basis and are based on the torus shell pressure transients measured during the QSTF plant unique tests. In magnitude, they are equal to the net force applied to the submerged portion of the torus during pool swell divided by the torus horizontal cross section area. The spatial variation of the average submerged pressure transient magnitude is based on both the 1/4 scale and 1/12 scale tests. The variation in the torus circumferential direction is based on the 1/4 scale 2D test results; the variation in the torus longi-tudinal direction is based on the 1/12 scale 3D test results. The spatial variation of the average submerged pressure magnitude varies with time during the pool swell transient. For this reason, the spatial variation depends not only on position on the submerged portion of the torus sheli, but also on time. Only the magnitude of the average submerged pressure transient varies at different positions on the torus shell; the timing of the pressure transient is the same at all submerged locations. Finally, the airspace pressure transients are based on the 1/4 scale tests. Representative pres-sure histories for the average submerged pressure transient and airspace pressure transient are given in Figures 4.3.2-1 and 4.3.2-2.

4.3.2-1 Revision 0

NEDO-21888 The assumptions made when determining the torus shell pressure histories are:

a. The normalized longitudinal variation of the submerged pressure history as measured in the 1/12 Scale 3D tests is applicable for all Mark I plants,

b. Effects due to the torus mitre joint are assumed to be negligible.

c. The loads derived from testing in the relatively rigid QSTF are assumed to provide a conservative load definition for the entire range of structural stiffnesses of the actual Mark I tori.

The longitudinal variation of the torus shell pressure magnitudes is primarily the result of the spacing of the downcomer pairs along the length of the torus. The downcomer spacing is not exactly the same in all of the plants, but is quite similar. Therefore, the results from the 1/12 scale tests can be applied to all the Mark I plants.

The effects of the torus mitre joints on the pool swell phenomenon are dis-cussed in Reference 4.3.2-2, where it is concluded that the effects are negligible.

4.3.2.2 Load Definition The Plant Unique Load Definitions contain the plant unique average submerged torus pressure histories and airspace pressure histories. In all these transients, the initial static pressure has been subtracted out so that only the dynamic pressure history is shown. Those figures which correspond to plant operating conditions at which multiple tests were run were con-structed such that the loads represent mean (average) loads. The figures for zero aP (for those plants whose normal operating conditions include drywell/wetwell AP) are based upon a single test.

The longitudinal and azimuthal spatial variations are implemented by means of multiplicative factors M and M .g These factors were determined by 4.3.2-2 Revision 0

NED0-21888 O

C dividing the pressure measured at a specific location by the integral of the pressure over the submerged portion of the torus shell. Figure 4.3.2-3 shows the coordinate system used in the calculation of the multipliers. The normalized length, z/4, is the torus centerline distance from the centerline of the main vent divided by the torus centerline length from the main vent to the center of the nonvent bay.

Table 4.3.2-1 contains the factors which adjust the magnitude of the average submerged pressure transient to correspond to any submerged posi-tion on the torus shell. Those factors are given for non-dimensional posi-tion on the torus shell and time during pool swell and are applicable to all Mark I plants. At points on the shell between those where the factor is defined, linear interpolation may be used to determine the appropriate factor. The submerged pressure transient changes only in magnitude at different points on the torus shell; the timing remains the same at all points. The airspace pressure transient is the same at all points on the unwetted portion of the torus shell.

The torus submerged pressure longitudinal and azimuthal multipliers are applied as follows. If, P

avg. s b(t) = The average submerged pressure magnitude at time t, P (t) = the local submerged pressure magnitude at time time t, l

M,M = the azimuthal and longitudinal submerged pressure '

multipliers corresponding to the submerged location '

and time of interest, Then, Ploc. sub(t) = Pavg. sub( } * "z

O Oq 4.3.2-3 Revision O

~

, .p, - v

NEDO-21888 For the load definitions corresponding to normal plant operating conditions, multiple tests were run in the QSTF. The submerged and airspace pressure histories defined in the Plant Unique Load Definitions correspond to the mean values of the loads measured in these multiple tests.

The submerged pressure transients defined above do not include the static pressure on the torus shell resulting from the water in the torus. The water pressure at any submerged location on the torus shell can be calcu-lated by:

P = .433 h w

where P = static water pressure at a submerged location on the torus shell, psi h = vertical head of water above submerged location on the torus shell, ft.

The static water pressure calculated by the above equation should be included in the submerged pressure transients after the transients have been adjusted to the desired spatial position on the torus shell.

Adding the static water pressure to the dynamic submerged pressure trace produces the total submerged pressure transient on the inside of the torus shell in psi relative to the initial wetwell airspace pressure. To put the total submerged pressure transient and the airspace pressure transient in terms of absolute pressure, the initial wetwell airspace pressure should be added to the total submerged and airspace pressure transients.

O 4.3.2-4 Revision 0

Table 4.3.2-1 TORUS SHELL PRESSURE HISTORY MULTIPLIERS M AND M O

j LONGITUDINAL VARIATION, M z/1 Time 0.0 0.361 0.552 0.895 1.000 Start of Accident 1.0 1.0 1.0 1.0 1.0 Peak Download 0.852 0.925 1.040 1.134 1.134 Zero Download 1.059 0.971 0.990 0.999 0.999 Peak Upload 1.109 0.992 0.996 0.945 0.945 Reduced Upload 1.077 0.979 1.005 0.966 0.966

. Zero Upload 1.0 1.0 1.0 1.0 1.0 O

AZIMUTHAL VARIATION, M O

1 165* 150' 135' 120' 90*

Time 180' 195* 210' 215' 240' 270*

Start of Accident 1.0 1.0 1.0 1.0 1.0 1.0 Peak Download 1.205 1.197 1.083 0.889 0.638 0.638 Zero Download 0.947 0.952 0.940 1.015 1.145 1.145 Peak Upload 0.908 0.935 0.940 1.023 1.186 1.186 Reduced Upload 0.963 0.983 0.972 1.010 1.067 1.067 i

Zero Upload 1.0 1.0 1.0 1.0 1.0 1.0 O

s_J i

4.3.2-5 Revision 0 l

i l

l l

"O 30 -

l l

3 l

l $"

w z

O 10

' ' ' I !

0 0 200 400 600 800 1000 1200 1400 TIME AFTER BREAK INITIATION (msac) l l

l 4 I

l Figure 4.3.2-1. Typical Average Submerged Torus Pressure History h l

1 j 4.3.2-6 Revision 0

NED0-21888 O

20 -

A w

g 20 -

U E

s O

10 -

0 '

I I I I I O 200 400 600 PA 1000 1200 1400 TIME AFTER BREAK INITIATION (msec) 1 i

1 Figure 4.3.2-2. Typical Torus Airspace Pressure Hurory I

l 4.3.2-7 Revision 0

l NED0-21888 l

9' N

n

) m n s _

i o

-->- Zlt o

O

_ o 270 -

N

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l

-l too l

l l

Figure 4.3.2-3. Submerged Location on Typical Torus Shell O

4.3.2-8 Revision 0 l

l NEDO-21888-4 REFERENCES FOR SECTION 4.3.2 1

~ 4.3.2-1 J. M. Humphrey, " Mark I Containment Program, 1/4 Scale Two-Dimensional Plant Unique' Pool Swell Test Report," General Electric Company, Report ,

I No. NEDE-21944-P, to be published.

i, i i ;

I 4.3.2-2 The Electric Power Research Institute, "Three-Dimensional Fool Swell i

Modeling of a Mark I Suppression System," EPRI NP-906, October 1978.

~

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4.3.2-9/4.3.2-10 Revision-0 i

i w y , ,,w,-- yy---,w er.. - . . w-, - - - + , ,-,.p.r -s.-4 sm--,-,2-rwm.m.,,,,,.-,.-m, .-w. m vw -.-r-,,w ._ ,- ev-w.,-.e,-e- . - , ,, .w.-n-, -e%.-e.,-..-...-------y--

(/ 4.3.3 Vent System Impact and Drag As the pool surface rises, it impacts structures in its path. The resulting loading condition of primary interest is the impact on the vent system.

The impact phenomenon consists of two events, the impact of the pool on the structure, and the drag on the structure as the pool flows past it follow-ing impact. The entire process is referred to simply as " impact" in this document; however, the load definition includes both the impact and drag portions of the loading transient.

The vent system components which are potentially impacted during pool swell include the downcomers, the vent header, and the main vents. Figure 4.3.3-1 shows the coordinate system which will be used to discuss the vent system. As for the torus vertical loads (Sections 4.3.1 and 4.3.2), only a sector of the vent system will be considered, since the same geometry repeats around the entire torus. A sector includes the vent system from the

(~'s centerline of a main vent to the position midway to the adjacent main vent.

V The sequence of impact of the pool on the vent system follows the same pattern in all the plants. The pool rises fastest where the downcomers are closely spaced (z/t = 1.0), and slowest where the downcomers are widely spaced (z/2 = 0.0). Impact first occurs on the angled portion of the downcomers and then on the vent header near z/2 = 1.0. From there the impact point moves along the vent header toward-the main vent. At any longitudinal (z/t) posi-tion on the header, impact first occurs at a particular 0 and spreads out in the e direction from that point. For plants without vent header deflectors 4

the initial impact occurs at 0 = 0*. For plants with deflectors, the initial impact is plant unique but is generally near 0 = 30*.

4 . The load definition for the vent system impact is specified in a different form for each of the three major components of the vent system.

For.the downcomers a generic impact pressure transient is specified which is assumed to apply to the entire impacted surface of the downcomer. This

( f- load definition bounds the loads for all plants.

~

4.3.3-1

NEDO-2.888 The load definition for the vent header is more complex than that for the downcomers. A number of local impact pressure transients, which correspond to the average downcomer spacing impact velocity, are defined for specific locations on the vent header. A pl. unique vent header impact velocity distribution curve forms the basis for adjusting the impact pressure transients at points on the vent header where the impact velocity is differ-ent from that of the average downcomer spacing. Finally, definitions of plant unique longitudinal and circumferential time delays specify when the local impact pressure transients occur in time relative to each other.

The impact load definition for the main vent is found by using the procedure outlined in the section on Impact and Drag on Other Structures Above the Pool, Section 4.3.4. The impact load definition for vent header deflectors is specified in Section 4.3.9.

4.3.3.1 Bases and Assumptions The plant unique 1/4 scale 2D pool swell tests (Reference 4.3.3-1) provide the primary basis for the vent system impact and drag load definition. In the 1/4 scale tests the plant unique vent header, vent header deflector (if applicable), and downcomer design and operating conditions, were accurately simu-lated. In these tests impact pressure transducers were mounted on the lower sur-face of the vent header and downcomers. These pressure transducers measured the impact pressure transient applied to the vent header and downcomers as the pool first impacted and then flowed past these structures. The 1/4 scale impact pressure transients were scaled to full scale (using Froude's Law of Com-parison which requires scaling of pressure by the geomeeric length ratio and scaling of time by the square root of the geometric length ratio) and then assigned to positions on the full scale plant header corresponding to the locations of the pressure measurements in the 1/4 scale tests.

The 1/4 scale vent header load cell measurements were used as the basis for the downcomer impact / drag load definition. The small load measured just before vent header impact was converted to an equivalent pressure on the bottom 50* included angle on the slanted portion of the downcomers adjacent to the vent header, 4.3.3-2 Revision 0

NEDO-21888 g

(,,) The 1/4 scale tests were also used as a basis for determining the plant unique vent header impact velocity and circumferential time delay. The impact velocity is necessary for determining the normalized impact velocity distribution and the longitudinal time delays.

The 1/12 scale 3D tests (Reference 4.3.3-2) were the basis for specifying the longitudinal vent header impact velocity distribution and the longi-tudinal impact time delays. The 1/12 scale tests were performed with a torus and vent system geometry of a typical Mark I plant, which resulted in a longitudinal distribution of the pool swell displacement and velocity corresponding to the typical plant , ant system configuration. This dis-tribution is applicable to all the Mark I plants. The 1/12 scale data were interpreted such that the longitudinal impact velocity distribution

and longitudinal time delays correspond to the plant unique height of the header above the initial pool surface for each plant.

The 1/4 scale 2D tests were performed at both the nominal plant operating

( conditions, and at the zero drywell/wetwell AP condition if it differed-from the nominal operating conditions. In general, multiple tests were run at the operating conditions, while a single test was run at the zero aP condition. The loads defined for the nominal plant operating conditions correspond to the mean (average) of the loads measured during the multiple tests. The zero aP condition loads correspond to the loads measured in the single zero AP test.

The following assumptions are made when defining the vent system impact and drag loads:

a. The impact pressure is proportional to impact velocity squared.

b. The duration of the impact pressure transient is inversely pro-portional to the impact velocity.

c. The plant unique timing of impact on the vent header and the varia-tion of impact velocity along the vent header z axis is that predicted

'by the 1/12 scale 3D tests.

.t-)

Q) 4.3.3-3 Revision 0

d. The impact loads derived from tests with the relatively rigid 9

QSTF vent header are assumed to provide a conservative load definition for the actual Mark I vent headers.

Assumptions a. and b. imply that the impact impulse (integration of impact pressure over the time of impact) is directly proportional to the impact velocity. In the 1/4 scale 2D generic sensitivity test report (Reference 4.3.3-3) the peak force, impulse, and duration of the net vent header force are plotted versus V , F, and 1/V respectively. The trends indicated there provide justification for Assumptions a. and b.

The justification for Assumptic,n c. is that the spacing of the downcomer pairs along the header, while not exactly the same in any two plants, is quite similar in all the plants. Therefore, the results from the 1/12 scale tests modeling a typical Mark I plant can be applied generically.

4.3.3.2 Load Definition

a. Downcome rs 9

In the 1/4 scale tests the measured impact / drag loading on the downcomers was found to be of very low magnitude. Figure 4.3.3-2 provides the generic downcomer load definition. The downcomer pressure transient defined in Figure 4.3.3-2 should be applied unifor ly over the bottom 50' of the angled portion of the down-comer only, as shown in Figure 4.3.3-3. The impact pressure transient begins as soon as the rising pool reaches the lower end of the angled portion of the downcomer as determined by the pool swell displacement / velocity procedure in Section 4.3.4 The transient ends at the time of maximum pool swell.

b. Vent Header The Plant !!nique Load Definitions provide the local impact pres-sure transients for each of the Mark I plants, as well as figures which show the locations of the local impact pressure transients.

Figure /.3.3-4 shows the general form of the vent header local O

4.3.3-4 Revision 0

NED0-21888 7 .,

impact pressure transients. The local impact pressure transients an. given in units of psid relative to the torus airspace pres-sure at the time of vent header impact. The pressure difference between the inside of the vent header and the torus airspace at the time of vent header impact is specified on a plant unique basis in Table 4.3.3-1.

The impact pressure transients in the Plant Unique Load Definitions (PULD) correspond to the average downcomer spacing impact velocity.

Since the impact velocity varies along the length of the header, the local impact pressure transients need to be adjusted for impact velocity. This is accomplished via the figures in the PULD which specify the vent header impact velocity distribution normalized to the average downcomer spacing impact velocity.

For a normalize:. impact velocity equal to 1.0, no adjustment of the local impact pressure transient is necessary. For other impact velocities both p the impact pressure magnitude and the time duration of the transients need d to be adjusted according to the following:

P=V at = 1/V where P = local impact pressure' magnitude, psi V = normalized impact velocity given in the PULD, f t/see at = time duration of the local impact pressure transient.

All points of the transient should be adjusted relative to the start of the transient, sec.

The Plant Unique Load Definitions also contain the plant unique longitudinal time delays, at

, and the plant unique circumferential time delays, at,g.

f .

v 4.3.3-5. Revision 0

NEDO-21888 The impact pressure transient at any point on the vent header begins at a certain time after the pool first impacts the header. This time is the sum of two quantities: the longitudinal time delay, which accounts for the longitudinal (z) distance from the position of first impact to the point of concern; and the circumferential (3) time delay, which accounts for the time it takes for the point of impact to " sweep" around the header in the cir-cumferential direction. Therefore, the time of impact at any point on the vent header is specified by:

t = t g + at + at g where t = time of impact at any particular point on the header, see t

g

=

time of first impact on the header as determined by the pool swell displacement / velocity procedure in Section 4.3.4, see at = ; lant unique longituckinal time delay for point of interest on the header, see acg = plant unique circumferential time delay for point of interest on the header, sec.

c. Main Vent The impact and drag load on the main vents should be calculated using the procedure discussed in Section 4.3.4 (Impact and Drag on Other Structures Above the Pool). Since the main vent is a large structure and is oriented at an angle to the pool surface, it should be subdivided into smaller sections and the impact and drag loads calculated separately on each subdivision.

O 4.3.3-6 Revision 0

NEDO-21388 Table 4.3.3-1 INTERNAL VENT HEADER / TORUS AIRSPACE PRESSURE DIFFERENTIAL AT TIME OF VENT HEADER IMPACT Pressure Differential (psi)

Plant Operating AP Zero aP Brunswick 1,2 11.4

Browns Ferry 1, 2, 3 5.2 5.8 Cooper 8.9 11.3 Dresden 2,3 7.4 8.9 Duane Arnold 6.1 8.6 Fermi 2 6.5 7.1 Fit = patrick 6.9 9.9 Hatch 1 6.9 10.1 Hatch 2 9.4

Hope Creek 1,2 6.3

8.2 (q/ Millstone Monticello 11.2 12.2 13.6

~

Nine Mile Pcint 10.5 12.2 Oyster Creek 8.3 11.2 Peach Bottom 2,3 10.6 11.0 Pilgrim 9.7 11.4 Quad Cities 1,2 7.4 8.9 Vermont Yankee 7.1 12.7

Normal operating condition is zero drywell/wetwell pressure differential.

~

4.3.3-7 Revision 0

_ _ _ _ _ _ . _ _ , . . . - _ _ - _ _ _ . - - - - - - - - - - t

i k

( MIDWAY BETWEEN 4 MAIN VENT Q . TWO MAIN VENTS MAIN VENT &

VENT HE ADEH 8 ME ASUREO IN y

ESTHER DIRECTION %

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3

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O a s

l TIME WHEN POOL TIME OF TIM E (msec)

REACHES LOWER MAXIMUM END OF ANGLED POOL SWE LL PORTION OF DOWNCOMER l

l Figure 4.3.3-2. Downcomer Impact and Drag Pressure Transient 4.3.3-9 Revision 0

I VENT HE ADER A

DowNCOMER (ANGLED SECTION) 's ,

1 i

,z,-'-'-"z

A IMPACT PRESSURE TRANSIENT l ,' ,' APPLIED TO SHADED REGION v'

/

l l i, /

50 DOWNCOMER (VERTICAL SECTION) l l

l Figure 4.3.3-3. Application of Impact / Drag Pressure Transient to Downcomer 4.3.3-10 Revision 0

.. . - . - . . . . - . . . _ - - - . - . . ~ . - - - - . . - - . . . . . - _ . _ . - . - . - - - . - . - . - - - - - - - - , .

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k l P3 4

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l Vent Header Local 1:npact Pressure Transient Figure 4.3.3-4.

1 4.3.3-11/4.3.3-12 Revision 0 1

U REFERENCES FOR SECTION 4.3.3 4.3.3-1 J. M. Humphrey, " Mark I Containment Program 1/4 Scale Pressure Suppression Pool Swell Test Program: LDR Load Tests - Generic Sensitivity," General Electric Company, Report No. 23545-P, December 1978.

4.3.3-2 The Electric Power Research Institute, "Three-Dimensional Pool Swell Modeling of a Mark I Suppression System," EPRI NP-906, October 1978.

A V

l l

r O :

4.3.3-13/4.3.3-14 Revision 0 i

4.3.4 Impact and Drag on Other Structures Above the Pool During the pool swell transsent, the rising pool will impact structures above the initial pool surface and below the maximum pool swell height. The loading on structures above the pool, other than the vent system (which includes the vent header, vent header deflector, and downcomers), is specified in this section.

As the pool surface rises during pool swell, it impacts structures located within its range of travel. Loads are generated due to both the impact forces and drag forces. The timing and amplitude of the loading on a particular structure depends on the velocity of the pool surface as it impacts and flowt past the structure. To facilitate the calculation of these impact and drag loads, the first part of this section provides a method for calculating the pool swell impact velocity at any point in the torus. Then, the two procedures for calculating the impact and drag loads are given.

4.3.4.1 Bases and Assumptions V

The primary input to the calculatio'n of the pool swell impact and drag loads is the pool swell impact velocity. The pool swell impact velocity calculation is based on analysis of high speed motion picture data obtained during tests performed at the 1/4 Scale 2D Test Facility (QSTF) and at the 1/12 Scale 3D Test Facility described in References 4.3.4-1 and 4.3.4-2, respectively.

Plant unique QSTF pool swell tests were performed to obtain 2-D pool swell transients corresponding to the average downcomer spacing in the torus.

For each plant the torus and vent system geometry and the plant operating conditions were modeled. Downcomer spacing in the test model was scaled to be equivalent to the average downcomer spacing in the plant being tested.

Motion pictures of the 1/12 scale 3-D pool swell tests were used to obtain the displacement and velocity profiles along the longitudinal axis (z) of the torus. A single representative geometry was used in conjunction with selected pressurization rates to represent generic plant belutvior, bi V

4.3.4-1 Revision O l

The assumptions made when defining the pool swell velocity and displacement behavior are as follows:

a. The pool swell phenomenon is sy= metric from sector to sector (a sector is defined in Figure 4.3.4-1).

b. The normalized pool swell 2-D displacement and velocity profiles obtained from the QSTF tests are representative of the displacement and velocity profiles at every longitudinal position in the Corus.

c. The longitudinal distribution of displacement and velocity obtained from the 1/12 scale 3-D tests is applicable to all Mark I plants.

Reference 4.3.4-2 discusses the 1/12 scale 3-D pool swell tests. These tests indicated that pool swell is symmetric f rom sector to sector, justifying Assumption a.

O The pool swell chape obtained in the QSTF tests (Reference 4.3.4-1) applies to the average cell spacing. Assumption b. is justified because even in regions with different downcomer spacings, the shape and velocity profiles will not be substantially different.

Assumption c. is based on the fact that the downcomer spacing geometry for the typical plant modeled in the 1/12 scale 3-D tests is similar to that in each of the Mark I plants.

The assumptions employed in the load definition for pool swell impact loads are as follows:

a. The impact load is calculated assuming the impacted structure is rig id ,

b. The impact pressure is parabolic in time. The duration of impact is calculated based on assuming that impact occurs over the bottom 50' included angle of th: impacted structure.

4.3.4-2 Revision 0

c. The hydrodynamic mass factor, g , is assumed to be 0.2 for all struc" ares of concern.

d. Based on the results reported in Reference 4.3.4-3, the impact load on gratings is defined as negligible (however, a drag load is specified).

The value of the hydrodynamic mass factor, Kh, was determined by comparing the measured impact impulses on the QSTF vent header during the Part 1 test series (Reference 4.3.4-4) to the impact calculated using the method explained in this section. K was determined by the comparison. In all cases K was calculated h h to be less than 0.20; K hwas conservatively assumed to be 0.20.

The assumptions employed in the load definition for pool swell drag are as follows:

a. Standard drag coefficients are applicable.

b. The maximum velocity at the (x,z) position of the structure is used

, for the standard drag calculation, instead of the actual pool swell velocity at structure location (x,y,z); it is assumed that such a drag load bounds the sum of the actual standard drag and acceleration drag at the same location at all times.

4.3.4.2 Determination of Pool Swell Displacement and Velocity.

To calculate the pool swell im;act and drag loads on other structures above the pool, the pool swell displacement and velocity transients must be known at the (x,z) location of the structure of interest (Figure 4.3.4-1 shows the coordi-nate system used for this analysis). This subsection describes a procedure for calculating the pool swell displacement and velocity transient at any point in the torus.

C\

- V 4.3.4-3 Revision 0 m

Figure 4.3.4-2 is a typical plot of the pool surface displacement longitudinal dis tribution. This describes the pool swell displacement at any z position normali:ed to the displacement corresponding to the average downcomer spacing.

In other words, Figure 4.3.4-2 is a plot of 6y,(2,t) ay,(t) where ay,(z,t) = the pool displacement at time t above the initial pool surface ay,(t) = the same parameter corresponding to the average downconer spacing.

Figure 4.3.4-3 is a plot of a typical plant pool surface displacement in the xy plane, ay,(x,t), for the average downcomer spacing. The pool swell dis-placement at any (x,z) location in the torus and at any time t is given by the product ay,(2, c) _

x ay*(x,t),

ay,(t)

A procedure similar to that indicated above is used to find the pool swell velocity transient at any point in the torus. Figure 4.3.4-4 shows the typical longitudinal velocity distribution and Figure 4.3.4-5 shows z. typical plant velocity transient in the (x,y) plane for the average downcomer spacing.

The pool swell velocity at any (x,:) location in the torus and at any time t is given by the product x v(x,t) v(t)

O 4.3.4-4 Revision 0

NEDO-21888 m

(JI where is from Figure 4.3.4-4 v(c) v(x,t) is from Figure 4.3.4-5 for a typical plant.

By using Figures 4.3.4-2 and 4.3.4-4 along with the appropriate plant unique figure corresponding to Figures 4.3.4-3 and 4.3.4-5, the pool swell displace-ment and velocity transients can be constructed for any (x,z) location in a Mark I plant.

4.3.4.3 Load Definition for Impact The procedure for calculating the pool swell impact loads on structures (other than the vent system) above the pool also uses the pool swell impact velocity as a primary input.

U During impact each structure must stop or redirect the momentum of the water which impacts it. An upper bound value of this momentum is given by the product of the hydrodynamic mass of the structure and the impact velocity.

The momentum arrested during impact is also equal to the impact impulse.

In fact, the impulse is less than the product of hydrodynamic & ass and impact velocity because much of the impacting momentum is redirected r.ther than totally arrested. In equation form the above becomes I=bbY 8e n

where I = impact impulse, Ibf-sec g = hydrodynamic mass of impacted structure. lbo (Table 4.3.f.-1)

V = impact velocity normal to the structure surface being p impacted, ft/sec.

v 4.3.4-5 Revision 0

NEDO-21888 9

g = gravitational constant = 32.4

,, lbm-ft 2

lbf-sec K

proportionality factor = 0.2.

h The impact impulse can be expressed in terms of force per unit projected 4.rea of the structure by dividing by the projected area of the structure. In addi-tion, the impact impulse is given as a ratio of impulse per impact velocity by dividing by the impact velocity. The result is I bb I*Ag n

144 c

where I = impact impulse per unit area, psi-sec A = projected area of the structure, ft .

The duration of impact is calculated assuming impact occurs ovo.r the bottom 50' included angle of the structure. If the structure's cross section is not circular, a circumscribing circle is assumed for this calculation. The impact duration is assumed to be equal to the time necessary for the impact velocity to travel through the 50' included angle, i.e.,

at = d/v n where d = distance corresponding to 50' included angle, f t (see Figure 4.3.4-6) at = impact duration, sec.

The impact load transient is assumed to be parabolic in time. Together with the definition of the impulse and duration, this completes the definition of the impact load on the e~ructure.

O 4.3.4-6 Revision 0

NEDO-21888 4.3.4.4 Load Definition For Drag The pool swell drag load on other structures, which follows the impact load and continues until the time of maximum pool swell, is calculated using the standard drag equation and the maximum pool swell velocity at the (x,z) location on thc structure of interest. The maximum pool swell velocity is found using the procedure outlined for determining the pool swell displace-ment / velocity characteristics (Section 4.3.4-2). The drag load is then given by F = pAV D n where F = drag load on structure, lbf D

p = density of pool water = 62.4 lbm/ft A = projected area of structure, ft V = impact velocity normal to the impacted surface, f t/sec CD = standard drag coefficient for structure of interest g = gravitational constant = 32.2 lbe-ft e

Ibf-sec 4.3.4.5 Application of Impact and Drag Loads on Structure The impact and drag pressure transient should be distributed uniformly over the surface. The load should be applied in the upward direction most critical to

-the structural evaluation of the structure within the specified load range.

The range of directions which should be considered includ' a variation of :10*

from the upward vertical.

./O L]

4.3.4-7/4.3.4-8 Revision 0

^s' (G.

Table 4.3.4-1 HYDRODYNr!IC l' ASS .CD ACCELEMTION DRAG '70LU:ES FOR WO-DI:1ENSIONAI STRUCTUML CO3OONENTS (LENGTH L FCR ALL STRUC-~3ES)

SECTICN THROUGH SCOY AND UNIFORM HYORODYNAMIC MASS ACCE! RATION ORAG SODY FLOW OIRECTION (PATTON.1965) VOLUME V 4 ,

)

R awA2L 2rR 2 L CIRCLE : ;

l ELLIPSE : z- s wa2L wa fa +bl L 1

i I

l b

ELLIPSE : I gro2 1 ab(a+bil.

g a me

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$ 1.21 a ra2L att4b+1.21 ral

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- 1/10 2.23 awa2L aL(4b+2 23ral e2 N 2 0 85 ora 4L all2b4 85was I A Of AMONO : L- 2a 1 0 76 awa2L a L(2b+0. 76 ra s I

1/2 0 67 ,,.2L aL(2b4 67ea) 14 5 0 61 awa2L aL(2b 0 61tal

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4.3,;-9 Revision 0

Table 4.3.4-1 (Continued)

HYDRODY::A: IC !O.SS AND ACCELERATIO!: DR.A.G VOLDES FOR r.?C-DDE:: SIC::AL STRUCTURAL CC:90: E::T5 (LD:GTH L FCR ALL STRUCTURES)

SECTICN THROUGH B00V AND UNIFORM HYORODYN AMIC MASS ACCELERATION ORAG BCDY F LOW OIR ECTION (PATTON.1965) VOLUME V 4 y a/c-2.6. b/c-3 6 1-8 E AM M C 2a

' # '* ** * #~

al P

2b M T- -

t y a.c-2 6t: c-26 16EAM M C 2a

'" # ## L i' ** ' ** ~## i L ek 2b HYORODYNAMIC MASS AND ACCELERATION ORAG VOLUMES FOR THREE OIMENSIONAL STRUCTURES BODY AND FLOW HYORODYN AMIC MASS ACCELER ATION ORAG OESCRIPTION DIRECTION (PATTON,19651 VOLUME V, f

CIRCULAR 8/3R3 8/3R3 R

bla e or/6ba2 re6ba2 E LLIPTICA L b OISK - 3 03 ar,6ba2 0 9=i6ba2 g 2 0 826 ar/6ba2 0 826r,6b,2

1.5 0 748 ari6ba2 0. 748 e;6ba2 10 0 637 ar6/ba2 0 637e/6ba2 4.3.5-10 Revision 0 l-

NEDO-21888 i%

U Table 4.3.4-1 (Continued)

HYDRODYNAMIC MASS AND ACCELERATION DRAG VOLUMES FOR TWO-DIMENSIONAL STRUCTURAL COFTONENT3 (LENGTH L FOR ALL STRUCTLRES)

BODY AND FLOW HYDRODYNAMIC MASS ACCELERATION DRAG DESCRIPTION DIR ECTION (PATTON,1965) VOLUME V A y b/s RECTANGULAR gllllD N ~

PLATE a 1 0.4 78 av/4a2h 0.478w/4a2b b

7 ._ g,, 1.5 0.680 aw/4a2b 0.680s/4a2b

2 0.840 aw/4a2b 0.840s/4a2d

/

2.5 0.953 aw/4a26 0.953 w/4a2b y 3 pr/4a2b w/4a2b pw/4a2b w/4a2b TRIANGULAR PLATE e a

,, 3 (TAN 81 3/2 , 3 (TAN 8)3/2

/ & y a w w y

s. y e SPHERE

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Figure 4.3.4-2. Typical Pool Surface Displacement Longitudinal Distribution 1

4.3.4-13 Revision 0 l

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NEDO-21888 l 8.0 '

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Figure 4.3.4-3. Typical Plant Pool Surface Displacement in 9:

XY Plane 4.3.4-14 Revision O

l l

NEDO-21888 lO t/tMAX

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Figure 4.3.4-4. Typical Pool Surface Velocity Longitudinal Distribution 4

4.3.4-15 Revision 0

NEDO-21888 9

60 s0 -

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5

=

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I ! l l I 0 0.2 0.4 0.6 0,8 1.0 1.2 HORIZONTAL OBSTANCE (x/R)

Figure 4.3.4-5. Typical Plant Velocity Transient in XY Plane 4.3.4-16 Revision 0 t

- - . - -...._-....n-.,---.,,,,,.a- , , , . , , , , , .--,.._,,,_---,.-.-,,a, ,,--,.n-,-,m.,,w,.,

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STRUCTURE OF CONCERN OR CIRCUMSCRIBED CYLINDER e

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REFERENCES FOR SECTION 4.3.4 4.3.4-1 J. M. Humphrey, " Mark I Containment Program 1/4 Scale Two-Dimensional Plant Unique Pool Swell Test Report," General Electric Company, Report No. NEDE-21944-P, to be published.

4.3.4-2 The Electric Power Research Institute, "Three-Dimensional Pool Swell Modeling of a Mark I Suppression System," EPRI NP-906, October 1978.

4.3.4-3 T. R. McIntyre, et al., " Mark III Confirmatory Test Program One-Third Scale Pool Swell. Impact Tests - Test Series 5805," General Electric Company, Report No. NEDE-13426-P, August 1975.

4.3.4-4 D. L. Galyardt, et al., " Mark I Containment Program 1/4 Scale Pressure Suppression Pool Swell Test Program: Scaling Evaluation,"

General Electric Company, Peport No. NEDE-21627-P, January 1978.

i 4.3.4-19/4.3.4-20 Revision 0

NCDO-21888 h.-

s 4.3.5 Froth Impingement Loads During the LOCA pool swell transient one of the loading conditions in the torus is due to froth impingement. This section describes these loads.

Froth is an air-water mixture which rises above the pool surface and may impinge on the torus walls and structures within *.he torus airspace. ~After-ward the froth will fall back, creating froth fallback loads. There are two mechanisms by which froth may be generated:

a. As the rising pool strikes the bottom of the vent header and/or the vent header protection device, froth is formed which travels upward and to both sides of the vent header.

b. A portion of the water above the expanding air bubble becomes detached from the bulk pool; this water is influenced only by its

,, - - own inertia and gravity. The water breaks into froth which rises E

s-up into the air space beyond the maximum bulk pool swell height.

4.3.5.1 Bases and Assumptions The basis for the froth impingement load definition is motion-picture data obtained from experimental pool swell tests performed on the 1/4 scale two-dimensional test facility (QSTF) (Reference 4.3.5-1). Although the QSTF tests are not scaled after bubble breakthrough, they do give an approximation of the froth formation phenomenon.

The following assumptions were made when defining the froth-impingement loads:

a. No froth velocity is assumed in the torus longitudinal direction.

b. The first froth formation mechanism described in Section 4.3.5 results in negligible loads on the torus shell.

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c. For the first froth formation mechanism described in Section 4.3.5 the froth density is assumed to be 10% of full water density.

.For the second froth formation mechanism described in Section 4.3.5, the froth density is assumed to be 100% of full water density for structures or sections of structures smaller than 1 ft in all linear dimensions, and 25% water density for structures larger than 1 ft. The exception to this are structures located above the vent header, for which the froth density is assumed to be 10% of full water density.

Ten and twenty-five percent of full water density is also used for the froth fallback density for structures above and to the side of the vent header, respectively.

Assumption a. is based on analysis of the 1/12 scale 3D pool swell test movies (Reference 4.3.5-2), which showed little froth velocity in the torus longitudinal direc tion.

Assumption b. concerns the froth which sprays off the vent header as the pool impacts that structure. This phenomenon was noted in the QSTF movies (Refer-ence 4.3.5-1). The froth spray which forms is of relatively low density; if this froth spray reaches the torus wall, it will have diffused to such an extent so as to produce a negligible load on the torus shell.

Assumption c. addresses the froth density. As discussed above, the froth density of the first froth formation mechanism appears relatively small in the QSTF movies. For small structures Assumption c. conservatively assumes 100% water density for the second froth formation mechanism, and a more realistic 25%

water density for larger structures. Structures above the vent header are shielded from froth (this was indicated in the QSTF movies, Reference 4.3.5-1) and therefore a 10% water density for the froth is assumed in this region.

4.3.5.2 I. cad Definition Procedure Observation of the QSTF motion-pictures indicated two predominant mechanisms for 'roth generation and also identified the regions where thase loads apply.

4.3.5-2 Revision 0

_O Two regions of the torus are identified such that within each region the loading is the result of one of the two mutually exclusive froth formation mechanisms.

Region I, as indicated in Figure 4.3.5-1, is the region where loading results from the pool surface impacting the vent header and/or deflector. Region II, as indicated in Figure 4.3.5-2, is the region where loading results from froth formed when the water above the rising bubble separates from the bulk pool and rises into the airspace.

For those structures in Region I, the froth impingement pressure is calculated using the following equation:

pg V f"g x 144 where Pg = froth impingement pressure, psi og = froth density. Equ'ls a 10% water density, lbm/ft V = Pool surface velocity at elevation of lower surface of vent header, ft/sec lba - ft g = 32.2 e

ibf - see Figure 4.3.5-1 shows a typical structure and the specified direction of appli-cation of the froth impingement loading within Region I. Figure 4.3.5-3 shows the time history assumed for the froth load.

1 For those structures in Region II the froth impingement velocity, Vg , is calculated assuming the froth travels vertically to the structure of concern under the influence of gravity only, from the height of the maximum pool swell underneath the structure at the velocity of the pool surface at the time of I

u)

{

4.3.5-3 Revision 0 l

l i

NEDO- 21888 O

maximum pool swell. This velocity is determined by the procedure described in Section 4.3.4.2. The froth impingement pressure is then calculated by l

n og V' g

f"g x 144 where Vf = froth dmpingement velocity, ft/see p g = froth density, determined by the criteria described in 3

Section 4.3.5-1, 1bm/ft The characteristics of the loading time history are shown in Figure 4.3.5-4 at gis the time needed for the froth to rise from the maximum pool swell height to the structure. Figure 4.3.5-5 specifies the range of possible directions of load application for structures in Region II. The same range of directions of application shown in Figure 4.3.5-5 for the torus (x,y) plane (245' from vertical) should be considered for the torus (y,z) plane. The g

calculated pressure should be applied to the structure of concern as a rectangular pulse 100 milliseconds in length in the direction most critical to the capability of the structure.

The froth fallback pressure is calculated by 2

p ,

A ff Yff ff 144 g where pfg = froth fallback pressure, psi Vgg = froth fallback velocity, f t/see p

gf = froth fallback density = 25 : of full water density except for the region directly above the vent header, where it equals 10*. of full water density, 1bm/ft3 4.3.5-4 Revision 0

NEDO-21888 The froth fallback velocity is calculated by allowing the froth to freefall from the maximum froth height (to which the froth rises under the influence of gravity alone) or the height of the upper torus shell, whichever is lower.

The froth fallback pressure is applied uniformly to the upper projected area of the structure of concern in the direction most critical to the behavior of the structure. The froth fallback is specified to start when the froth impingement load in Region II ends, and lasts for 1.0 seconds. Figure 4.3.5-6 indicates the directions of application which should be considered. The same range of directions of application shown in Figure 4.3.5-6 for the torus (x,y) plane (245* from vertical) should be considered for the torus (y,z) plane.

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[ 4.3.5-5/4.3.5-6 Revision O I

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FROTH REGION I (TYP OF BOTH VENT SIDES OF HEADER)

HEADER

-- TYPICAL STRUCTURE I

45 DIRECTION OF LOAD APPLICATION O !

TORUS l

l Figure 4.3.5-1. Definition of Froth Impingement Region I 4.3.5-7 Revision 0

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REGION tl 4 0.6R E MAXIMUM POOL SWELL PROFILE

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U TORUS Figure 4.3.5-2. Definition of Froth Impingement Region II l 4.3.5-8 Revision 0

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TIM E (msec)

TISAE OF VENT HE ADER IMPACT Figure 4.3.5-3. Froth Ioading History - Region I 4.3.5-9 Revision 0 l

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3

$ Pg -

a

$ O 100 msec 4 y y

TIME (msec; TfME OF MAXIMUM POOL SWE LL

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l Figure 4.3.5-4 Froth Loading History - Region II i

4.3.5-10 Revision 0 l

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STRUCTURE POS$18LE DIRECTIONS OF LOAD l

VENT HEADER fh TORUS 4

4 4

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Possible Directions of Load Application t 4.3.5-11 Revision 0 i .

,-,.. . -. , - __ ,_ .._ ... ._, _. ,-_.-._-__ ~ ____. ... _._, _ . . _ _ _--_. .-___._._----__ - i

POSSIBLE DIRECTIONS OF LOAD TORUS VENT $

HEADER STRUCTURE Figure 4.3.5-6. Possible Directions of Froth Fallback Load Application 4.3.5-12 Revision 0

,. . -. . , . _ . . . - - . . , . . . . . . . . ~ . . . . . . . . . - =. . . . . . . _ . . -

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! NEDO-21888 REFERENCES'FOR SECTION 4.3.5

~ 4.3.5-1 J. M.' Humphrey, " Mark I Containment Program 1/4 Scale Two-Dimensional i Plant Unique Pool Swell Test Report," General Electric Company, I' Report No. NEDE-21944-P, to be published.'

k i

4.3.5-2 The Electric-Power Research Institute, "Three-Dimensional Pool Swell i

Modeling of a Mark I Suppression System," EPRI NP-906, October 1978.

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NEDO-21888 q

O 4.3.6 Pool Fallback Loads Following the pool swell transient, the pool water falls back to its original level and in the process generates fallback loads. After the pool surface has reached its maximum height due to pool swell, it falls back under the influence of gravity, creating drag loads on structures inside the torus shell which are between the maximum bulk pool swell height and the downcomer exit level. The fall-back load starts as soon as the pool reaches its maximum height and ends when the pool surface falls past the structure of concern. The fallback load on the torus shell is included in the definition of the torus vertical load due to pool swell (Sections 4.3.1 and 4.3.2).

4.3.6.1 Bases and Assumptions The basis for the pool fallback load definition is the pool swell data obtained from analysis of movies from the QSTF tests described in Reference 4.3.6-1 and from the 1/12 scale 3D tests described in Reference 4.3.6-2.

The procedure for determining fallback loads utilizes the following major assumptions:

a. The fallback load on structures below the downcomer exit level is negligible.

b. Fallback loads on the vent header, downcomers, and main vents are considered negligible.

c. Standard drag coefficients apply.

d. The drag force is the sum of standard drag (proportional to velocity l squared) and acceleration drag (proportional to acceleration). ,

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4.3.6.2 Load Definition Procedure Fallback loads are applied to the top surfaces of the structures affected.

If the top surface area of the structure of concern is large, the area should be subdivided and the fallback loads calculated individually for ecch smaller ar ea . As a guideline, if any linear dimension of a surface is greater than 1 foot, the area should be divided into smaller areas.

The procedure requires determinining the maximum pool swell height above the top surface of the structure. The pool swell displacement / velocity procedure described in Section 4.3.4 is used for this. Freefall of the bulk fluid from this height produces both standard drag and acceleration drag, with ine total drag given by the sum. The fallback load is applied uniformly over the upper projected surface of the structure in the direction most critical to the behavior of the structure. Figure 4.3.6-1 indicates the possible directions of application to be considered. This range (245' from the vertical) applies to the torus (y,z) plane as well as the torus (x,y) plane shown in Figure 4.3.6-1.

O.I 4.3.6-2 Revision 0

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POSSIB LE I VENT DIRECTIONS HEADER OF LOAD f \ / .

O TORUS STRUCTURE l

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NEDO-21888 1

REFERENCES FOR SECTION 4.3.6 4.3.6-1 J. M. Humphrey, " Mark I Contain=ent Program 1/4 Scale 2-D Plant Unique Pool Swell Test Report," General Electric Company, Report No.

NEDE-21944-P, to be published.

1 4.3.6-2 The Electric Power Research Institute "Three-Dimensional Pool Swell Modeling of a Mark I Suppression System, EPRI NP906, October 1978.

4.3.6-3 F. J. Moody, " Analytical Model for Estimating Drag Forces on Rigid Submerged Structures Caused by LOCA and Safety Relief Valve Ramshead Air Discharges," General Electric Company, Report No. NEDO-21471, 4 September 1977.

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4.3.6-5/4.3.6-6 Revision 0 l

NEDO-21888 4.3.7 LOCA Jet Load As the drywell pressurizes during a postulated LOCA, the water slug (if any) initially standing in the submerged portion of each downcomer is accelerated downward into the suppression pool. As the water slug enters the pool, it forms a jet which could potentially load structures which are intercepted by the discharge. The calculation of these Joads is dascribed in this section.

Jet loading affects structures which are enclosed by the jet boundaries and lasts from the time that the jet first reaches the structure until the time when the last particle of the water slug passes the structure.

4.3.7.1 Load Definition The calculation procedure described to obtain LOCA jet loads is based on both experimental data obtained from tests performed at the 2: Scale 2D Test Facility (QSTF) and also on an analytical model described in Reference 4.3.7-1.

O Plant unique downcomer clear / g 1 !?rmation was obtained experimentally during 4

the QSTF testing in the forn of LOCA jet fluid velacity and acceleration histories.

The major assumptions incli,3ed in the methodology are as follows:

1. When a structure is engulfed in the jet path, the force on the structure can be calculated using a standard velocity-squared drag equation. If a structure partially or fully intercepts the jet, a momentum balance is used to define a proportionality factor to be ussd in place of the standard drag coefficient.

2. The jet is completely dissipated when the last fluid particle leaving the downcomer catches up to the front end of the jot; I the dissipated jet provides no more drag.

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NED0-21888 4.3.7.2 Evaluation Procedure The procedure used to calculate forces on submerged structures in the LOCA jet path is presented below.

First, the following plant unique LOCA jet information is obtained:

r = an identifier for each fluid particle in the jet, equal to the time at which the fluid particle was discharged from the downcomer.

VD (t) = jet discharge velocity for particle r, ft/sec dV D

d:

"D(T) jet discharge acceleration for particle , f t/sec 2 Next, the c:aximum jet penetration,1, is determined by i . VD 2 (rmax'), ge O p a D (Imax}

where max denotes the last ti=e step, i.e., vent clearing time.

If I < L , there is no jet load on the structure.

p S If 1 > L,, there is a jet load on the structure.

L, = vertical (X) distance (f t) from the bottom of the downcomer to che structure of concern. Note that the X coordinate defined here is different from that defined previously for the torus.

The plant unique LOCA jet infor=ation is used to construct a plot of the jet front position vs. time. This is done by plotting the position (denoced as Xj)

O 4.3.7-2 Revision 0 i

/' '

l d vs time (elapsed time following start of DBA) of a number of fluid particles which exit the downcomer during the clearing transient. This time-space plot (t,x) is based on the following equation:

l x -

V3 (t) (t-t) 3 Whenever a line of constant t overtakes an earlier line, the earlier line ceases to extend and the overtaking line continues. The time-space plot enables one to enter a location x, and time t, read the corresponding value of f , and then obtain the associated jet velocity from the plant unique LOCA jet data.

As the jet travels through the pool, the particles at the rear, which were discharged from the downcomer at higher velocities, catch up with particles at the front, which were discharged at lower velocities. When this "over-taking" occurs both particles continue on at the higher velocity. When the last fluid particle leaving the downcomer catches up to the front of the jet, the jet dissipates, which shows that as the rear particles cctch up to the particles in front, the jet becomes shorter and wider.

For a structure at position x, values of t and t can be obtained from the time-space plot explained above. Then the above equation for Xj and the plant unique LOCA jet data can be employed to determine jet area at position x for various values of time, t. The jet area is given by:

A ADV D('}

D 1- I 1 ) (dVDC') b ,I 1 \ fdYD ('}\

dr 1 dt *j VD (t) j ) VD (I)/

where A is the discharge area.

D The area of the jet front is approximated by the area the jet possesses one initial discharge diameter D = (4AD /n) ! behind the actual jet front.

ol w,

4.3.7-3 Revision 0

HEDO-21888 If L, < D, the jet front area should be approximated as x = L,. The jet radius is calculated in order to determine which structures overlap the jet, and how much overlap actually occurs.

The standard drag force on the submerged scructure is computed based on the normal component of velocity intercepting the structure, ethe projected area of the structure intercepted by the normal component of velocity, and the jet area.

If a structure is fully submerged inside the jet boundary at time t, a velocity-squared drag calculation is performed:

CgpA D(V (t) cos G)2 p F (t) =

2g

, lb f

c where C = applicable standard drag coefficient D

A p = the projected area of the structure intercepted by the normal component of velocity, ft' h

0 = angle between the normal to surface A and the vercical plane p = density of pool water = 62.4 ft lbm-ft g" = 32.2 lbf-sec V (t)= discharge velocity at T for the structure position D L,, ft/sec.

Note that since F (t) is evaluated at a given X (that of the structure), the variables t and T are not independent of each other, but are related by the time-space plot. For this reason V (t) is equal to the jet velocity at the D

structure at time t, and therefore, use of V D(T) in the above equation results in F (t).

O 4.3.7-4 Revision 0 l

1 l

l

NEDO-21888 .

If a structure fully or partially intercepts the jet (no jet flow around

. structure) the following equation applies (VD(T) COS 0) o F (t) = K A j P 2g

, 1b f

d e

= 4 if the structure q where K) = 2 for momentum stoppage with no momentum or K turns the jet back on itself. A represents the projected area intercepted by

. P the jet.

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O I 4.3.7-5/4.3.7-6 Revision 0 eT.- -

--e-t- <-,-,-,.y-= y e , ,,y. 4,-.yy-,,,--y.#,3-,---, v. ew- -g -v,-g,.,,wwwren,- -

,-,,,,,.-3%,,-wy,-y-w, ,,,w..., w, -

--y-- ,-w,wmwy - - - - -

NEDO-21888 O REFERENCES FOR SECTION 4.3.7 i

4.3.7-1 F. J. Moody,"Analytica1Model for Liquid Jet Properties for Predicting Forces in Rigid Submerged Structures," General Electric Company, Report No. NEDE-21472, September 1977.

i

(

O i

l lO l 4.3.7-7/4.3.7-8 Revision 0 i i

(,,/ 4.3.8 LOCA Bubble - Induced Drag Loads on Submerged Structures During the initial phase of the DBA, pressurized drywell air is purged into the suppression pool through the submerged downcomers. After vent clearing, a single bubble is formed around each downcomer. For the DBA the duration of the LOCA bubble is typically 0.2 seconds from its initial formation until it breaks through the pool surface. It is during the bubble growth period that unsteady fluid motion is created within the suppression pool. During this period all submerged structures below the pool surface will be exposed to transient hydrodynamic loads.

4.3.8.1 Bases and Assumptions The bases of the flow model and load evaluation for the LOCA bubble-induced submerged structure load definition are derived from the model in Reference 4.8.3-1 and are summarized briefly below.

The major assumptions used to develop the analytical models are listed below.

O V

a. Bubble Dynamics - The initial LOCA bubble pressure is the same as the drywell pressure at vent clearing.

This is established from the QSTF test data (maximum DBA PDW}*

- Air is an ideal gas.

- The bubble is spherical and characterized by the Rayleigh equation.

-b. Flow Fields - The flow fields are established by using inviscid, irrocatio'nal potential flow theory with the bubble considered as a fluid source.

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4.3.8-1 Revision O i

NED0-21888

- The presence of boundaries is considered by using the method of images (sink for the free surface and source for the solid walls). The circular cross section of the torus shell is modeled b/ rectangular cross sections. The basis for this is contained in the model evalua-tion report (Ref. 4.3.8-2),

c. Drag Loads - Submerged structures are assumed rigid to maxi-n Stmetures mi e the load.

- Sum of the standard drag and acceleration drag is the total drag force on a structure. The basis for these two components of the drag force was established in References 4.3.8-1 and 4.3.8-2.

A series of tests on submerged structures was performed in the Mark I QSTF (Ref-erence 4.3.8-3). The tests measured LOCA bubble-induced loads on four cylindrical structures submerged both horizontally and vertically near the downcomer exists.

The entire blowdown event for each test was also filmed by a high speed movie camera. The main objective of the test series was to acquire quantitative total drag loading data on cylindrical structures exposed to an accelerating flow field in aa actual scaled prototypical configuration and to compare these results with the pred.ctions of the analytical model given in Ref-erence 4.3.8-1. Detailed *.est procedures and measured results are presented in Reference 4.3.8-3. The model evaluation report (Reference 4.3.8-2) com-pares the test data to the results predicted by the analytical model.

4.3.8.2 Load Definition Detailed procedures to cal. alate drag forces on submerged structures are described in Reference 4.3.8-1 which provides the methodology of utilizing the unsteady velocity and acceleration flow fields in the suppression pool to obtain drag loads on s bmerged structurei, such as pipes, beams, etc.

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NEDO-21888 v

The load dufinition method first requires establishement of the velocity and acceleration flow fields within the torus by simulating the DBA con-dicion at the vent exists. Then, drag loads on submerged structures are calculated based on these transient flow fields.

a. Flow Field Establishment An air charged bubble is considered as the fluid source having a source strength determined by the dynamic characteristics of a bubble. The Mark I torus LOCA bubbles are simulated by multiple sources in a finite pool. By using potential flow theory and the method of images for source and sink considerations to simulate solid walls and a free surface, the velocity and acceleration flow fields within the torus are established.

b. Drag Loads Evaluation

. The drag force on a submerged structure consists of two components:

the standard drag and the acceleration drag. The standard drag

() load is computed on a locally uniform flow field evaluated at the local velocity. The acceleration drag load is evaluated based on an inviscid, uniform but i . steady (accelerating) flow field. The sum of these two drag fore gives the total drag load on a sub-merged structure.

4.3.8.3 Selection of Key Parameters For Load Evaluation Procedures This section outlines the key parameters which influence the magnitude and s

characteristics of the LOCA bubble drag loads on submerged structures. The procedure used to specify the input for the submerged structure drag load model is also discussed.

The plant specific suppression pool geometry should be identified first as follows:

a. Torus shell dimensions

b. Torus water depth

(

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NED0-21888

c. Location of submerged structure considered (Figure 4.3.S-1)

d. Locations of douncomers (Figure 4.3.8-1).

Other plant specific data which need to be identified are:

a. DBA transient drywell pressure

b. Thermodynamic properties of drywell air and wetwell water

c. Initial dr"well/wetwell pressure differential

d. Overall vent system friction factor

e. Downcomer inside diameter

f. Downcomer submergence

g. Initial bubble surface velocity

h. Undisturbed pressure at bubble position before bubble appears.

The submerged structure considered is divided into the appropriate number of sections. The coordinates of each section center, orientation of the struc ,

ture axes and the structural projected cross-sectional area are identified.

The acceleration drag volume and the standard drag coefficient for each section of the submerged structure are then obtained from tables given in Reference 4.3.8-1.

With the input parameters identified and calculated, the submerged struc-ture drag load model is used to evaluate the resultant transient drag forces on each structure element being analyzed. In addition, transient bubble dynamic characteristics within the pool can be found. Tveical results are shown in Figures 4.3.8-2 and 4.3.8-3. These figures show the 7.0CA bubble drag forces on a vent header sypport column in the radial (X) and tangential (Z) directions as a function of time.

4.3.3-4 Revision 0

)gY VENT HEADER SUPPORT COLUMN a = INITIAL SUBMERGENCE 4- b ->- / b c =

= DISTANCE BETWEEN DOWNCOMER AND TORUS q WATER DEPTH d = DOWNCOMERID me =w

. >-x

< l i e >

' 1

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C 4

5 'I I

SECTION A-A A-A TORUS 9

Y.

MAIN VENT %

? *O' VENT HEADER I 4 Z

m k, w [,< /

) g

) DOWNCOMMER

9. 4 -

( ) ( ,.

D b VENT HEADER SUPPORT COLUMN if X

SECTION A-A x Figure 4.3.8-1. Torus and Typical Submerged Structure Geometry 4.3.8-5 Revision 0 i

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NEDD-21888 O

-600

-500 -

1

-400 -

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o -300 -

+

X

.J s- 9 200 -

-100 -

0

' ! ! I O 0.01 0 02 0.03 0.04 0.05 TIME (sec)

Figure 4.3.8-2. Sample His:.ary of Total X-Forces on Vent Header Support Column 4.3.8-6 Revision 0

-250 l

-200 -

j ~150 -

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u.

O a O -100 -

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l I I I I O

O 0.01 0.02 0.03 0 04 0.05 TIME (m) l l

l Figure 4.3.8-3. Sample Time History of Total Z-Forces on Vent Header Support Column l 4.3.8-7/4.3.3-8 Revision O l

O U

REFERENCES FOR SECTION 4.3.8 J

4.3.8-1 F. J. Moody, " Analytical Model For Estimating Drag Forces on Rigid Submerged Structures Caused by LOCA and Safety Relief Valve Ramshead Air Discharges", General Electric Company, Report No. NEDO-21471, September 1977.

. 4.3.8-2 " Submerged Structures Model Evaluation Report", General Electric Company, Report No. NEDE-21983-P, October 1978.

i 4.3.8-3 "l/4-Scale Test Report: Loads on Submerged Structures Due to LOCA Air Bubbles and Water Jets", General Electric Company, Report No. NEDE-23817-P, September 1978.

l.

9 J

O' 4.3.8-9/4.3.8-10 Revision 0

aum a --+a e-~ . .-,- s-m+ ~ - m- -

4 -m - - -s ,- w ..- - - -.- -wa-I NEDO-21888 1

1 4.3.9 Vent Header Deflector Loads (Later) i i

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NED0-21888 4.4 CONDENSATION OSCILLATION LOADS 4.4.1 Torus Shell Loads (Later) 4.4.2 Loads on Submerged Structures (Later) 4.4.3 Lateral Loads on Downcomers (Later) 4.4.4 Vent System Loads (Later) l t

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NED0-21888 4.5 CHL'GGING LOADS 4.5.1 Torus Shell Loads (Later) 4.5.2 Loads on Submerged Structures (Later) 4.5.3 Lateral Loads on Downcomers (Later) 4.5.4 Vent System Loa'ds (Later)

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i i SECTION 5 SAFETY / RELIEF VALVE DISCHARGE LOADS i

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NEDO-21888 5.0 SAFETY / RELIEF VALVE DISCHARGE LOADS This section specifies the methodology to be used to define loads due to S/RV actuations. Section 5.1 gives an introduction to the S/RV phenomena and loadings. Section 5.2 describes the loads due to S/RV actuatians through T-Quencher discharge devices. Section 5.3 describes the loads due to S/RV actuations through Ramshead discharge devices.

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5.1 INTRODUCTION

When a S/RV actuates, pressure and thrust loads are exerted on the S/RVDL piping and discharge device (Ramshead or T-Quencher). In addition, the expul-sion of wate,r and then air into the suppression pool through the discharge device results in pressure loads on the submerged portion of the torus shell and drag loads on submerged structures. This section describes procedures for determination of loads

resulting from water / air discharge through both T-Quencher and Ramshead devices following S/RV actuation. Procedures are presented for the definition of the following S/RV discharge loads:

e S/RVDL pressure and temperature e Thrust loads on S/RVDL piping e Thrust loads on T-Quencher arms e T-Quencher internal pressure e Water jet loads on submerged structures e Torus shell pressure distribution e- Air bubble induced drag loads on submerged structures In addition, methodology for determining the S/RVDL water reflood height and timing for input to subsequent S/RV actuation analyses is provided.

.A series of analytical models and hand calculational procedures will be used to develop S/RV line unique loads. Flowcharts-depicting the loads / parameters evaluated by each analytical model or procedure and how each methodology inter-acts with the others are presented in Figures 5.1-1 and 5.1-2.

All loads are applicable to S/RVDL's ending in either Raushead or.T-Quencher discharge devices, unless specifically indicated otherwise.

5.1-1 Revision 0 l

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NEDO-21888 These loads / parameters are to be evaluated under each of the following initial conditions:

A. First actuation B. First actuation - leaking S/RV (Ramshead only)

C. Subsequent actuation Section 5.2 presents phenomenological descriptions, load definition proce-dures, assumptions and bases, for each load resulting from a S/RV discharge through a T-Quencher device. Section 5.3 presents similar information for loads resulting from S/RV discharges through Ramshead discharge device .

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- S/RVDL CLE ARING MODEL

' MARK I LOAD DEFINITION REPORT CALCULATES: _

S/RVOL PRESSURE I y PRO *J. DES CALCULATIONAL OCTAIN S/RVDL AND TORUS GEOMETRY AND OPERATING S/RVDL THRUST LOADS CONDITIONS FOR THE CASE S/RVDL TEMPERATURE TO BE EVALUATED T4UENCHER PRES $URE .

SUBMERGED STRUCTURES WATER CLE ARING VELOCITY ,

YWATER CLEARING ACCELERATION C LOADS ON T4UENCHER ARMS DISCH ARGE PRESSURE O

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S/RVDL REFLOOD MODEL 1 P CALCULATES:

S/RVDL REFLOOD HEIGHT TORUS SHELL LOAD MODEL AND TIMING (FOR SUBSEQUENT S/RV CALCULATES:

ACTUATION ANALYSES)

AIR BUBBLE PRESSURE % AIR BUBBLE DRAG MODEL

-. AIR BUBBLE R ADIUS C CALCULATES:

AIR BUBBLE VERTICAL POSITION /

AIR BUBBLE INDUCEO DRAG TORUS SHELL PRESSURE LOADS ON SUBMERGED

n STRUCTURES DISTRIBUTION Q

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a O Figure 5.1-1. T-Quencher Load Definition Scheme

S/RVOL Cl E ARING MODEL MARK I LOAD DEFINITION REPORT

+ CALCULATES:

S/RVOL PRESSURE OBTAIN S/RVOL AND TORUS S/RVDL THRUST LOADS GEOMETRY AND OPERATING .

PROVIDES CALCULATIONAL CONDITIONS FOR THE CASE " PROCEDURE FOR:

WATER CLEARING ACCELERATION TO BE EVALUATED AND VELOCITY : 1 WATER JET LOADS ON

" SUBMERGED STRUCTURES DISCH ARGE PRESSURE HORIZONTAL POSITION OF BUBBLES 1 P Z 0

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HEIGHT AND TIMING AIR BUBBLE VERTICAL ;

POSITION 1 r S/RV AIR BUBBLE PRESSURE AIR BUBBL E DRAG MODEL ATTENUATION MODEL CALCULATES:

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NEDO-21888 5.2 T-QUENCHER '0_ ADS

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5.2.1 S/RV Discharge Line Clearing Transient Loads 5.2.1.1 Load / Parameter Descriptions

a. S/RVDL and T-Quencher Pressure When a S/RV opens, the pressure within the S/RVDL undergoes a transient prior to reaching a steady state value. A transient pressure wave travels back and forth in the line as the pressure continues to increase until the inertia of the water slug in the submerged portion of piping is overcome. During the water clearing transient, the pressures within the discharge pipe and the T-Quencher device reach their maximum values. Following expulsion of the water slug, the peak pressure in the discharge pipe decreases to a quasi-steady state value which is a function of the S/RV steam flow rate and friction along the line upstream of the entrance to the T-Quencher

[} device. Similarly, the T-Quencher internal pressure increases and then decreases to a quasi. steady state value which is a function of the steam flow rate and pressure losses resulting from flow through the holes in the T-Quencher device. Figures 5.2.1-1 and 5.2.1-2 present sample analytical model predictions of S/RVDL and T-Quencher internal pressure.

b. Thrust Loads on the S/RVDL The high flow of steam into the discharge line when an S/RV opens results in the development of a pressure wave at the entrance to the line.

This pressure wave travels to the air / vater interface where it is reflected and returns to the entrance of the line. The wave con-tinues to travel back and forth in the discharge line until a quasi-steady state condition is reached as the pressure differential across the wave approaches zero. The pressurization of the S/RV line results in an acceleration and expulsion of the water in the submerged portion of piping.

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5.2.1-1 Revision 0

NEDO-31888 During the early portion of this transient, a substantial pressure differential exists across the pressure wave. Therefore, when the wave is within a S/RV pipe segment between a pair of elbows there exists a substantial difference in the pressure applied to the inte-rior surface of the elbows on each end of the segment. This pressure differencial, plus momentum effects from steam (or water in initially submerged pipe runs) flowing around elbows in the line, results in transient thrust loads on the S/RV discharge pipe segments. These loads should be considered in the design of S/RV pipe restraints, the connection of the S/RV to the main steam line, and the T-Quencher support system. Figures 5.2.1-3 and 5.2.1-4 presents sample analytical predictions of the transient thrust loading on S/RV dis-charge pipe segments initially filled with gas and water, respectively.

Figure 5.2.1-5 presents the positive sense of the thrust loading on a pipe segment.

c. S/RV Air Bubble Charging Pressure Following an S/RV actuation, the air-steam mixture initially in the S/RVDL is compressed prior to discharge into the pool through the T-Quencher device. The pressure at which the air is discharged to the pool is proportional to the maximum pressure which occurs at the air / water interface. Therefore, the maxiaum predicted pressure (predicted by the S/RVDL clearing model) at the air / water interface is used as input to the evaluation of S/RV air bubble and shell pressures.

d. Mass Flow Rate and Acceleration of Water During S/RV Discharge As the pressure increases in the S/RVDL following S/RV actuation, the water slug in the submerged portion of piping is expelled into the suppression pool. The mass flow rate and acceleration of the water slug during this transient are evaluated using the S/RV discharge line clearing model. One or both of these transient parameters are used as input for evaluation of water jet drag loads and thrust loading 9

5.2.1-2 Revision 0

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(,) on the S/RVDL discharge device. Figures 5.2.1-6 and 5.2.1-7 present examples of analytically predicted water mass flow rate and acceleration, respectively.

5.2.1.2 Bases and Assumptions The S/RV discharge line clearing loads described in the previous section are evaluated using a first principles analytical model (S/RV Discharge line Clearing Model). The model has been shown to' conservatively predict full scale data recorded during extensive in-plant testing, Reference 5.2.1-1. The analytical bases and appropriate model-data comparisons for the S/RV Line Clearing Model are documented in Reference 5.2.1-2.

To ensure that a conservative load definition is obtained, the following assumptions are employed when applying the S/RV Line Clearing Model.

a. The S/RV flow rate is assumed to be 1.225 times the ASME rated S/RV

f. flow. This factor was obtained as follows: A factor of (1/0.9) is

(,,/ applied to the ASME rated flow rate to account for the fact that this value is by definition only 90 percent of the expected S/RV flow (Reference 5.2.1-3). In addition, a factor of 1.05 is applied to account for the allowable uncertainty on the S/RV loss coefficient, Kp, (Reference 5.2.1-3). Finally, a factor of 1.05 is applied for conservatism. The 1.225 multiplier is the product of the factors described above, i.e., (1/0.9) x 1.05 x 1.05 = 1.225.

b. The S/RV main disk stroke time is assumed to be 0.02 second. The S/RV loading most significantly affected by the main disk stroke time is the transient wave thrust load. Shorter stroke times result in higher loading. The value of 0.02 second represents a lower bound of main disk stroke times measured during performance testing of S/RV's of similar design to those installed in Mark I plants.

c. The suppression pool water level is at the maximum value allowed by technical specifications. This assumption results in the maximum

,- initial water leg in the S/RV discharge line which in turn results I,,) in the highest water clearing loads on the S/RVDL and discharge device.

5.2.1-3 Revision 0

i NEDO-21888 i

d. The S/RVDL vacuum breaker does not leak. By assuming the vacuum breaker does not leak, a lower value of S/RVDL to wetwell AP is lh calculated

which results in a longer initial water leg in the discharge line. Since a longer water leg results in h -h . water clearing loads, the assumption of a non-leaking vacuum breaker is conservative,

e. For subsequent S/RV actuation conditiens, the initial water leg in the S/RVDL is assumed to correspond to the maximum S/RVDL reflood**

which is predicted to occur after the minimum predicted time ***

between S/RV actuations (see figure below). The S/RVDL clearing increases as the initial water leg lengths increase. Therefore, the highest water leg that can reasonably be expected to occur coincident with a subsequent S/RV actuatic , is assumed.

USE OF THIS'/ALUE WITHOUT PLANT UNIQUE ANALYSIS MINIMUM TIME BETWEEN ACTUATIONS FROM 3a PLANT UNIQUE ANALYSIS

= 4 L g USE THl3 VALUE WITH PLANI UNIQUE ANALY3 S j

p POOL WATER LEVEL s Y

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TIME FOLLCWING VALVE CLOSURE

=Following the previous S/RV closure, air from the drywell enters the S/RV vacuum breaker and pressurizes the S/RVDL to the drywell pressur<, minus the vacuum breaker set point. If the vacuum breaker leaks, additional drywell air will enter the S/RVDL increasing its internal pressure and further depressing the water leg.

- Prediction =ade by the S/RVDL reflood analytical model (See Section 5.2.3).

- - Prediction made by plant unique analysis. If an analysis is not utilized to determine timing between actuations, the subsequent actuation is assumed to occur at the time corresponding to the maximum predicted reflood.

3.2.1-4 Revision 0

NEDO-21888 5.2.1.3 Load Definition Procedure

)

A summary of key parameters which influence S/RV discharge line clearing loads is shown below.

a. S/RVDL initial water leg

b. S/RVDL air volume

. c. S/RVDL diameter

d. S/RV steam flow rate f

e. S/RVDL initial temperature

f. S/RVDL configuration and hydrodynamic losses

g. S/RV main disk stroke time.

I Using the assumptions presented in Section 5.2.1.2, unique values for these parameters are established for each S/RVDL in the plant under initial conditions A (First Actuation) and C (Subsequent Actuation).

These values are then used as input to the S/P.VDL line clearing analytical model (Reference 5.2.1-2) to obtain each tressient S/RVDL clearing load and parameter.

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. ) REFERENCES FOR SECTION 5.2.1 2

5.2.1-1 R. A. Asai, et al. , "!! ark I Containment Program Final Report -

Monticello T-Quencher Test, General Electric Company," 2eport No. NEDE-21864, Addendum 1 July 1978.

5.2.1-2 A. J. Wheeler, " Comparison of Analytical Model for Computing

' Safety / Relief Valve Discharge Line Transient Pressures and Forces to Monticello T-Quencher Test Data." General Electric Company, Report No. NEL2-23749-P, Addendum 1, to be published.

5.2.1-3 ASME Boiler and. Pressure Vessel Code,Section III, A

Subsection NB-7000. Ava',lable from the American Society of Fechanical Engineer, New York, New York.

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5.2.2 Torus Shell . essure 5.2.1.1 Load Description Prior to the initial actuation of an S/RV caused by a normal operational transient, the S/RV discharge lines contain air at atmospheric pressure and suppression pool water in the submerged portion of piping. Following S/RV actuation, steam enters the S/RVDL compressing the air within the line expelling the water slug and discharging the air into the pool. The compressed air charges bubbles which expand _ resulting in an outward motion of the surrounding pool water. The outward momentum of the pool water causes the pressure within the bubbles to drop below the ambient pool pressure. The negative bubble pressure slows and reverses the bubble exr>ansion and the pool water begins to move inward. The inward momentum of the water results in a compression of the bubbles to a pressure above ambient.

The expansion and compressicn of the air bubbles continues until they rise and break through the pool surface. The positive and negative dynamic pressures developed within these bubbles result in an oscillatory, attenuated pressure loading on the torus shell.

J Figure 5.2.2-1 is an example of the analytically predicted torus shell pressure transient resulting from a S/RV discharge through a T-Quencher device. Fig-ures 5.2.2-2 and 5.2.2-3 are. examples of analytically predicted torus shell longitudinal and radial pressure distributions, respectively.

5.2.2.2 Bases and Assumptions The torus shell pressure loading is evaluated using a first principles analytical model (Torus Shell Load Model). The model incorporates information obtained from extensive full scale in-plant T-Quencher testing (Reference 5.2.2-1) as well as parametric testing of a 1/4 scale T-Quencher discharge system (Reference 5.2.2-2). The model has been shown to conservatively predict avail-able full scale shell pressure data and to correctly predict the parametric trends measured at 1/4 scale. The analytical and/or empirical bases for the Torus Shell Lord Model are documented in Reference 5.2.2-3.

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5.2.2-1 Revision 0

NEDO-21888 To ensure that a conservative load definition is obtained, the same assumptions are employed when applying the Torus Shell Load Model for design as were g

used with the S/RVDL Clearing Model (See Section 5.2.1.2) with the following additions:

a. For multiple S/RV actuations, it is assu=ed that the pressure loading at a given location on the torus shell is equal to the square root of the sum of the squares (SRSS) of the peak pressure loads pre-dicted to occur at that location for each individual S/RV actuation.

Review of torus shell pressure data obtained during extensive in-plant S/RV testing (Reference 5.2.2-1) provides justification for this assumption. During in plant T-Quencher testing, single valve accuations resulted in higher peak shell pressures than did the actuation of three adjacent S/RVs.

In addition, the following conservative assumptions / methods are applied when evaluating the torus shell loading for each single valve actuation prior to combining the results to obtain multiple valve loads.

(1) Under test conditions, the analytical models conservatively pre-dict available measured data (single valve actuations).

(2) The design S/RV flow rate is 1.225 times the ASME rated flow.

(3) The pool water level is assumed to be at its maximum value allowed by technical specifications.

(4) The pool temperature is assumed to be at its maximum value for the condition under evaluation.

b. The predicted frequency of the torus shell pressure waveform is accurate to 15 percent for first actuations and 230 percent for leaking valve and subsequent actuations.

l 5.2.2-2 Revision 0 l

NEDO-21888 These uncertainties are based on consideration of scatter observed in full scale testing (Reference 5.2.2-1) and the accuracy with which the Torus Shell Load Model has been shown to predict shell loading frequencies over the range of S/RVDL geometries and operating conditions (Reference 5.2.2-3).

5.2.2.3 Load Definition Procedure The following are key parameters which influence the torus shell pressure losding.

a. S/RV air bubble charging pressure (from S/RV discharge line clearing model)

b. Suppression pool temperature

c. Suppression pool geccetry

d. T-Quencher location in the suppression pool.

O The assumptions in Section 5.2.2.2 are used to establish plant, S/RV line, and initial condition unique values for these parameters. These values are then input to the Torus Shell Load Model (Reference 5.2.2-3) to evaluate the torus shell pressure distribution and transient pressure waveform for a single valve actuation for initial conditions A and C. The peak pressure values pre-dicted at each shall location (due to single valve actuations) are added by the SRSS method to obtain the integrated torus shell loading for each,of the load case / initial condition combinations shown in Table 3.0-1.

The structural evaluation is performed using the predicted pressure waveform within the appropriate uncertainty range (stretched or compressed time scale) which is expected to result in the mnvimum structural response.

5.2.2-3/5.2.2-4 Revision 0

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5.2.2-7/5.2.2-8 Jevision 0

NEDO-21888 REFERENCES FOR SECTION 5.2.2 5.2.2-1 R. A. Asai, et al., " Mark I Containment Program Final Report -

Monticello T-Quencher Test," General Electric Company, Report No. NEDE-21864, July 1978.

5.2.2-2 C. T. Sawyer, " Mark I Containment Program Final Report - 1/4 Scale T-Quencher Test," General Electric Company, Report No. NEDE-24549, December 1978.

5.2.2-3 P. Valandani, " Mark I Containment Program - Analytical Model for Computing Air Bubble and Boundary Pressures Resulting from a S/RV Dischar:1e Through a T-Quencher Device," General Electric Company, Report No. NEDE-21878, December 1978.

5.2.2-4 E. A. Buzek, et al., " Final Raport - In-Plant Safety / Relief Valve Discharge Load Test - Monticello Plant," General Electric Company, Report No. NEDC-21581, June 1977.

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,3 5.2.3 S/RVDL Reflood Transient b

5.2.3.1 Description of Phenomena Following closure of a S/RV the steam pressure in the S/RVDL decreases rapidly as the steam flows into the suppression pool. At a sufficiently low steam pressure, pool water will reenter the S/RVDL. A rapid depressurization of the line then occurs as steam in the line is condensed by the inflowing water.

This depressurization causes the water to reflood into the S/RVDL and the vacuum breaker valve on the S/RVDL to open, allowing drywell atmosphere to enter the line. The reflooding water may rise in the line to a level somewhat above its initial pre-S/RV actuation level before equilibrium is reestablished.

The actual reflood water level depends primarily on the size of the S/RVDL vacuum breaker valve to allow a rapid repressurization of the line.

A subsequent actuation may occur following closure of the S/RV. The loads developed during the subsequent actuation will depend on the conditions in the S/RVDL at the time of the actuation. Subsequent actuation loads on the S/RVDL g and discharge device are proportional to the initial water level in the S/RVDL.

U Therefore, the peak reflood water level predicted to occur at a point in time af ter the minimum predicted time

between S/RV actuations is used as the initial water leg in the S/RVDL for subsequent actuation load calculations.

An example of analytically predicted S/RVDL reflood behavior is illustrated in Figure 5.2.3-1. which presents water column length in the S/RVDL as a function of time. It is seen that the water column length increases continuously from the time the reflood transient begins until the first peak is reached. The delay between valve closure and the initiation of reflood is due to pressure decay in the S/RVDL following S/RV closure.

Prediction made by plant unique analysis. If an analysis is not utilized to determine timing between actuations, the subsequent actuation is assumed to occur at the time corresponding to the maximum predicted reflood.

5.2.3-1 Revision 0

NEDO-21888 5.2.3.2 Bases / Assumptions The S/RV discharge line reflood transient. is evaluated using a first principles analftical medel (Reflood Model). The medel has been shewn to conservatively predict the maximum S/RVDL reflood measured during full scale in-plant testing.

The analytical bases and appropriate model data comparisons for the reflood model are documented in Reference 5.2.3-1.

The following assumptions will ensure that a conservative reflood prediction is obtained:

a. The S/RVDL vacuum breaker is located in the drywell at a point above the maximum S/RVDL reflood level.

b. The suppression pool water level is at the maximum value allcwed by technical specifications. This assumption maximizes the pool pressure (elevation head) at the discharge device elevation resulting in the maximum net force driving the water slug up the S/RVDL.

c. The delay time between complete S/RV closure and the beginning of the reflood transient is a linear function of the entire S/RV discharge line volume. The delay time used for the evaluation of reflood timing is based on measured in-plant S/RVDL depressurization transients following S/RV closure (Ref erence 5.2.3-2). This delay time is mul-tiplied by the ratio of the S/RVDL volume to the volume of the line tested to provide line unique delay times.

5.2.3.3 Procedure for Evaluating the S/RVDL Reflood Transient A summary of key parameters which influence the S/RVDL reflood transient is shown below:

a. Drywell and wetwell pressure

b. S/RVDL vacuum breaker effective flew area l

d. Vacuum breaker opening time g

5.2.3-2 Revision 0

d. Vacuum breaker set point

e. S/RVDL geometry and hydrodynamic losses

f. Vacuum breaker location along the S/RVDL

g. Whether air or steam enters the vacuum breaker

h. Condensation heat tra sfer coefficient on the interface of the reflooding water slug

i. Pool water temperature

j. Submergence depth of the discharge device.

The assumptions presented in Section 5.2.3.2 are used to establish S/RV dis-charge line unique values for these parameters. These values are then used as input to the S/RVDL reflood analytical model (Reference 5.2.3-1) to evaluate the discharge line reflood transient.

Once the reflood transient predictions have been obtained, a delay time is added at the beginning of each transient to account for time required for the S/RVDL to depressurize following S/RV closure and therefore allow the reflood transient to begin. This delay time is determined using the following relation:

Delay Time = C V see where V =

Volume of S/RVDL under consideration, ft C = 0.008,

- 3 determined from Monticello full scale T-Quencher ft tests (Reference 5.2.3-1)

The maximum S/RVDL reflood, which is predicted to occur at a time in the transient af ter the minimum predicted time between S/RV actuations (based on plant unique analyses), is used as the initial water leg for subsequent S/RV actuation evaluation. Optionally, if plant unique analyses ars not perfor=ed,

("T the maximum reflood occurring at any time during the reflood transient is to be

-\ "'] used as the initial water leg for such evaluations.

5.2.3-3/5.2.3-4 Revision 0

O O O i5 i

GHAPH OF S/HVDL REFLOOD io -

WATEH t EVEL VERSUS TIME t O. S/RV CLOSURE

s g. INITIAL ENTRY OF WATER INTO i

DISCHARGE DEVICE. STAHT OF ANALYTICAL MODEL PflEDICTION

' 8 ~ '2 DISCliARGE DEVICE HEFLOODED.

6 -

WATER REENTHY INTO S/RVDL

=

J o

vs $

ta es 3 o A POOL WAT Eli L E VEL O e o u u j H s w

< u a w

.! .hJ <

w s i S -5 -

-10 -

i

-is l l l l o '2

te 2 4 6 a

, 30 j $ TIME AFTER S/HV CLOStfilE (seci r

in r

O

s i Figure 5.2.3-1. Sample Prediction of S/ItVDL Reflood Transient g

i

t NEDO-21888 S

1 REFERENCES FOR SECTION 5.2.3 5.2.3-1 .A. J. Wheeler, " Mark I Containment Program, Analytical Model for

[ Computing Transient S/RV Discharge Line Reflood," General Electric

. Company, Report No. 3EDE-23898-P, October 1978 1

1

5.2.3-2 R. A. Asai,. et al. , " Mark I Containment Program Final Report -

Monticello T-Quencher Test," General Electric Company, Report No. NEDE-21864-P, July 1978.

i 4

1-

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i .

r-1 i

i i

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it a

4 i..

i 5.2.3-7/5.2.3-8 Revision 0 i'

- , , . . - , _ _ , . , , , . . , _ - . . . _ . - . . - - _ . - . _ . - _ ~ . _ . _ . _ . - . ,_ .._ _ .. _ _ ._ ___ , _ -... _ _ __ _ _ _ . _ _ _ . _ _ - .._,

NEDO-21888 I 5.2.4 T-Ouencher Water Jet Loads on Submerged Structures 3.2.4.1 ' Load Description When an S/RV is actuated, water initially contained in the submerged portion of the S/RV discharge line (S/RVDL) is forced out of the T-Quencher arm through ths ata holes forming orifice jets. Some distance downstream the orifice jets will merge to form column jets. Further downstream, the column jets will merge to form the quencher arm jets. As soon as the water flow through the arm holes ceases, the quencher arm jet velocity will decrease rapidly and the jet will penetrate a limited distance into the pool. Figure 5.2.4-1 shows the phases of quencher jet formation and decay. These T-Quencher water jets will create drag loads on nearby submerged structures which are within the jet path.

5.2.4.2 Bases / Assumptions The bases for establishing the water jet dynamics are empirical mathematical D models (Reference 5.2.4-2) obtained from stead? -state submerged jet theory at various jet zones (orifice jet, column jet, quencher arm jet and jet penetration). Reference 5.2.4-1 provides the first principles methodology from which the jet properties for water jets from the quencher arms and endcap holes can be conservatively determined. ;

To ensure that a conservative T-Quencher water jet load definition is obtained, the following assumptions are employed when applying the T-Quencher v.:er jet model.

a. The quencher arm is divided into six sections along the axis of the arm. Each section assumes a uniform hole pattern. For arms with endcap holes, there is an sdditional endcap section (see Figure 5.2.4-2).

0 V

l 5.2.4-1 Revision 0

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

b. As the air / water interface reaches the first column of holes on the arm, the water acceleration through all the holes is assu=ed constant and equal to the value when the interface is at the first column of holes.

c. The maximum value of the water velocity through the quencher arm holes found in a section of the quencher arm is conserva-tively used to calculate velocity characteristics for the water jet from that section.

d. The t.bove water velocity is further conservatively assumed constant with time and applied in the calculation throughout the clearing period of that section.

Higher . rater velocities through the arm holes result in higher =ocentum in the water jet and larger drag loads on the structures.

l 5.2.4.3 Load Definition Procedure

~

A summary of the key parameters which influence the T-Quencher water jet G

loads is shown below;

a. Maximum water velocity through the quencher arm holes.

This quantity depends on the S/RVDL clearing transient at the initial conditicos to be analyzed. The S/RVDL water clearing model is used to determine the water jet velocity and acceleration up to the time that the first column of holes is uncosered. The water jet momentum depends on this velocity.

t l b. Distance from the structure te the quencher arm.

The velocity at the structure locatf on depends on this distance because the water jet velocity decreases rapidly with the distance from the arm. The drag load on the structure is proportional to the square of the local velocity at the structure.

5.2.4-2 Revision 0 O

NEDO-21888 O c. Structure geometry.

The drag load is proportional to the structure's projected area and the drag coefficient which are functions of the structure geometry.

O l

O 5.2.4-3/5.2.4-4 Revision 0

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(s/ !

5.2.4-5 Revisicn 0 1

l l l l l l 1

/

w , j SECTION NO F ,,

= (E fdDCAP JLT) m"

" g r

+ 0 co 8 SECTION SECTION SECT ION SECTIOfJ SECTION SECTION NO 1 NO 2 NO 3 NO. 4 NO $ NO. 6 NOTE FOH T GUEhCHEH ARM WITH NO ENDCAP tsOL ES. THE HE IS NO SECTION NO. 7

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Figure 5.2.4-2. Jet Sectic,s Along the Quenclier Arta on Torus Plan View O O O

MEDO-21888

.O REFERENCES FOR SECTION 5.2.4 5.2.4.1 " Analytical Model for T-Quencher k'acer Jet Loads on Submerged Structures", Prepared for General Electric Company by Nuclear a

Services Corporation and Southwest Research Institute Report,

Report No. 24589-P, to be published.

5.2.4-2 F. J. Moody, " Analytical Model for Liquid Jet Properties for Predicting Forces on Rigid Submerged Structures", General Electric Company, Report No. NEDE-21472, September 1977.

1 m

i I

I. l I

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5.2.4-7/5.2.4-8 Revision 0

() 5.2.5 T-Ouencher Bubble-Induced Drag Loads on Submerged Structures Oscillating bubbles resulting f rom S/RV actuation create an unsteady three-dimensional flow field and ther2 fore induce acceleration and standard drag forces on the submerged structures in the suppression pool.

This section addressos the load definition procedures for loads on sub-merged structures due to T-Quencher S/RV bubbles. The following subsections discuss the methodology to define the loads; the bases and assumptions on which the analytical models are developed and verified; and the selection of key parameters from typical plants for which the design load on sub-merged structures can be obtained.

5.2.5.1 Bases and Assumptions The bases of the analytical model are presented in Reference 5.2.5-1. The major assumptions are summarized as follows:

(O s,) a. The total drag on submerged structures is the sum of the acceleration and standard drags,

b. The flow fields caused by T-Quencher S/RV bubbles are described by fluid sources. The presence of boundaries is accounted for by using the method of images.

c. Source strengths are determined from the T-Quencher bubble dynamics model-(Reference 5.2.5-2).

5.2.5.2 Load Definition Detailed procedures to calculate drag forces on submerged structures are described in Reference 5.2.5-1 which provides the methodology of utilizing the unsteady flow field generated by the T-Quencher S/RV bubbles in the s 1 5.2.5-1 Revision 0 l

l

?

l 4#

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<g A%

/////

4'RQ

'+h.C TEST TARGET (MT-3) l.0 !# E4 D3a O

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ll1.8 l.25 1.4 l 1.6 8

7 MICROCOPY RESOLUTION TEST CHART

S(,;;

Ni!,', zzz;z*' , .

i

NED0-21888 suppression pool, to obtain drag loads on submerged structures, such as pipes, beams, etc.

The load definition first requires the establishment of the unsteady flow field within the torus by simulating the S/RV bubbles by point sources with strengths determined from the T-Quencher S/RV bubble dynamics model (Reference 5.2.5-2). Then, drag loads on submerged structures are calcu-laced based on the previously established unsteady flow field,

a. Flow Field Establishment An oscillating T-Quencher S/RV bubble is considered as a fluid source with source strength determined from the T-Quencher S/RV bubble dynamics mod.1 (Reference 5.2.5-2). The T-Ouencher S/RV bubbles in the :'. ark I torus are simulated by multiple sources in a finite pool. By using potential flow theory and the method of images to account for the effects of solid walls and the free surface, the flow field within the torus is estab: _shed.

b. Drag Loads Evaluation O

The drag force on a submerged structure consists of two compo-nents: the standard drag and the acceleration drag. The standard drag load is computed on a locally uniform flow field evaluated at the local velocity. The acceleration drag load is evaluated based on an inviscid, uniform but unsteady (accelerat-ing) flow field. The sum of these two drag forces gives the total drag load on a submerged structure.

5.2.5.3 Selection of Key Parameters for Load Evaluations This section outlines the key parameters which influence the magnitude and characteristics of the T-Quancher S/RV bubble drag loads on submerged struc-tures. The procedure used to specify the input for the submerged structure drag load model is also discussed.

O, 5.2.5-2 Revision 0

.( ) The plant specific suppression pool geometry should first be identified as follows: ,

a. Torus shell dimensions

b. Torus water level

c. Relative locations of S/RV discharge lines in the suppression pool.

Then, the initial conditions of each S/RV discharge line are specified as follows:

a. Initial pipe pressure

b. Initial volume of air-steam mixture.

Next, the S/RVDL Clearing Model and the T-Quencher S/RV Bubble Dynamics Model ,

are applied as discussed in Sections 5.2.1 and 5.2.2 to determine the following:

a. S/RVDL discharge pressure

b. Clearing time

c. T-Quencher bubble pressure and bubble radius history.

Af ter dividing the structure into appropriate sections for more precise calculation, information on the structure section locations, orientations, acceleration drag volumes, projected areas and the standard drag coefficients are obtained.

With the above parameters identified, the dimensions of the pool model, the coordinates of the initial locations of the T-Quencher S/RV bubbles, and l

the structure sections are calculated. '

O V

5.2.5-3 Revision 0

EEDO-21888 With the input parameters identified and calculated, the submerged structure drag load model is used to evaluate the resultant transient drag forces on each structure sections being analyzed. Typical results are shown in Figures 5.2.5-1 and 5.2.5-2. These figures show the T-Quencher bubble drag forces on a downcemer in the radial (X) and tangential (Z) directions as a function of time.

O O

5.2.5-4 Revision 0 l

l

i 5

=

0 E

2o -

O i

~

1

-s -

~

1 I I I I 0 046 0.12 0.18 0.24 0.30 0.36 TIME AFTER BUS 8LE ENTERS POOL (sec)

Figure 5.2.5-1. Sample Predicted Time History of Total X-Forces on Downcomer 5.2.5-5 Revision 0 l _

NEDO-21888 O l l

22s I

150 -

A n -

=

8 0

o 8o -

! 9 5

S

-n -

-tso

! I I I I

- 22s 0.36 O O 06 0.t 2 0.18 0.24 0.30 TIME AFTER 8U58LE ENTERS POOL (sec) l l ,

l l

Figure 5.2.5-2. Sample Predicted Time liistory of Total Z-Forces )

on Downcomer 1 0

5.2.5-6 Revision 0

T NEDO-21888 ,

3 l l( ) - REFERENCES FOR SECTION 5.2.5 F 3.2.5-1 F. J. Moody, et al., " Analytical Model for Estimating Drag Forces I

on Rigid Submerged Structures Caused by LOCA and Safety Relief Valve Ramshead Air Discharges Supplement for T-Quencher Air '

-Discharges", General Elcetric Company, Report No. NEDO-21471, to be published.

5.2.5-2 P. Valandani, " Mark I Containment Program, Analytical Model for Computing Air Bubble and Boundary Pressures Resulting from a S/RV Discharge Through a T-Quencher Device", General Electric I- Company, Report No. NEDE-21878, December 1978.

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LO 5.2.5-7/5.2.5-8 Rev'ision 0

() 5.2.6 Thrust Loads on T-Quencher Arms

. 5.2.6.1 Load Descriptions

a. Thrust Loads Along Axis of T-Quencher Arms Following S/RV actuation the pressurization of the S/RVDL causes the water initially in the T-Quencher and submerged portion of piping to be accelerated and expelled through the T-Quencher arm holes into the suppression pool. The redirection of flow of water 90 degrees out the holes and the internal pressure of the arms results in thrust loads on the arm endcaps. By assuming a difference in the water flow rate into each arm of the T-Quencher a not thrust load is obtained acting along the axis of the T-Quencher device.

One arm of the T-Quencher may have holes in the endcap. For quenchers of this design, an additional component of endcap thrt st

() is produced 'due to differences in momentum of the fluid (water ac steam) leaving the device.

An example of the predicted individual end cap thrust loads and the net thrust loading on a T-Quencher with endcap holes is shown in Figure 5.2.6-1.

b. Thrust Loading Perpendicular to the T-Quencher Arms As the water is expelled from the T-Quencher device, irregularity in the shape of the gas / water interface (e.g. the interface may act be perpendicular to the quencher arm) may result in a net thrust load perpendicular to the arm due to momentum imbalance. An example of the predicted thrust loading perpendicular to the quencher arm is shown in Figure 5.6.2-2.

A N,)

5.2.6-1 Revision 0 1

NEDO-21888 5.2.6.2 Bases and Assumptions

a. Thrust Loading Along Axis of T-Quencher Arms The thrust load on the T-Quencher along the arm axis is determined by performing a momentum balance (in the direction of the arm axis) on each arm separately to obtain the individual endcap thrust loadings.

The difference in the endcap loads is then computed to determine the net thrust load on the T-Quencher device as a whole.

The analyses is separated into three phases:

(1) Before the gas / water i;~arface reaches the first row of holes (2) After the gas / water interface reaches the first row of holes (3) After water clearing.

During the first phase the S/RVDL Clearing Model (see Section 5.2.1) is used to determine the water velocity and acceleration as it enters the T-Quencher arms.

The velocity and acceleration of fluid flow from the holes during this period is determined using conservation of mass and an assumed flow split between the arms.

The inputs required to perform a momentum balance on each quencher arm during the second phase of the analysis are determined using the following assumptions:

(1) Slug flew is assumed.

(2) The acceleration of the water flowing from the quencher holes is assumed to remain constant in time at the value predicted S.2.6-2 Revision 0 O

(3

when the interface reaches the first row of quencher holes.

The gas driving the interface will begin to escape through the holes into the pool at this point. Therefore, the driving pressure will decrease, which results in a reduction in water clearing acceleration from the holes.

(3) In evaluatihg the internal pressure on the endcaps, the pressure in the arm is assumed to remain constant at the value predicted when the interface reaches the first row of holes. This is a bounding assumption since the pressure will actually drop significantly as the gas escapes through the quencher holes.

(4) The water velocity through the holes is the same for all holes for any point in time.

(5) The water acceleration through the holes is the same for all holes for any point in time.

(6) Following water expulsion, the steam flew through the end cap holes is assumed to be choked.

b. Thrust Loading Perpendicular to the Quencher Arms

The thrust loading perpendicular to the T-Quencher arms is determined by performing a momentum balance on each arm in the direction perpendicular to the arm axis. The analysis is performed using the following assumptions:

(1) The velocity and acceleration of water out the holes is computed using the assumptions in Section 5.2.6.2(a).

(2) As the gas / water interface travels down the quencher art a small length of the quencher arm has water exiting one side while gas exits the other side of the arm. This assumption s

is based on evaluation of movies taken of water / gas discharges s/

from T-Quencher devices during the 1/4 scale T-Quencher Test Program (Reference 5.2.6-1).

5.2.6-3 Revision 0

NED0-21888 (3) The gas comentum flow through the holes in the region of the arm immediately behind the gas / water interface is neglected.

As the interface travels down the arm the gas exiting the holes immediately behind the interface will not have sufficient time to establish a choked flow condition before water clears the holes directly opposite them. By assuming the gas velocity equals zero, a minimum gas momentum out one side of the arm is obtained and the net thrust loading is thus maximized.

5.2.6.3 Load Definition Procedure A summary of key parameters which influence the thrust loads is given below.

a. T-Quencher arm peometry (i.e., with or without endcap holes)

b. Water-mass flow rate and cceeleration as a function of time

c. Time at which,the air / water interface reaches the holes.

O These parameters are evaluated for initial condition C (see Table 3.0-1) and the quencher arm geometry under consideration.

a. Thrust Loads Along T-Quencher Arm Axis Since the gas / water interface location is being followed in the transient, the time at which the interface reaches the first row of holes should be determined from the S/RV discharge line clearing model output. This time separates the two phases of the method-ology. The water velocity and acceleration up to this time are readily available from the model output (see Section 5.2.1) and can be applied in the calculation of the thrust load 'or the "before" phase. The water velocity and acceleration predicted at the time the interface reaches the holes becomes the initial '

conditions for the "after" phase. The water velocity through the holes, the interface location, velocity and acceleration, and the 5.2.6-4 Revision 0 O

NEDO-21838 1

O(j average water acceleration in the arm are determined as a function lof time as the water clears the arm. Following water expulsion, steam is assumed to flow through the end cap holes. The thrust load on each

endcap and the net thrust load on the device is then determined as a function of time. Finally the three phases are combined to obtain the i

j entire thrust load transient on the T-Quencher device along the arm axis.

. b. Thrust Loads Perpendicular to the T-Quencher Arms Using the assumptions in Section 5.2.6.2(h), the net momentum flow per-

! pendicular to the arm is evaluated as the gas / water interface travels

! down the arm. From this, the net thrust load application point is deter-mined as a function of time.

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5.2.6-5/5.2.6-6 Revision 0

NED0-21888 O

40.000 NET THRUST ON T4UENCHER ENO CAP HOLES 00,000 -

L )

C \

20.000 -

g 10,000 -

O f I I I ' '

)

l 70,000 THRUST LOAD ON END CAP WITH HOLES END CAP HOLES l

60,000 -

r2 t'

50,000 -

(48% FLOW) 1

= )

$ 40 ,000 - '

o l 3 00,000 -

j 20,000 -

E l 10,000 -

' I ! ! I I 0

I l

70,000 !

THRUST LOAD ON END CAP WITHOUT HOLES 60,000 -

UJ 50,000 -

(52% FLOW) 40,000 -

30,000 -

20,000 -

10,000 -

0 ' I I I I I O 0.04 0.08 0.12 0.16 0.20 0.24 0.28 0.32 TIME FROM S/RV OPEN (sec)

Figure 5.2.6-1 Thrust Loads on Arm Endcaps 1

O S.2.6-7 Revision 0 l

1

END OF HOLE PATTERN

- 8.0 s

- 7.0 1;i

= U e o j - 6.0 e o e c E 2 o z

O

z y h

- 5.0 fo l e :

5 0 l 5 E l 9 t~

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- 1.0 l i 1 I l 0 0 0.05 0.1 0.15 0.2 0.25 0.35 TIME FROM S/RV OPEN (sec)

Figure 5.2.6-2. Generalized Shape of Thrust Loading Transient and Application Point with Time 5.2.6-8 O

NEDO-21888 REFERENCES FOR SECTION 5.2.6 5.2.6-1 C. T. Sawyer, " Mark I Containment Program, 1/4 Scale T-Quencher Test," General Electric Company, Report No. NEDE-24549, December 1978.

5.2.6-9/5.2.6-10 Revision 0

- 5.2.7 Maximum S/RVDL and Discharge Device Pipe Wall Temperature The S/RV discharge pipe and discharge device are subjected to thermal expansion loading during a S/RV discharge which bounds any other thermal loading resulting from normal or accident conditions. The maximum (bulk pipe wall) temperature at any axial location along the discharge pipe is a function of the steady state internal pressure at that location during steam flow. Figure 5.2.7-1 presents an example of the predicted S/RVDL steady state temperature distribution during an S/RV discharge.

5.2.7.1 Lases / Assumptions The maximum S/RV discharge pipe temperature (upstream cf the discharge device) at a given axial location along the line is assumed to be the saturation tenperature corresponding to the steady state pressure predicted by the S/RVDL clearing model at that location. The analytical and empirical bases for :he S/RVDL clearing model are documented in Reference 5.2.7-1.

The maximum discharge device temperature is assumed to be the saturation

(_) temperature corresponding to the maximum steady state stagnation pressure of steam expected to occur within the device.

The maximum steam stagnation pressure is assumed to be that which.would be required to develop the maximutt expected S/RV steam flow (for all Mark I S/RV's) through the discharge device exit holes under choked flow conditions.

The steam pressure density product required to do this was evaluated using the following relation:

l

~

k+1 ,

t l 1 m 2 2(k-1) 2 lbf lbm po,o ,

kg e 0.6 A, ,

k+1 ft 3

l P =

stagnation steam pressure in discharge device, lbf/ft 2 o

g

=

stagnation steam density in discharge device, lbm/ft ;

1 -

v l 5.2.7-1 Revision 0

NED0-21888 ,

0.6 =

exit hole discharge coefficient

=

A, total exit hole area, ft k =

ratio of specific heats g = 9 gravitational constant, (lbca-f t)/(1bf-sec-)

& =

steam flow through S/RV throat. -

see Comparison of the steady state S/RVDL temperature distribution and the dis-charge device temperature measured during extended S/RV discharges perforrad during full scale in-plant testing, Reference 5.2.7-2, indicates that this approach provides a bounding thermal loading definition.

5.2.7.3 Load Definition Procedure The S/RV discharge line clearing model is used to evaluate the steady state pressure distribution along the discharge pipe applying procedures presented in Section 5.2.1.3. This pressure distribution is then converted to a temperature distribution using the steam tables and the assumption that the steam is saturated. The resulting temperature distribution is applied from the S/RV exit to the discharge device entrance.

The discharge device temperature is defined generically as follows:

T-Quencher - 370*F Ramshead - 390*F O

5.2.7-2 Revision 0

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5.2.7-3/5.2.7-4 Revision o l

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4 NEDO-21888 i REFERENCES FOR SECTION 5.2.7 5.2.7-1 A. J. Wheeler, " comparison of Analytical Hodel for computing safety /

Relief Valve Discharge Line Transient Pressures and Forces.co Monticello T-Quencher Data," General Electric Company, Re}' ort No. NEDE-23749-P, Addendum 1, ver. ember 1978.

[

5.2.7 R. A. Asai, et al., Mark I Containment Program Final Report, Monticello T4uencher Test," General Electric Company, Report No. NEDE-21864, July 1978.

1-J J

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5.2.7-5/5.2.7-6 Revision 0

NEDO-21888 5.2.8 ratigue Cycles

[

The actuation of S/RV's is an expected occurrence in plant operation. During the plant lifeti=e the containment will be subjected to nanerous S/RV actuatien transients. Therefore, fatigue loading due to S/RV discharge must be considered in the structural evaluation of the components affected.

The number of fatigue cycles for each of the various S/RV discharge loads dis-cussed in the previous sections is a function of the number of S/RV actuations assumed to occur during the plant lifetime. 'The total number of S/RV actuations x

is a function of the number or-feactor transients which occur and the number of S/RV actuations for each. i The total number of transients from nonaccident operating conditions shall be based on extrapolation of plant operating history. In addition, S/RV actua-tions due to accident (SBA or IE') conditions shall be considered.

-~ .

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5.2.8-1/5.2.8-2 Revision 0

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-5.3 RAMSHEAD LOADS l- The following sections define the S/RV discharge loads for a Ramshead l discharge device. Since many of the loads are similar to those previously i defined for a T-Quencher discharge device, some of the Ramshead sections

- will refer to the appropriate T-Quencher section.

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5.3.1 S/RV Discharge Line Clearing Transient Loads The analytical codel for S/RV discharge line clearing calculates S/RVDL loads resalting from S/RV actuation. In addition, it calculates various parameters for use as input to other S/RV loading analyses. The analytical basis for this model is documented in Reference 5.3.1-1 and verified against test data contained in Reference 5.3.1-2.

5.3.1.1 Load / Parameter Descriptions

a. S/RVDL Pressure When an S/RV opens, the pressure within the S/RVDL undergoes a transient prior to reaching a steady state value. A transient pressure wave travels back and forth in the line as the pressure continues to increase until the inertia of the water slug in the

,O submerged portion of piping is overcome.

During the water clear-k.s/ ing transient, the pressures within the discharge pipe reach their maximum values. Following expulsion of the water slug, j- the peak pressure in the discharge pipe (the peak S/RVDL pressure occurs at the S/RV exit) approaches a quasi-steady state value which is a function of the S/RV steam flow rate and friction along the line from the S/RV to the upstream entrance to the Ramshead device. Figure 5.3.1-1 presents a sample of a pre-diction of the S/RVDL pressure transient at the S/RV exit.

b. Thrust Loads on the S/RVDL WhenaSIRVopens,thehighflowrateofsteamintothedischarge line results in the development of a pressure wave at the entrance to the line. This pressure wave travels to the air / water inter-face where it is reflected back to the entrance to the line.

This wave continues to travel back and forth in the discharge line until a steady state condition is reached when the pressure differential across the wave approaches zero.

O O

5.3.1-1 Revision 0

NEDO-21888 During the early por. ton of this transient, a substantial pressure differential exists across the pressure wave. There-fore, when the wave is within a S/RV pipe segment there exists a substantial difference in the pressure applied to the interior surface of the elbows on each end of the segment. This pressure differential, plus momentum effects from steam flowing around eloows in the line, results in transient thrust loads on the S/RV discharge pipe segments. These loads are considered in the design of pipe restraints, the connection of the S/RV to the main steam line, and the Ramshead support system. Figure 5.3.1-2 presents a sample analytical model prediction of the thrust loading on a typical S/RV discharge pipe segment.

As the water slug in the submerged portien of the S/RVDL is ex-pelled following a S/RV actuation, the redirection of water flew in the Ramshead device results in a thrust load transient on the last pipe segment. Figure 5.3.1-3 presents the location atd posi-tive sense of this loading. Figure 5.3.1-4 presents sample pre-diction of the water clearing thrust load transient on the last pipe segment.

c. Ramshead Discharge Pressure Following a S/RV actuation, the air / steam mixture initially in the S/RVDL is compressed prior to discharge into the pool through the Ramshead device. The predicted pressure after water clearing at the Ramshead discharge is used as input to the analysis of S/RV air bubble drag loads and torus shell pressures.

d. Mass Flow Rate of Water During Discharge When the S/RVDL and Ramshead are modeled in the S/RVDL clearing model, the mass flow rate of water through the Ramshead exit is calculated as a function of time.

O

5. 3.1- 2 Revision 0

NEDO-21888 (v~'y .

The calculated mass flow rate of water :s used as input to water jet load analyses as well as analysis of the water clearing thrust loads on elbows in the S/RVDL. Figure 5.3.1-5 presents an example of a water mass flow rate transient.

e. Mass Flow Acceleration cf Water Through the Ramshead The mass flow acceleration through the Ramshead exit is computed by the S/RVDL clearing analytical model in a manner similar to the mass flow rate as outlined in Item d. above.

The mass flow acceleration is used as input to the analysis of water clearing thrust loads on S/RVDL elbows. Figure 5.3.1-6 presents an example of the water mass acceleration transient.

5.3.1.2 Bases and Assumptions The bases and general assumptions that are applicable to the line clearing analytical model for Ramshead discharge devices are identical to those for the T-Quencher device. See Section 5.2.1.2 for details.

5.3.1.3 Load Definition Procedure

a. Outline of Procedure The following is a general outline of the procedure involved in evaluating the loads and parameters calculated by the line cl. car-ing analytical model.

Inputs for the line clearing model are prepared using Initial Conditions A through C for each S/RVDL under consideration. See Table 3.0-1. The initial water leg in the discharge line A

V 5.3.1-3 Revision 0

NEDO-21888 is an input to the line clearing model. The water leg must be O

determined by use of the reflood analytical model (see Section 5.3.3) prior to using the line clearing model for subsequent actuation cases.

Next, the line clearing model is used to obtain transient values for the following parameters or loads: 1) S/R7DL pressure,

2) thrust loads on S/RVDL pipe segments, 3) Ramchead discharge pressure, 4) water slug mass flow rate, and 5) water slug mass acceleration.

The values obtained for Ramshead discharge pressure and/or mass flow rates and acceleration are used as input to evaluate the:

1) Torus Shell Pressure Distribution (Section 3.3.2), 2) Air Bubble Drcg Loads (Section 5.3.5), 3) Water Jet Loads (Sec-l tion 5.3.4), and 4) Maximum Discharge Pipe and Ramshead Envire ental Temperature (Section 5.3.6).

b. Key Inputs O

The key inputs for Ramshead line clearing loads are identical to those for the T-Quencher device. See Section 5.2.1.3 for details.

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O . S/RV DISCHARGE PIPE SEGMENT 1r THE TRANSIENT THRUST LOAD IS THE NET LOAD CN THE PIPE SEGMENT TO ANO IS POSITIVE IN THE OIRECTION SHOWN DISCHARGE DEVICE Figure 5.3.1-3. Sense of Thrust Loading O

5.3.1-7 Revision 0

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1 l

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NEDO-21888 I !

j. i REFERENCES FOR SECTION 5.3.1 l

5.3.1-1.'" Mark I containment Program Analytical Model For Cc:puting Transient ;

j Pressures and Forces in the Safety / Relief Valve Discharge Line,"

a f

[ General Electric Company, Report No. NEDE-23749-P. !

i >

t

{ 5.3.1-2 E. A. Buzek, et al., " Final Report - In-Plant S/RV Discharge Load l l Test - Monticello Plant (Ramshead)," General Electric Company, S 4

Report No. NEDC-21581-P, August 1977. l 1- _

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-5.3.1-11/5.3.1-12. ' Revision 0

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5.3.2 Ramshead Torus Shell Pressures Loads on the torus shell resulting from air discharge from a Ramshead device are calculated using two separate analytical models. The bubble dynamics model calculates the air bubble pressure, radius and location as a function of time. The torus shell load model calculates the maximum and minimum pressures and the spatial pressure distribution on the torus shell.

The torus shell load definition procedure involves determining the vall pressure amplitude and frequency of an actual system, and applying these values to an experimentally measured wall pressure / time history that has been normalized. The procedure to conservatively predict loads on the torus shell involves the application of experimentally determined multipliers to the analytical prediction of the peak pressures from the torus shell pressure model. The adjusted magnitude of the shell load is then applied to an

(~) experimentally det. ermined torus shell pressure / time history. Additionally,

\' #

the frequency of the experimental pressure / time history is adjusted to account for line unique variations of the actual bubble frequency.

5.3.2.1 Lead Description The actuation of an S/RV results in the discharge of water from the S/RV discharge line followed by air and then steam. Subsequent to the water discharge, the air discharge from the S/RVDL forms two high pressure oscil-lating bubbles. The bubbles will rise as a result of buoyancy effects.

l The positive and negative dynamic pressures developed within these bubbles

)

are transmitted to the torus shell. Figure 5.3.2-1 presents an example of i the torus shell pressure transient resulting from an S/RV discharge through l a ramshead device. Figures 5.3.2-2 and 5.3.2-3 present examples of torus shell radial and longitudinal pressure distributions, respectively.

O)

\,.

5.3.2-1 Revision 0

NEDO-21888 5.3.2.2 Bases and Assumptions 9

The theoretical basis for the mathematical models which predict the torus shell pressure loads is described in Reference 5.3.2-1. Additicnally, experimental verification and empirica) djustment of the predicted torus shell pressure loads are used to justify the model and load definition procedure (See Reference 5.3.2-1).

When evaluating inputs for the bubble dynamics and torus shell pressure distribution analytical models the following general assumptions apply:

a. The bubble distance from the centerline of the Ramshead discharge device is 6.25 ft. This was empirically determined from the j measured pressure distributions (see Reference 5.3.2-1).

b. The Mark I torus suppression pool is modeled (for the torus shell pressure distribution model) as an equivalent rec-tangular cross section. This equivalent geometry is based on conservation of the pool water volume and pool depth. The results of this procedure are in good agreement with Monticello test data as reported in Reference 5.3.2-1. Reference 5.3.2-1 (Appendix D) contains the detailed basis and justification for the " equivalent rectangular geometry" assumption.

c. The pool water level is at the technical specification high water level. This assumption results in the maximum initial water leg in the S/RVDL which in turn results in the highest loads on the S/RVDL and discharge device.

d. The S/RVDL vacuum breaker does not leak, i.e., for first actuations, the initial water leg in the S/RVDL is determined by the pressure differential between the S/RVDL and wetwell air space. It is assumed equal to the drywell to wetwell 2P minus the vacuum breaker set point. By assuming the O

5.3.2-2 Revision 0

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k-)

s vacuum breaker does not leak, a lower value of S/RVDL to vetvell AP is calculated which results in a longer initial water leg in the dis-charge line. Since longer water legs result in higher loads, the assumption of a non leaking vacuum breaker is conservative,

e. The maximum S/RVDL reflood, which is predicted to occur at a time in the transient after the minimum predicted time between S/RV actuations (based on plant unique analyses), is used as the initial water leg for subsequent S/RV actuation evaluation. Optionally, if plant unique analyses are not performed, the maximum reflood occurring at any time during the reflood transient is to be used as the initial water leg for such evaluations.

5.3.2.3 Load Definition Procedure

a. Outline of Procedure The following is a general outline of the steps involved in obtaining fh -

(_) the pressure loads on the torus shell. Figure 5.3.2-4 presents the flowchart delineating this procedure.

First, the necessary inputs to the line clearing analytical model are prepared for Initial Conditions A through C (Table 3.0-1) for each S/RVDL under consideration (see Section 5.2.1.3).

Next, the line clearing analytical model is used (for each Initial Condition) to obtain the discharge pressure at the Ramshead after water clearing.

The necessary inputs are then prepared for each of the Initial Conditions (for each S/RVDL) for the bubble dynamics model.

General assumptions used in preparing the inputs are given in Section 5.3.2.3(b).

O 3.3.2-3 Revision 0

NEDO-21888 Next, inputs are prepared and the torus shell pressure distribution model is used to obtain the torus shell pressure transient ;c selected h

locations on the shell for a valve actuation from each S/RVDL dis-charge device for each Initial Conditien.

It is then necessary to compute a multiplier

to be applied to the

, torus shell pressures, which are output from the torus shell pres-sure distribution model, to adjust for condition dependent (first actuation, leaky valve or subsequent actuation) and location dependent (azimuthal attenuation) effects. See Reference 5.3.2-1 for the values of the multipliers.

The adjusted maximum positive and maximum negative shell pressures are then applied to the amplitude scale of the appropriate empirically derived pressure time history traces (see Figures 5.3.2-1, 5.3.2-5 and 5.3.2-6).

The next step is to adjust the time scale of the appropriate condi-tion depend'ent pre.ssure history to account for the actual bubble frequency due to S/RVDL differences. This is accomplished by determining a time scaling parameter (ts)*

The time scaling parameter (Ts) to be applied to the appropriate empirical pressure / time history (see Figures 5.3.2-1, 5.3.2-3, and 5.3.2-6) to stretch or compress the time scale is computed from the following equations.

f (time scaling parameter) 1 Ts " - f-1

The basis and justification for the multipliers sre contained in Reference 5.3.2-1.

O 5.3.2-4 Revision 0

NEDO-21888 FN where O

f, . 6.9 Hz (from equation below using Monticello test data).

The frequency predtetion for the given S/RVDL is:

3kgc P,(144) f = r.n gp

, = L It should be noted that the air bubble frequency prediction is derived from the Rayleigh equation assuming adiabatic compression and expansion. See Reference 5.3.2-1 for a description of this equation.

The calculated frequency of the torus shell pressure transient is assumed accurate to :15 percent for first actuation, leaky valve, and subsequent actuation cases. This uncertainty is based on consideration of data scatter observed in full scale testing (Reference 5.3.2-1), and the accuracy with which the above frequency

) equation has~been shown to predict torus shell pressure loading frequencies over the range of operating conditions. Due to the above uncertainty, the structural evaluation is performed uting the calculated frequency within the specified uncertainty range, which is expected to result in the maximum structural response.

Each load case is then evaluated under the appropriate Initial Condition given in Table 3.0-1. For multiple valve actuations, the peak positive and negative shell pressures predicted at a given location for eac! h6 discharge device are to be combined by calculating the square root of the sum of the squares (SRSS) of the peak pressure loads predicted to occur at that location for each individual S/RV actuation (see Section 5.2.2.2).

Following evaluation of the torus shell pressure distributions for the Load Case / Initial Conditions shown in Table 3.0-1, the pres-sure loadings are to be combined with other torus shell loads rm occurring during the same time period per the load combination bar charts presented in Section 3.0.

( s) 5.3.2-5 Revision 0 i

b. Key Inputs There are several parameters that must be evaluated in order to apply the bubble dynamics analytical model. These include the thermo-dynamic state of the gas in the S/RVDL and geometric considerations.

Such parameters include:

Pressure in the discharge device during air discharge e Initial air mass in the S/RVDL e Initial gas temperature and pressure in the S/RVDL e Initial gas volume in the S/P.VDL e Suppression pool temperature and pressure at the discharge device exit e Suppression pool geometry e Configuration of Ramsheads in the suppression pool Similarly, the key inputs that must be evaluated to apply the torus shell pressure distribution model are as follows:

O e Torus shell geometry e Configuration of Ramsheads in the suppression pool e Maximum bubble radius e Minimum bubble radius e Maximum bubble pressure e Minimum bubble pressure e Bubble oscillation frequency e Bubble azimuthal location 1

9 5.3.2-6 Revision 0 l

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Ramshead [. Ramshead Located on Torus Centerline.

5.3.2-8 Revision 0

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5 DISCHARGE MAXIMUM BUB 6LE Pf1 ESSURE CHAHACTEfliSTICS MAX POS AND NEG m

TORUS SHE t t

PIPE CLEARING BUBBLE -

PHESSuftES MODEL DYN AMICS MODEL P

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POSITIVE PHESSUHES AHE NORMALIZED TO THE MAXIMUM POSITIVE PHESSultE AND NEGATIVE PHESSURES AHE NORMALIZED TO liiE MAXIMUM NEGATIVE PRESSUHE l 06 -

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REFERENCES FOR SECTION 5.3.2 5.3.2-1 " Safety Relief Valve System Analytical Models for tlse with Ramshead Discharge Devices," General Electric Company, Report No. NEDE-23803-P, July 1978.

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5.3.3 S/R'IDL Reficed Transient (Ramshead)

The reflood analytical model is used to determine the transient water rise in i.a S/RVDL following S/RV closure. Outputs from the reflood model are used to provide inputs to the models which calculate loads during a sub-sequent S/RV actuation. For details of the procedure used to calculate reflood parameters refer to Section 5.2.3.

5.3.3.1 Bases The reflood procedura and analytical model for the Ramshead discharge device are identical to those used for the T-Quencher discharge device.

The only differences are model input parameters dependent on the gecretry of the discharge device (e.g., loss coefficients, volumes, inertial lengths, and average areas). The analytical model is described in detail in Reference 5.3.3-1.

This reference also shows that the analytical model conservatively predicts

() test data which is documented in Reference 5.3.3-2.

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NEDO-21888 REFERENCES FOR SECTION 5.3.3 5.3.3-1 A. J. Wheeler, et al., " Analytical Model for Computing 'Jacer Rise in a Safety / Relief Valve Discharge Line Following Valve Closure,"

General Electric Company. Report No. NEDE 23898, October 1978.

-5.3.3-2 E. A. Buzak, et al., "In-Plant S/RV Discharge Load Test - Monticello Plant (Ramshead),". General Electric Company, Report No. NEDC-21581-P, August 1977.

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5.3.3-3/5.3.3-4' Revision 0

NEDO-21888 5.3.4 Ramshead Water Jet Loads on Submerged Structures

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5.3.4.1 Load Description An S/RV actuation causes the discharge of a water column into the suppression pool. As the water column leaves the Ramshead discharge device and enters the suppression pool, it forms a jet which may load structures inside the torus which are fully or partially intercepted by the discharge. The calculation of these loads is described in this section.

l Jet loading lasts from the time that the jet reaches the structure until the time when the last particle of the water slug passes the structure. Fig-ure 5.3.4-1 presents a sample predition of this load.

5.3.4.2 Bases for !kthodology The calculation procedure described to obtain Ramshead jet loads is based on the analytical model described in Reference 5.3.4-1. S/RV line unique dis-l

() . charge jet velocities and accelerations are obtained from the S/RVDL clear-ing model an described in Section 5.3.1. The major assumptions included in i the methodology are as follows:

i

a. When a structure is engulfed in the jet path, the force on the r

structure can be calculated using a standard velocity squared drag equation. If a structure partially or fully intercepts the jet, a momentum balance is used to define a proportionality factor to be used in place of the standard drag coefficient.

b. The jet is completely dissipated when the last fluid particle leaving the Ramshead reaches the jet. front; the dissipated jet provides no more drag.

j 5.3.4.3 Load Definition Procedure

-The procedure used to calculate forces on submerged structures in the Ramshead

- jet path is presented below.

5.3.4-1 '

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NEDO-21888 First, from the S/RVDL clearing model described in Section 5.3.1 the time dependent discharge velocity and acceleration are obtained. The parameters g

involved are:

! = An identifier equal to the time at which a jet particle is discharged from the Ramshead, see VD (T) = Jet discharge velocity at time r, ft/sec dVD (*}

AD (I} " dr

= Jet discharge acceleration at time T, ft/sec 2 Next the maximum jet penetration distance is determined by D IImax) , gg max where T denotes the . time at which the water jet clears the Ramshead. If 1 < Ls , there is no jet load on the structure. If f. 3 L,, there will be jet loads on the structure. La = distance (x) from the Ramshead discharge to the structure of concern.

From the Ramshead clearing data, a plot of the jet front position versus time is constructed. This is done by plotting the position (denoted as x ) versus time (t being the elapsed time following the start of the S/RV discharge history of a number of fluid particles which exit the Ramshead during the clear-ing transient). This time-space plot (t,x) is based on the following equation:

x = V (T) (t - t), ft D

Whenever a line of constant t overtakes an earlier line, the earlier line ceases to extend, and the overtaking line continues. The time-space plot enables one to enter a location x) read the corresponding value of T. e nd then obtain associated jet velccity f rom the Ramshead clearing.

O 5.3.4-2 Revision 0

('T For a structure at position x), values of r and t can be obtained from the time space plot. Then the equation below and the Ramshead clearing data is employed to determ , the correspondine jet area. The jet area is given by:

A A(x ,t) =

D =

D D(* '

y 3 , ft' 7

1y AD (I) (E~I) l~f V 'k(T)*j where A is the discharge area of the S/RV line.

D The area of the jet front is approximated by the area the jet possesses one initial discharge diameter (d = (4AD /w)1/2) rem ved fr m the front. An additional computation of the jet radius determines which structures are overlapped by the jet, and to what degree overlap occurs.

The standard drag force on the submerged structure is based on the normal component of velocity intercepting the structure, the projected area inter-cepted by the normal component of velocity, and the jet area profile.

If a st- cture is fully submerged inside the jet boundary at time t, a velocity squared drag calculation is used.

Fy (t) =

CD ^P 2g

, lbf c

Where pA represents the projected area intercepted by the jet.

If a structure fully or partially intercepts the jet (no jet flow around structure) the following equation applies:

I Fy (t) = Ky Ap V(r) , lbf c

Where yK = 2 for momentum stoppage, yor K = 4 if the structure turns the jet back on itself.

A 5.3.4-3/5.3.4-4 Revision 0

NED0-21888 20 000 O

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0 # I l 0.18 0.19 0.20 0.21 TIME AFTER JET CONTACTS STRUCTURE (sec)

Figure 5.3.4-1. Sample Prediction of a Typical Structure Loading History O

1 5.3.4-5/5.3.4-6 Revision 0

f 1

NEDO-21888 !

1 1

REFERENCES FOR SECTION 5.3.4 5.3.4-1 F. J. Moody, " Analytical Model for Liquid Jet Properties for Predicting Torus on Rigid Submerged Structures." General Electric Company, Report No. NEDE-21472, September 1977.

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NEDO-21888 f^.

V 5.3.5 Ramshead Bubble-Induced Drag Load on Submerged Structures Oscillating bubbles resulting from S/RV actuation create an unsteady three-dimensional flow field and therefore induce acceleratim. and standard drag forces on the submerg-td structures in the suppression pool.

This section addresses the load definition procedures for loads on sab-merged structures due to Ramshead S/RV bubbles. The following sub?ections discuss the methodology to define the loads; the bases and armumptions on which the analytical models are developed and verified; and the selection of key parameters from typical plants for which the design load on sub-merged structures can be obtained.

5.3.5.1 Bases and Assumptions The bases and assumptions of the analytical model are derived from the work in Reference 5.3.5-1. The major assumptions are summarized as follows:

a. The total drag on submerged structures is the sum of the acceleration and standard drags.

b. The flow fields caused by Ramshead. S/RV bubbles are described by fluid sources. The presence of boundaries is accounted for by using the method of images.

c. Source strengths are determined from the Ramshead bubble dynamics model (Reference 5.3.5-2).

5.3.5.2 Load Definition Detailed procedures to calculate drag forces on submerged structures are described in Reference 5.3.5-1 which provides the methodology of utilizing the unsteady flow field generated by the Ramshead S/RV bubbles in the suppression pool to obtain drag loads on submerged structures such as pipes, beams, etc.

5.3.5-1 Revision 0 i

NEDO-31888 The load definition first requires the establishment of the unsteady flow field within the torus by simulating the S/RV bubbles by poirt sources with strengths determined from the Ramshead S/RV bubble dynamics model (Reference 5.3.5-2). Then, drag loads on submerged structures are calculated based on the previously established unsteady flow field.

a. Flow Field Establishment An oscillating Ramshead S/RV bubble is considered as a fluid source with a source strength determined from the Ramshead S/RV bubble dynamics model (Reference 5.3.5- G The Ramshead S/RV bubbles in the Mark I torus are simulated by multiple sources in a finite pool. By using potential flow theory and the method of images to account for the effects of solid walls and the free sur-face, the flew field within the torus is established.

b. Drag Loads Evaluation The drag force on a submerged structure consists of two components:

the standard drag and the acceleration drag. The standard drag load is computed on a locally uniform flow field evaluated at the local velocity. The acceleration drag load is evaluated based on an inviscid, uniform but unsteady (accelerating) flov field. The sum of two drag forces gives the total drag load on a submerged structure.

5.3.5.3 Selection of Key Parameters for Load Evaluations This section outlines those key parameters which influence the magnitude and characteristics of the Ramshead S/RV bubble drag loads on submerged struc-tures.

The procedure used to specify the inputs for the submerged structure i

drag load model is also discussed. The plant specific suppression pool geometry should first be identified as follows:

a. Torus shell dimensions

b. Torus water level O

5.3.5-2 Revision 0

~

k<,,s) c. Relative locations of the S/RV discharge lines in the suppression pool.

Then, the initial conditions of each S/RV discharge lines in the suppression as

a. Initial pipe pressure

b. Initial volume of air-steam mixture.

Next, the S/R'.*DL clearing model and the Ramshead S/RV bubble dynamics model are applied ;.s discussed in Sections 5.3.1 and 5.3.2 to determine the following inputs:

a. Discharge pressure

b. Clearing time

() c. Ramshead bubb'le pressure and' bubble radius history.

Af ter dividing the structure into appropriate sections for more precise calculation, information on the structure section locations, orientations, acceleration drag volumes, projected area and the standard drag coefficients should also be obtained.

With the above parameteiJ identified, the dimensions of the pool model, the coordinates of the initial locations of the Ramshead S/RV bubbles and the structure sections can be calculated.

With the input parameters identified and calculated, the submerged structure drag load model is used to evaluate the resultant transient drag forces on each atructure sections being analyzed. Typical results are shown in Fig-ures 5.3.5-1 and 5.3.5-2. These figures show the Ramshead bubble drag forces on a downcomer and a vent header support column in the radial (X) and tanhential (Z) directions as a function of time.

s

{V 5.3.5-3/5.3.5-4 Revision 0

50.0 d COWNCOMER O VENT HEADER SUPPORT COLUMN A

37.5 -

25 0 -

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Figure 5.3.5-1. Sample Pre.ficted Time History of Total X-Forces on Downectaer and Vent Header Support Column O

5.3.5-5 Revision 0 w , _ , - - -

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NED0-21808 O

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Figure 5.3.5-2. Sample Predicted Time History of Total Z-Forces on Downcomer and Vent Header Support Ce1.umn 5.3.5-6 Revision G

NEDO-21888 REFERENCES FOR SECTION 5.3.5 5.3.5-1 F. J. Moody, et al., " Analytical Model for Estimating Drag Forces on Rigid Submerged Structures Caused by LOCA and Safety Relief Valve Ramshead Air Discharges," General Electric Company, Report No.

NEDO-21471, September 1977.

5.3.5-2 " Mark I' Containment Program, Safety-Relief Valve System Analy :al '

Models for Use with Ramshead Discharge Devices", General Elec ric Company, Report No. NEDE-23803-P, December 1978.

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-5.3.6 Maximum S/RVDL'and Discharge Device k'all Temperature

' The procedure to determine the thermal loads ass'ociated with the S 'RV dis- l

2. charge line and the Ramshead device is identical to that.for the T-Quencher-device. See Section 5.2.7 for details of the procedure.

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The procedure _to. determine the fatigue usage factor for lines terminating

{- in'a Ramshead discharge device-is identical to that for a T-Quencher discharge

~ device . See Section 5.2.8 for-details of the procedure.

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SECTION 6 OTHER CONSIDERATIONS J

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NEDO-21888 6.0 OTHER CONSIDERATIONS

~

Section 6 discusses LOCA and S/RV loads not addressed in Sections 4 and 5.

t These loads are of secondary importance because they are negligible when

, ~ compared to other-loads defined for the structures affected. Each potential

-loading condition is described and justificatio'n is given for its classifica-4 tion as negligible.

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NEDO-21888 d 6.1 SEIS}iIC SLOSH LOADS In the event of an earthquake, the Mark I torus will be subjected to both horizontal and vertical oscillatory motions which will produce waves on the surface of the suppression pool (seismic slosh) . These waves will generate pressure loads on the torus shell. Unsteady motions in the pool water will also generate drag loads on submerged structures.

Scale model tests of the Mark I torus were conducted to determine the seismically induced slosh amplitude and slosh pressures (Reference 6.1-1). The tests were performed over the range of Mark I Final Safety Analysis Report (FSAR) seismic criteria. Calculated slosh amplitudes for the full scale Mark I plants are small, as shewn in Table 6.1-1. These results are conservative since the seismic respense spectrum used in the seismic slosh tests is an envelope of the individual responses for all the Mark I plants. The magnitude of the maximum local slosh pressures on the full scale torus shell was found to be less than 0.8 psi. This pressure is insignificant when compared to the other torus wall loads occurring during a postulated LOCA.

w.) '

The drag loads on submerged structures due to the vertical pool motion can be conservatively evaluated by using the following equation:

P= 2 0.0005n cos(1.684c)lcos(1.684c)l

- 0.001 nDsin (1.684t) sin 0 where P = net drag load on structure (psi). Applied normal to the structures' longitudinal axis.

n = maximum slosh amplitude, in.

I t = time, see i

%)

6.1-1 Revision 0

NED0-21888 D = diameter of cylinder circumscribing r.he structure, in.

r = distance from center of torus to object, ft R = Initial water level plus the maximum slosh amplitude, ft 0 = angle between the vector normal to the structures' longitudinal axis and a line drawn from the center of the torus to the point of application (see Figure 6.1-1), rad. -

The above equation was devel, ped assuming sinusoidal motion and used the maximum drag coefficients for oscillatory flow from Reference 6.1.2. It is assumed to apply to any structure below the height defined by the addition of the torus high water level and the maximum slosh amplitude. Large struc-tures should be divided into smaller sections for force calculationa.

Figure 6.1-1 illustrates the application of the above equation.

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O 6.1-2 Revision 0

NEDO-21888 O Table 6.1-1 MARK I CONTAINMENTS SEISMIC SLOSH RESULTS Maximum Sloth Amplitudes at Torus Wall Plant (inches)

Browns Ferry 1, 2, 3 21.0 Brunswick 1, 2 15.5 Cooper 19.8 Dresden 2, 3 21.2 Duane Arnold 11.2 Fermi 2 16.7 Fitzpatrick 15.5 Hatch 1, 2 14.9 Hope Czeek 1, 2 20.8 Millstone 16.4 Monticello 11.1 Nine Pdle Poin,t 1 O Oyster Creek 10.2 20.7 Peach Bottom 2, 3 12.6 Pilgrim 14.3 Quad Cities 1, 2 25.0 Vermont Yankee 12.9 i

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INITIAL POOL SURFACE y TCRUS CENTER

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v STRUCTURE LOAD MAGNITUDE = P -

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g Figure 6.1-1. Application of Seismic Slosh Drag Lead i O 6.1-4 Revision 0

NEDO-21888 REFERENCES FOR SECTION 6.1 6.1-1 S. Arain, Seismic Slosh Evaluation, General Electric Company, Report No. NEDC-23702-P, March 1978.

6.1-2 T. Sarpkaya, " Forces on Cylinder Near a Plane Boundary in a Sinusoidally Oscillating Fluid," Jour".cl of Fluids Engineering, September 1976.

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l 6.1- 5 Revision 0 1

k' 6.2 POST POOL SWELL WAVES Following a LOCA, random pool surface waves are present in the torus as a result of two sequential wave formation mechanisms:

a. The fallback of the bulk pool swell after the DBA and the subsequent pool motion which is generated and dies out as the vent flow reaches a quasi-steady stata condition

b. The pool surface disturbance resulting from steam blowdown through the vent system during the subsequent condensation oscillation and chugging.

The basis for the evaluation of the wave behavior during the first regime is observed wave loads from the QSTF, Reference 6.2-1. The loading frcm these waves during and immediately after fallback is included in the pressure data directly incorporated into the torus shell pressure histories (Section 4.3.2).

Although the QSTF results are not scaled precisely for system behavior following bubble break through, they do give the correct order of magnitude for wave behavior. By review of the pool swell vertical loading generated during this time period, it is concluded that the wave loading is insignificant when compared with other vertical loads.

Loads resulting from waves generated during the second regime have been evaluated based on FSTF data (Reference 6.2-2). As FSTF torus shell pressure transducers measured these loads during condensation oscillation and chugging, they are inherently included in the load definition for these phenomena.

I O 6.2-1/6.2-2 Revision 0 l

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REFERENCES FOR SECTION 6.2

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1 6.2-1. J.-M.:Humphrey, Mark I Containment' Program, "l/4-Scale Two-Dimensional Plant Unique Pool Swell Test Repo r t" , General Electric 1 Company, Report No. NEDE-21944-P, to be published, j i

.6.2-2 .V. E. Torbeck,'" Mark I Containment Program Full Scale Test Program", ;

{' General _ Electric Company, Report No. NEDE-24539-P, to be published. j u

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NEDO-21888 O 6.3 ASY10!ETRIC VENT PEPJORMANCE During a LOCA, geometric effects in the >! ark I vent system may cause differences in pool swell behavior around the circumference of the torus. As discussed in this section, these variations are already included in the loads presented in this report.

The circumferential variations in pool swell loads were studied in the 1/12. scale three-dimensional tests performed at EPRI (Reference 6.3-1). Because all plants have similar downcomer spacing, one model was used to define the variations in longitudinal pool swell behavior for all plants. In that model, one specific plant vent system and torus were accurately modeled to include the effects of nonuniform downcomer spacing and the flow distribution between dowcomers. It is the difference in the fL/D between downcomers which causes the variation in ficw rate, and the test facility was designed to accurately simulate the fL/D values of the prototype vent system.

Near a downcomer pair, the pool behavior is determined by the expanding bubbles.

The pool water is pushed upward more rapidly directly over the bubbles, locally generating higher peol swell velocities and pool swell heights. At locations where the downcomer spacing is close, the resulting pool swell velocity and height are higher than at those locations where the spacing is further apart.

The EPRI test results were used as a basis for including the longitudinal variation of pool swell behavier. The following pool swell loads have been determined including that behavior:

a. Torus shell pressure histories (Section 4.3.2)

b. Vent system impact and drag loads (Section 4.3.3)

c. Impact and drag loads on structures above the pool (Section 4.3.4)

d. Froth impingement loads (Section 4.3.5)

e. Pool fallback loads (Section 4.3.6),

a i

i 6.3-1/6.3-2 Revision 0

NEDO-21808 O

REFERENCES FOR SECTION 6.3 6.3-1 The Electric Power Research Institute, "Three-Dimensional Pool Swell Modeling of a :tark I Suppression System," EPRI NP-906, October 1978.

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NEDO-21888 6.4 DOWNCOMER AIR CLEARING LATERAL LOAD l l

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I,J 6.5 SONIC WAVE With the postulated instantaneous rupture of a steam or recirculation line, a pressure wave traveling at sonic velocity would expand from the break location into the drywell atmosphere and through the vent system. The wave amplitude, at the break location and at the time of the break initiation, would be substantially below the reactor operating pressure and would attenuate rapidly as it expanded out from the break location into the large drywell volume. The wave amplitude af ter entering the vent system would be nearly uniform, but greatly attenuated from its initial value at the break location.

The insignificance of the sonic wave was determined from the tests performed on the FSTF (Reference 6.4-2) and QSTF (Reference 6.5-1).

FSTF data traces show some evidence of a shock wave at break ir.itiation but this is due in part to the arrangement of the discharge pipe configuration for the simulated break flow. The sonic wave produced by the rupture disc was not proto-

~~g typical since the rupture dise led into a long, empty line in which a strong pres-

\-) sure wave would be expected. In addition to this, the break flow is discharged to one location in the cylindrical FSTF drywell which has no other structures contained in it. In the Mark I plants the large drywell volume would greatly attenuate any shock wave, and scattering of the shock wave due to the equipment and piping located inside the drywell would further reduce the effects of any sonic wave.

The FSTF had a rupture disc arrangement upstream of the drywell which simulates the instantaneous rupture of the steam or recirculation pipe. In the FSTF dis-charge pipe section, the rupture disc was located downstream of the orifice which simulates the break area. It should be noted that in the QSTF discharge pipe arrangement, the rupture disc was located just upstream of the orifice used to simulate the break area. Further, in FSTF the critical flow area of the rupture ;

disc which generates the shock wave was much larger (by a factor of 3) than the r

6.5-1 Revision 0 l

NEDO-21888 throat section of the orifice (simulated break area). Therefore, the initial shock wave generated in FSTF is expected, but is not prototypical for DBA conditions.

Additionally, the FSTF arrangement as discussed earlier, did not allow for the prototypical attenuation of this sonic wave as would occur in the Mark I plants.

The QSTF data confirms that the sonic wave generated in FSTF was not prototypical.

Examination of the QSTF data shows no evidence of a sonic wave in the drywell, vent system, or wetwell. Since in QSTF the rupture disc was located upstream of the flow orifice which simulates the break area, the strength of the sonic wave is better simulated than in FSTF. Note also that QSTF also had an empty discharge line downstream of the orifice and a drywell with no structures to scatter the wave.

Considering the above discussion, it is concluded that the loading on the vent system and torus due to sonic waves is negligible.

6.5-2 Revision 0 O

4-NEDO-21888

'l 4

l REFERENCES FOR SECTION 6.5 ;

6.5-1 J.~M.'Humphrey, " Mark I Containment Program 1/4 Scale Two-Dimensional Plant Unique Pool Swell Test Report," General Electric' Company, '

Report No. NEDE-21944-P, to be published.

1

.6.5-2 J. E. Torbeck, " Mark I Containment Program Full Scale Test Program,"

i General Electric Company, Report No. NEDE-24539-P, to be published.

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(~'S 6.6 COMPRESSIVE WAVE

%)'

The rapid bulk pressurization of the drywell immediately following a postulated DBA prior to vent clearing would theoretically result in a weak compressive wave being generated in the downcomer water leg that would travel across the pool to the torus shell.

The insignificance of the compressive wave was determined from tests performed on both the QSTF and the FSTF, see References 6.6-1 and 6.6-2, respectively.

If the wave phenomena were significant, its presence would have been observed in the QSTF, because the drywell pressure response and wetwell pressure response were accurately acaled. Additionally, the pressure transducers used in the QSTF were capable of measuring response at frequencies up to 200 Hz, which would have been sufficient to detect a compressive wave. Examination of wetwell submerged pressure test data indicatee no evidence of a compressive wave. The same observation has been made in data obtained from the FSTF which also has instrumentation capable of detecting the presence of such a wave.

I)

\j The energy of the presumed compressive wave is dissipated in accelerating the water slug in the downcomers into the pool. It is also attenuated as the wave travels from the downcomer exit to the torus wall. Therefore, the resulting pressure on the torus wall generates negligible loads as verified by the QSTF and FSTF tests.

Even though the presence of the compressive wava has not been specifically observed, its effect has already been included in the load definition for both the torus net vertical loads (Section 4.3.1) and the torus shall pressure histories (Section 4.3.2) because the QSTF was built and operated in accordance with the pool swell scaling laws. Thus, the potential compressive wave behavior is accurately included in the test results.

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REFERENCES FOR SECTION 6.6

6. 6-l' -J. M. Humphrey, " Mark I Containment Program, 1/4 Scale Two- ,

Dimensional Plant Unique Pool Swell Test Report," General Electric

Company, Report No..NEDE-21944-P, to be, published.

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l 6.6-2 J. E. Torbeck, " Mark I Containeent Program Full ScaleTest Program," '

General Electric Company, Report Nc; NEDE-24539-P, to be published.

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NEDO-21888 O 6.7 SAFETY / RELIEF VALVE STEAM CONDENSATION LOADS 6.7.1 T-Ouencher Discharge Following opening of an S/RV and the expulsion of the air and water initially in the line, a steady flow of steam enters the suppression pool through the T-Quencher exit holes. As steam is condensed within the pool, high-frequency, low-magnitude pressure loads are produced on the torus shell. I Steady state steam discharges through T-Quencher devices were performed during ,

the Monticello T-Quencher Test, Refer e 6.7.1-1. These tests were performed at two discrete mass fluxes throug) t luencher holes (150 lbm/sec-ft2 and 2

22 lbm/sec-ft ). During these tests t zimum steam condensation loading on the torus shell over a local pool temperature range of M to ll5*F was 20.6 psid with frequencies ranging from 75 to 230 Hz. The peak dynamic extreme fiber principle stress determined from strain data measured on the torus shell during steady steam discharge was negligible (less than 0.6 ksi).

The results of additional testing of the steam condensation performance of a cross quencher device are documented in Reference 6.7.1-2. During these tests the vertical and horizontal spacing, as well as the diameter of the exit holes, were identical to that on the T-Quencher. These parameters are of primary importance in the condensation performance of a quencher device.

Sonic steam discharges through cross quencher arms resulted in no discernible changes in steam condensation pressure magnitude over a local pool temperature ranga of 90 to 212*F.

The low magnitude and high frequencies (higher than those producing the major contribution to torus shell and support response) of these loads, indicate j that they are of relatively minor importance in the structural assessment

.of the Mark I torus and support sytem.

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' REFERENCES FOR SECTION 6.7.1

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6.7.1-1 R..A. Asai, et al., " Mark I Containment Program Final. Report - i Monticello T-Quencher Test," General Electric Company, Report No. NEDE-21864, July 1978.

6.7.1-2' T. Y. Fukushima, et al., " Test Eesults Employed by GE for EWR Containment and-Vertical Vent Loads," General Electric Company, Report No. NEDE-21078, 0ctober 1975.

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6.7.2 Ramshead Dischar2e Steady state steam discharges through Ramshead devices were performed during the Menticello Ramshead Test (Ref. 6.7.2-1). These tests were performed at a mass flux of 200 lbm/see-ft through the Ramshead exit areas. During these tests the maximum steam condensation loading on the torus shell over a local pool temperature range of 76 to 93*F was i5.0 paid. The frequency of these pressure oscillations ranged between 80 and 125 Hz. The peak dynamic extreme fiber principle stress determined from strain data measured on the torus shell during steady steam discharge was small (less than 3.0 ksi).

Additional data (Reference 6.7.2-2), provides the basis for the establishment of a steam condensation stability limit of 170*F (local pool temperature) for Ra=sheads when the mass flux exceeds 40 lbm/sec-f t . This data indicates that smooth condensation of steam from a Ramshead discharge can be expected for local pool temperatures below 160*2, f) Operation of Mark I plants within the technical specification limits on pool v

temperature, and prudent operator action following any postulated transients or accidents, will ensure that local temperatures above 160*F concurrent with mass fluxes above 40 lbm/sec-f t will not be reached. Therefore, steam condensation will remain smooth.

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