ML20249C603

Non-proprietary Rev 1 to WCAP-14271, Low-Pressure Integral Sys Test Facility Scaling Rept

ML20249C603
Person / Time
Site:	05200003
Issue date:	08/31/1997
From:	WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML20036E486	List: ML20249C600 ML20249C603 NSD-NRC-98-5718, Forwards non-proprietary Rev 1 to WCAP-14271, Low Pressure Integral Sys Test Facility Scaling Rept & Proprietary Rev 1 to WCAP-14270, AP600 Low-Pressure Integral Sys Test Facility Scaling Rept. Proprietary Encl Withheld
References
WCAP-14271, WCAP-14271-R01, WCAP-14271-R1, NUDOCS 9806300388
Download: ML20249C603 (500)

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WESTINGHOUSE NoN. PROPRIETARY CLASS 3 FACILITY SCALING REPORT 1

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l WCAP-14271 l Rev.I  ;

LOW-PRESSURE INTEGRAL SYSTEMS TEST FACILITY SCALING REPORT l

l l AUGUST 1997

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l-l l Westinghouse Electric Corporation I

l Energy System Business Unit l P.O. Box 355 l Pittsburgh, PA 15230-0355 l

l C 1997 Westinghouse Electric Corpcra: ion (A) l All Rights Reserved 3631w.l.non:Ib-080497

FACILITY SCALING REPORT l.

!- TABLE OF CONTENTS Section .Till_e Page EXECUTIVE

SUMMARY

1 ABSTRACT 6 ACKNOWLEDGMENTS 7

1.0 INTRODUCTION

1-1 l 1.1 Scaling Objectivet - 1-1 1.2 General Scaling Methodology 1-2 1.3 Evaluation of Scaling Analysis Methods 1-3 j 1.4 ' Rationale for Scaling Choices 1-12 1.5 References 1-17 L

l 2.0 EXPERIMENTAL OBJECTIVES AND GENERAL SCALING METHODOLOGY 2-1 2.1 Test Facility General Modes of Operation 2-2 2.2 Fundamental Scaling Requirements 2-3 i

.. 2.3 AP600 System Decomposition and Hierarchy 2-4 2.4' Initial Conditions for Long-Term Cooling 2 ! ~3.0 , . PHENOMENA IDENTIFICATION AND RANKING 3-1 3.1 . AP600 Design and Emergency Core Cooling 3-1

.J l < 3.2 Plausible Phenomena Identification Ranking Table 3-5 1 3.3 References 3-10 4.0 CLOSED LOOP . NATURAL CIRCULATION SCALING 4  ;

. 4.1 Single-Phase Natural Circulation Scaling Analysis 4-1

l. 4.2 Two-Phase Natural Circulation Scaling Analysis 4-13 4.3 Primary Loop Design Specifications 4-51 4.4 Evaluation of the Core Processes Specific Frequencies, Characteristic Time i

Ratios, and Scaling Distortions 4-51 L - 4.5 - Conclusions 4-54 4.6 References 4-55 l- '

5.0 - .OPEN SYSTEM DEPRESSURIZATION SCALING ANALYSIS 5-1 5.1 Description of the Depressurization Process 5-2 '

5.2 . Governing Equations for the Two-Phase Fluid System Depressurization 5-2 5.3 Top-Down Subsystem Level Analysis for the Depressurization of a 5-7 O- Break Flow Rate Dominated System  ;

CJ 5.4 Top-Down System Level Depressurization Scaling Analysis 5-16

[7 368Iw.l.non.1b-060897 iii l

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FACILITY SCALING REPORT TABLE OF CONTENTS (Cont.)

Section Title _P_ age 5.5 Scaling Synergistic Phenomena 5-24 5.6 Bottom-Up Scaling Depressurization Scaling Analysis 5-26 5.7 Evaluation of Depressurization Specific Frequencies, Characteristic Time Ratios, and Scaling Distortions 5-55 5.8 Conclusions 5-56 5.9 References 5-57 6.0 CORE MAKEUP TANK SCALING ANALYSIS 6-1 6.1 CMT Phenomena 6-2 6.2 Scaling Analysis for CMT Recirculation 6-4 6.3 Scaling Analysis for CMT Draining 6-13 6.4 Conclusions 6-30 6.5 References '6-31 7.0 VENTING, DRAINING, AND INJECTION SCALING ANALYSIS 7-1 7.1 Depressurization Scaling Requirements for Venting, Draining, 7-2 and Injection Processes 7.2 General Scaling Analysis for Tank Draining Processes 7-6 7.3 Accumulator Scaling Analysis 7-7 7.4 In-Containment Refueling Water Storage Tank Scaling Analysis 7-13 7.5 Safety Injection Line Scaling Analysis 7-17 7.6 Balance and Vent Line Scaling Analysis 7-26 7.7 Bottom-Up Scaling Analysis for Upper Core Support Plate Draining 7-36 7.8 Reactor Vessel Downcomer Scaling Analysis 7-39 7.9 Conclusions 7-48 7.10 References 7-49 8.0 LCS RECIRCULATION COOLING SCALING ANALYSIS 8-1 8.1 Top-Down Scaling for Analysis for LCS Recirculation Cooling 8-2 8.2 Bottom-Up Scaling Analysis for LCS Recirculation Cooling 8-3 8.3 Lower Containment Sump Scaling Analysis 8-4 8.4 Evaluation of Core Process Specific Frequencies, Characteristic Time Ratios, and Scaling Distortions 8-6 8.5 Conclusions 8-7 8.6 References 8-7 l

9.0 SCALING ASSESSMENT 9-1 9.1 References 9-3 L

i 3081W-i.ftutt A b0605U j9/

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FACILITY SCALING REPORT

(]

v TABLE OF CONTENTS (Cont.)

Section Title Pm

! 10.0

SUMMARY

OF RESULTS AND CRITICAL PHYSICAL ATTRIBUTES 10-1 10.1 Dominant Processes 10-1 l 10.2 Scaling Distortions 10-3 10.3 Critical Attributes l 10-4 l

10.4 Conclusions 10-4 APPENDIX A CONSTITUENT LEVEL CONTROL VOLUME BALANCE EQUATIONS A-1 FOR A TWO-PHASE FLUID l

l APPENDIX B STEADY-STATE ANALYSIS OF TWO-PHASE FLUID NATURAL i CIRCULATION B-1 APPENDIX C TRANSFORMATION LAWS FOR SCALING POLYNOMIALS C-1

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3681w-l.non:Ib-060897 Y

FACILITY SCALING REPORT LIST OF TABLES Table Title _P_ age Table 1-1 Rationale for Scaling Choices 1-19 l

Table 3-1 Plausible Phenomena Identification Ranking Table for AF600 Large

}

Break LOCA 3-11 t

Table 3-2 Plausible Phenomena Identification Ranking Table (PPIRT) for AP600 l Small Break LOCA 3-16 Table 3-3 Description of AP600 SBLOCA and Long-Term Cooling Phenomena Ranked (H) or (P) 3-19 Table 4-1 Steady-State Loop Balance Equations for Single-Phase Natural Circulation Flow 4-57 Table 4-2 Single-Phase Constituent Level Scaling Analysis: Control Volume l Balance Equations for the Core 4-58 I Table 4-3 Single-Phase Constituent Level Scaling Analysis: Non-Dimensionalized Balance Equations for the Core 4-59 Table 4-4 Single-Phase Constituent Level Scaling Analysis: Residence Times and Characteristic Time Ratios for the Core 4-60 Table 4-5 Single-Phase Constituent Level Scaling Analysis: Process Specific Frequencies for the Core 4-61 Table 4-6 Steady-State, Single-Phase Natural Circulation Loop Scaling Ratios 4-62 Table 4-7 Steady-State, Single-Phase Natural Circulation Loop Scaling Ratios:

Isochronicity 4-63 i Table 4-8 Constituent Level Scaling Analysis: Two-Phase Mixture Control Volume Balance Equations for the Core as Derived in Appendix A 4-64 Table 4-9 Constituent Level Scaling Analysis: Two-Phase Mixture Non-Dimensionalized Balance Equations for the Core 4-65 l

l \

l Table 4-10 Constituent Level Scaling Analysis: Two-Phase Mixture Residence l

l Times and Characteristic Time Ratios 4-66 '

Table 4-11 Two-Phase Constituent Level Scaling Analysis: Process Specific l

Frequencies for the Core 4-67 l

Table 4-12 Steady-State Loop Balance and State Equations for Two-Phase Natural Circulation Flow 4-68 3681w-1.non:lb-060897 vi

FACILITY SCALING REPORT

(~T! LIST OF TABLES (Cont.)

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Table Title Page Table 4-13 Equation for Core Inlet Fluid Velocity Under Two-Phase Natural Circulation Conditions 4-69 Table 4-14 Steady-State, Two-Phase Natural Circulation Loop Scaling Ratios for Saturated Conditions 4-70 Table 4-15 Steady-State, Two-Phase Natural Circulation Loop Scaling Ratios:

(With Property Similitude) 4-71 Table 4-16 - System Scaling Ratios for Steady-State Natural Circulation with Single-Phase and Two-Phase Flow Regions (Material Property Similitude and Fixed Length Ratio) 4-72 '

Table 4-17 System Scaling Ratios for Steady-State Natural Circulation with Single-Phase and Two-Phase Flow Regions (Property Similitude and Fixed Length Ratio) 4-73 Table 4-18 Summary of System Scaling Results for the 1/4 Length Scale Model Primary Loop (Property Similitude) 4-74 Table 4-19 Two-Phase Flow Transitions in the Loop Legs (Fluid Property (m'} Similitude) and Pressurizer Surge Line 4-75

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Table 4-20 APEX Core Heater Bundle Dimensions and Power 4-75 Table 4-21 Axial Power Fractions for the APEX Core 4-76 Table 4-22 OSU APEX Primary Loop and Core Scaling Ratios 4-77 Table 4-23 OSU APEX Primary Loop and Core Design Specifications 4-78 Table 4-24 Evaluation of Single-Phase Natural Circulation Residence Times, l

Characteristic Time Ratios, and Specific Frequencies (Isochronicity,  !

Pressure Scaled) 4-79 Table 4-25 Evaluation of Single-Phase Natural Circulation Residence Times, i Characteristic Time Ratios, and Specific Frequencies (Isochronicity,  !

Property Similitude) 4-80 l

l Table 4-26 Evaluation of the Two-Phase Natural Circulation Residence Times, Characteristic Time Ratios, and Specific Frequencies (Pressure Scaled) 4-81 Table 4-27 Evaluation of Two-Phase Natural Circulation Residence Times,

/N Characteristic Time Ratios, and Specific Frequencies (Fluid Property

) Similitude) 4-82 l

l 3681w.l.non:1b-060897 vii j

FACILITY SCALING REPORT LIST OF TABLES (Cont.)

Title Table P, ate Table 5-1 System Level Scaling Analysis: System Time Constant and Characteristic Time Ratios 5-58 Table 5 2 Scaling Ratios for System Depressurization Events Dominated by Break or Vent Path Flow Rate 5-59 Table 5-3 Scaling Ratios for System Depressurization Events Dominated by Volumetric Expansion 5-60 Table 5-4 Ideal Initial Conditions for a Depressurization Transient 5-61 Table 5-5 Homogeneous Equilibrium Model (HEM) Critical Mass Flux 0 SX350.2 5-62 Table 5-6 Break and Vent Path Flow Diameters 5-63 Table 5-7 Stored Energy of Reactor Vessel Structural Components 5-67 Table 5-8 Downcomer Stored Energy 5-69 Table 5-9 Heat Capacity Ratios and Total Energy Release Ratios for Reactor Vessel Stnictural Components 5-70 Table 5-10 Steam Generator Scaling Ratios and Dimensions 5-71 Table 5-11 Pressurizer Scaling Ratios and Dimensions 5-72 Table 5-12 Passive Residual Heat Removal Heat Exchanger Scaling Ratios and Dimensions (Single Heat Exchanger) 5-73 Table 5-13 Evaluation of Two-Phase Fluid Depressurization Residence Times, Characteristic Time Ratios and Specific Frequencies for a Two-Inch Cold Leg Break 5-74 Table 5-14 Evaluation of Two-Phase Fluid Depressurization Residence Times, Characteristic Time Ratios, and Specific Frequencies for a Double-Ended DVI Break 5-75 Tabr 4 15 Evaluation of Two-Phase Fluid Depressurization Residence Times, Characteristic Time Ratios, and Specific Frequencies for a One-Inch Cold Leg Break 5-76 Table 6-1 Top-Down Subsystem Level Scaling Analysis Dimensionless Equations for the CMT Draining Processes (Pre-Heated Walls) 6-32 368Iw-1.non;Ib-060897 viii

FACILITY SCALING REPORT l (3 LIST OF TABLES (Cont.)

V I l

Table Title Pate l Table 6-2 Top-Down Subsystem Level Scaling Analysis Dimensionless Equations for CMT Draining Processes (Cold Walls) 6-33 l Table 6-3 Model CMT Scaling Ratios and Dimensions 6-35 Table 6-4 Evaluation of CMT Draining Following Prolonged CMT Loop l Circulation (Hot CMT Walls) Residence Times, Characteristic Time l

Ratios, and Specific Frequencies 6-36 l Table 6-5 Evaluation of CMT Draining with Cold Walls. Residence Times, Characteristic Time Ratios, and Specific Frequencies 6-37 1

Table 7-1 Top-Down Subsystem Level Scaling Balance Equations for Safety Injection Systems 7-51 Table 7-2 Top-Down Subsystem Level Scaling Analysis: Control Volume Balance Equations for Safety Injection Tank Draining (With Simplifying Assumptions) 7-51 g3 Table 7-3 Set of Initial and Boundary Conditions Used to Non-Dimensionalize

( ) the Safety Injection Tank Balance Equations 7-52 .

w/ l Table 7-4 Non-Dimensionalized Balance Equations for Safety Injection Tank Draining 7-53 Table 7-5 Non-Dimensionalized Balance Equations for Accumulator Injection 7-54 Table 7-6 Model Accumulator Scaling Ratios and Dimensions 7-55 l Table 7-7 Model Accumulator Scaling Ratios and Dimensions that Satisfy the Transition Pressure Requirement 7-56 l Table 7-8 Accumulator Time Constants, Residence Time Ratios, and Property Ratios 7-56  !

l Table 7-9 Non-Dimensionalized Balance Equations for IRWST Injection 7-57 l Table 7-10 Model IRWST Scaling Ratio and Dimensions 7-59 6

Table 7-11 IRWST Time Constants, Process Specific Frequencies, Characteristic Time Ratios, and Distortion Factors 7-60 Table 7-12 Control Volume Balance Equations for the i* Section of a Safety Injection Line 7-61 (N Table 7-13 Scaling Ratios for Safety Injection Line Resistance 7-61 L/ )

3681w l.non:lb-080497 iX

FACILITY SCALING REPORT LIST OF TABLES (Cont.)

Table Title Page Table 7-14 Model Safety Injection Line Scaling Ratios and Dimensions 7-62 Table 7-15 DVI Line Scaling Ratios and Dimensions 7-65 Table 7-16 Control Volume Balance Equations for the i* Section of a Vent Line (Two-Phase Homogeneous Mixture) 7-65 Table 7-17 Scaling Ratios for CMT Balance Lines and ADS Vent Lines 7-66 Table 7-18 Balance Line Scaling Ratios and Dimensions 7-67 Table 7-19 ADS 1-3 Vent Line Single and Combined Train Dimensions 7-68 Table 7-20 ADS 4 Line Scaling Ratios and Dimensions 7-69 Table 7-21 ADS 1-3 Sparger Scaling Ratios and Dimensions 7-70 Table 7 22 Upper Core Support Plate Perforation Scaling Ratio and Dimensions 7-71 Table 7-23 Control Volume Balance Equations for the Reactor Vessel Downcomer 7-71 Table 7-24 Non-Dimensionalized Balance Equations for Downcomer Liquid Transport Processes 7-72 Table 7-25 Downcomer Scaling Ratios and Dimensions 7-74 Table 7-26 DVI Diffuser Dimensions 7-75 Table 7-27 E /aluation of Downcomer Fluid Residence Times, Characteristics Time Ratios, and Specific Frequencies 7-76 Table 8-1 Constituent Level Scaling Analysis: Two-Phase Mixture Control Volume Balance Equations for the Core as Derived in Appendix A 8-8 Table 8-2 Constituent Level Scaling Analysis: Two-Phase Mixture Non-Dimensionalized Balance Equations for the Core 8-9 Table 8-3 Constituent Level Scaling Analysis: Two-Phase Mixture Residence Times and Characteristic Time Ratios 8-10 Table 8-4 Two-Phase Constituent Level Scaling Analysis: Process Specific Frequencies for the Core 8-11 Table 8-5 Steady-State Recirculation Loop Balance and State Equations 8-12 368Iw-t.non:Ib-080497 X

( FACILITY SCALING REPORT l(O l

] LIST OF TABLES (Cont.)

Table Title M i

Table 8-6 System Scaling Ratios for Steady-State LCS Recirculation with Single- l Phase and Two-Phase Fluid Regions (Property Similitude and Fixed Length Ratio) 8-13 Table 8-7 Lower Containment Sump Recirculation Line Scaling Ratios and Dimensions 8-14 Table 8-8 Control Volume Balance Equation and Non-Dimensionalized Balance I Equation for Containment Sump Filling 8-15 )

1 Table 8-9 Containment Sump Scaling Ratios and Dimensions 8-16 l 1

l Table 8-10 Evaluation of Two-Phase LCS Recirculation Residence Times, j Characteristic Time Ratios, and Specific Frequencies l (Fluid Property Similitude) 8-17 l

l )

Table 9-1 Initial Conditions for APEX Test Facility to Model a Two-Inch Cold Leg Break 9-4 l,y Table 9-2 Scale Factors to Relate the AP600 Plant to OSU NOTRUMP Calculations 9-5 1> a

! Table 10-1 Summary of Characteristic Time Ratios and Residence Times for the Dominant 10-5 Table 10-2 Summary of Residence Time Constant Scaling (Desired Value 1 (Tcy)g = 0.5) 10-7 l Table 10-3 Distortion Factors for *he AP600 Dominant Processes Identified Using the H2TS Methodology 10-8 Table 10-4 Critical Attributes for the OSU APEX Test Facility 10-9 l

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1 3681w.l.non:lb-060897 Xi J

FACILITY SCALING REPORT LIST OF FIGURES FJure ' Title Page Figure 1-1 General Scaling Methodology 1-20 Figure 1-2 Flow Diagram for the Hierarchical, Two-Tiered Scaling Analysis (NUREG/CR-5809) 1-21 Figure 1-3 Decomposition Paradigm and Hierarchy (NUREG/CR-5809) 1-22 Figure 1-4 Friction Factor Ratio (fgf )pas a Function of Single-Phase or Two-Phase Reynolds Number for a 1/4 Length Scale Natural Circulation System with Fluid Property Similitude 1-23 Figure 1-5 Diameter Ratios Required to Satisfy Equation 1-20 1-23 Figure 1-6 Scaling Ratio Variation as a Function of Length Scale [The Diameter Ratio Represents the Minimum Required to Satisfy Equation (1-20)] 1 24 l Figure 2-1 General Scaling Methodology for the AP600 Test Facility 2-5 I

I Figure 2-2 AP600 Reactor Coolant System Decomposition and Hierarchy (Process) 2-6 l

l Figure 2-3 AP600 Passive Safety System Decomposition and Hierarchy 2-7 Figure 3-1 Diagram of the Westinghouse AP600 System 3-26 Figure 3-2 Flow Diagram for AP600 Passive Safety System Operation 3-27 Figure 3-3 AP600 Large Break LOCA Scenario 3-18 Figure 3-4 AP600 Small Break LOCA Scenario 3-29 Figure 3-5 AP600 LOCA PPIRT Development Methodology 3-30 Figure 4-1 Scaling Analysis Flow Diagram for Single-Phase Natural Circulation 4-83 Figure 4-2 Hot and Cold Leg Regions of Single-Phase Natural Circulation Flow Within a PWR 4-84 Figure 4-3 Scaling Analysis Flow Diagram for Single-Phase Natural Circulation 4-85 Figure 4-4 Regions of Single-Phase and Two-Phase Natural Circulation Within a PWR 4-86 Figure 4-5 Scaling Ratios for Steady-State Natural Circulation vs. AP600 System Pressure (Model Pressure = 375 psia) 4-87 Fic"re 4-6 Scaling Ratios for Steady-State Natural Circulation vs. AP600 System Pressure (Model Pressure = 300 psia) 4-87 36stw-1.non:ib.oso497 xii Rev.I

FACILITY SCALING REPORT

/* LIST OF FIGURES (Cont.)

.(

Figure Title Eagea i

Figure 4-7 Scaling Ratios for Steady-State Natural Circulation vs. AP600 l System Pressure (Model Pressure = 200 psia) 4-88 j Figure 4-8 Scaling Ratios for Steady-State Natural Circulation vs. AP600 System Pressure (Model Pressure = 100 psia) 4-88 Figure 4-9 Model Power Requirements vs. AP600 System Pressure for Steady-State L Natural Circulation (Model Pressure = 375 psia) 4-89 Figure 4-10 Model Power Requirements vs. AP600 System Pressure for Steady-State Natural Circulation (Model Pressure = 300 psia) 4-89 I

Figure 4-11 Model Power Requirements vs. AP600 System Pressure for Steady-State Natural Circulation (Model Pressure = 200 psia) 4-90 l

Figure 4-12 Model Power Requirements vs. AP600 System Pressure for Steady-State l Natural Circulation (Model Pressure = 100 psia) 4-90 l l Figure 4-13 Flow Regime Transition Boundaries for AP600 and OSU Model Hot Legs 4-91 A  !

Figure 4-14 j]

A Dimensionless Diameter (D*) for the AP600 and OSU Model Hot Legs 4-91 ,

l' l,

Figure 4-15 Dimensionless Diameter (D*) for the AP600 and OSU Model Cold Legs 4-92 Figure 4-16 Dimensionless Diameter (D*) for the AP600 and OSU Model Pressurizer Surge Line 4-92 Figure 4-17 Critical Heat Flux Similarity Criteria for the AP600 Test Facility 4-93 Figure 4-18 Critical Heat Flux Similarity Criteria for the AP600 Test Facility 4-93 Figure 4-19 Critical Heat Flux Similarity Criteria for the AP600 Test Facility 4-94 Figure 4-20 Critical Heat Flux Similarity Criteria for the AP600 Test Facility 4-94 Figure 4-21 Axial Linear Power Profile (Normalized) for the Model Core 4-95 Figure 4-22 Radial Power Distribution in Power Core 4-96 Figure 5-1 Scaling Analysis Flow Diagram for System Depressurization 5-77 Figure 5-2 Typical Depressurization Curve for Small Break Loss-of-Coolant Accidents 5-78 Figure 5-3 Control Volume Representation of the Primary System 5-79

.J 3681w-1.non:lN)60897 xiii

FACILrrY SCALING REPORT LIST OF FIGURES (Cont.)

! Figure Title h

' Figure 5-4 Comparison of Critical Mass Flux Ratios (G, /G,,) vs. Pressure Ratio (P/Po ) as Predicted by the HEM (Boxes) for Isentropic Expansion and by Equation 5-43 (Solid Line) 5-80 Figure 5-5 Comparison of Equation (5-57) (solid line) to Ma viken Data (boxes) and RELAPS Calculation (crosses) 5-81 Figure 5-6 Pressure Scaling Relationship Between the AP600 and the APEX Model 5-82 Figure 5-7 Fluid Property Scaling Ratios as a Function of APEX Model System Pressure 5-83 Figure 5-8 Natural Circulation Scaling Ratios as a Function of APEX Model Pressure for Depressurization Transients 5-84 Figure 5-9 AP600 Time Dependent Decay Power Based on 1979 ANSI Standard 5-85 Figure 5-10 APEX Decay Power Profile 5-86 Figure 5-11 Integrated APEX Core Power 5-87 Figure 6-1 AP600 Passive Safety System Design 6-38 O

Figure 6-2 AP600 Core Makeup Tank 6-39 Figure 6-3 AP600 CMT Piping Layout 6-40 Figure 6-4 AP600 SSAR Calculation of CMT Draining Flow for 2-Inch Cold Leg Break 6-41 Figure 6-5 Cold Leg Balance Line Void Fraction for 2-Inch Cold Leg Break 6-42 Figure 6-6 AP600 Mass Flow Rate in Cold Leg Balance Line 6-43 Figure 6-7 AP600 Plant Hot Water Layer Thickness in CMT 6-44 Figure 6-8 OSU Mass Flow Rate in the Cold Leg Balance Line 6-45 Figure 6-9 OSU Hot Water Layer Thickness in the CMT 6-46 Figure 6-10 Ratio of Rescaled OSU to AP600 Plant Mass Flow 6-47 Figure 6-11 Ratio of Rescaled OSU to AP600 Plant Hot Water Layer Thickness 6-48 Figure 6-12 Scaling Analysis Flow Diagram for CMT Condensation and Draining Processes 6-49 3681w-l.non:lb-060897 xiv

FACILITY SCALING REPORT l

Q LIST OF FIGURES (Cont.)

b Figure Title Page Figure 6-13 Idealized Model for CMT Cold Wall Condensation Behavior 6-50 Figure 6-14 AP600 CMT Wall Heat Up Rate for Different Fluid Volumes at 1080 psia 6-51 Figure 6-15 AP600 CMT Wall Heat Up Rate for Different Fluid Volumes at 800 psia 6-51 Figure 6-16 AP600 CMT Wall Heat Up Rate for Different Fluid Volumes ,

j at 400 psia 6-52 l i

Figure 6-17 AP600 CMT Wall Heat Up Rate for Different Fluid Volumes at 200 psia 6-52 Figure 618 AP600 CMT Wall Heat Up Rate for Different Fluid Volumes at 50 psia 6-53 Figure 6-19 APEX CMT Wall Heat Up Rate for Differeni luid Volumes at 385 psia 6-53 Figure 6-20 APEX CMT Wall Heat Up Rate for Different Fluid Volumes at 285 psia 6-54 Figure 6-21 APEX CMT Wall Heat Up Rate for Different Fluid Volumes at 142.6 psia 6-54 Figure 6-22 APEX CMT Wall Heat Up Rate for Different Fluid Volumes at 73.3 psia 6-55 Figure 6-23 APEX CMT Wall Heat Up Rate for Different Fluid Volumes at 17.8 psia 6-55 Figure 7-1 Control Volume Boundaries for Safety Tank Draining Analysis 7-77 Figure 7-2 Flow Diagram for the Accumulator Scaling Analysis 7-78 Figure 7-3 Scaling Analysis Flow Diagram for the IRWST Draining Process 7-79 Figure 7-4 Control Volume for Safety Injection Line (Actual piping will have various geometries and fittings) 7-80 Figure 7-5 Scaling Analysis Flow Diagram for CMT Balance Lines and ADS Vent Lines 7-81

/

l \s} -

I Figure 7-6 Control Volume for Sections of a Vent Line or Balance Line (Actual geometries will vary) 7-82 3681w.l.non:1b 080497 xy

FACILITY SCALING REPORT LIST OF FIGURES (Cont.)

! Ficure Title Page l

I Figure 7-7 Scaling Analysis Flow Diagram for Downcomer Phenomena 7-83 Figure 8-1 Long-Term Recirculation Cooling Mechanism 8-18 Figure 8-2 Scaling Analysis Flow Diagram for the Lower Containment Sump Recirculation Subsequent to Sump Flood-up 8-19 Figure 8-3 W GOTHIC Calculation for Containment Pressure During a 2-Inch Break in the AP600 8-20 Figure 8-4 W GOTHIC Calculation for Containment Pressure During a Double-Ended Guillotine Break of the AP600 DVI Line 8-21 Figure 9-1 AP600 Plant Pressurizer and Steam Generator Pressures for a Two-Inch Cold Leg Break 9-6 Figure 9-2 APEX Test Facility Pressurizer and Steam Generator Pressures for an Equivalent Two-Inch Cold Leg Break 9-7 Figure 9-3 Normalized Pressure Comparisons Between AP600 Plant and APEX Facility 9-8 Figure 9-4 Normalized CMTl Level for AP600 Plant and APEX Facility 9-9 Figure 9-5 Normalized CMT2 Level for AP600 Plant and APEX Facility 9-10 Figure 9-6 Normalized Accumulator 1 Level for AP600 Plant and APEX Facility 9-11 Figure 9-7 Normalize : Accumulator 2 Level for AP600 Plant and APEX Facility 9-12 Figure 9-8 Normalized ADS 1-3 Flows for AP600 Plant and APEX Facility 9-13 Figure 949 Normalized Break Flow for AP600 Plant and APEX Facility 9-14 Figure 9-10 Normalized System Mass for AP600 Plant and APEX Facility 9-15 Figure 10-1 Characteristic Time Ratios as a Function of Process Specific Frequency for Single-Phase and Two-Phase Natural Circulation and Long-Term Recirculation (Sections 4.0 and 8.0) 10-10 Figure 10-2 Characteristic Time Ratios as a Function of Process Specific Frequency for System Depressurization (Section 5.0) 10-11 Figure 10-3 Characteristic Time Ratios as a Function of Process Specific Frequency for all AP600 Transport Processes Identified by SBLOCA PPIRT 10-12 3681w 1.non-lb-08G497

. XVi

FACILITY SCALING REPORT EXECUTIVE

SUMMARY

Westinghouse uses passive safety systems to enhance the overall safety of the AP600 Pressurized Water Reactor (PWR) power plant. The passive systems perform the same safety function as the l current safety systems in operating PWRs; however, they perform this function using gravity forces and natural circulation to provide the flow through the reactor system. This eliminates the need for pump-driven systems to provide the core cooling needed to maintain a low reactor coolant temperature during postulated accidents.

Westinghouse and the U.S. Nuclear Regulatory Commission (USNRC) will evaluate the performance of the AP600 passive safety system using safety analysis computer codes, which have been developed for these advanced plants. The transients in which the passive systems are most active include the small break loss-of-coolant accident (SBLOCA) and long-term cooling.

The Advanced Plant Experiment (APEX) facility at Oregon State University (OSU) has been desiped to develop a database to validate the safety analysis codes that will be used to predict the performance of the AP600 passive safety systems.

To ensure that the APEX experiments provide valid thermal-hydraulic data and capture the key phenomena over the range of conditions for which the passive safety systems require validation, a detailed state-of-the-art scaling analysis has been performed to establish the facility design and test conditions. The following sections present a summary of the scaling process applied to the OSU APEX experiments.

PLAUSIBLE PHENOMENA To determine which of the AP600 features should be incorporated into the APEX design, it was first necessary to identify the phenomena that are important to AP600 passive safety system performance.

This was done by developing a Plausible Phenomena Identification and Ranking Table (PPIRT) for AP600 small break loss-of-coolant accidents (SBLOCAs). Existing data on standard PWRs, coupled with engineering judgement and calculations for the AP600, were used to determine which thermal-hydraulic phenomena might impact core liquid inventon or fuel peak clad temperature. These phenomena have been ranked and included in the AP600 SBLOCA PPIRT.

H2TS SCALING METHODOLOGY Having identified the important phenomena through the use of a SBLOCA PPIRT, the APEX scaling analysis was divided into four parts. Each part corresponds to a different phase of a SBLOCA:

= Natural circulation scaling analysis

• System depressurization scaling analysis 3681w-1.non:Ib-060897 1

FACILITY SCALING REPORT

. Venting, draining, and injection scaling analysis l

. Recirculation scaling analysis The Hierarchial, Two-Tiered Scaling Analysis (H2TS) method was used to analyze each phase of the SBLOCA transient. The H2TS method is an essential pan of the NRC's Severe Accident Scaling Methodology (SASM).

There are four basic elements of the H2TS analysis method:

• System decomposition

= Scale identification

. Top-down/ system scaling analysis

. Bottom-up/ process scaling analysis For the APEX scaling analysis, the AP600 is divided into a reactor coolant system (RCS) and a passive safety system. Each of these systems can be subdivided into interacting subsystems (or modules), which can be further subdivided into interacting constituents (materials). The interacting constituents can be further subdivided into interacting phases (liquid, vapor, or solid). Each phase can be characterized by one or more geometrical configurations, and each geometrical configuration can be described by three field equations (mass, energy, and momentum conservation equations). Each field equation can be characterized by several processes.

After dividing the system of interest, the next step is to identify the scaling level at which the similarity criteria should be developed. This is determined by the phenomenon being considered. For example, thermal-hydraulic phenomena involving integral system interactions, such as primary system depressurization or loop natural circulation, are examined at the system level. Phenomena such as passive residual heat removal (PRHR) decay heat removal, core makeup tank (CMT), accumu.lator (ACC), and in-containment refueling water storage tank (IRWST) passive safety injection, automatic depressurization and lower containment sump (LCS) recirculation cooling are examined at the subsystem level. Thermal-hydraulic phenomena important to individual components such as the reactor core, pressurizer, steam generators, hot legs, cold legs, coolant pumps, and interconnecting piping are examined at the component level. Specific interactions between the steam-liquid mixture and the stainless steel structure are examined at the constituent level.

The OSU APEX scaling study presents scaling analyses performed at different levels. The thermal-hydraulic phenomena of interest, the system at which the analysis will be performed, the control volume for the analysis (that is, the geometrical configuration), the applicable balance equations, and the processes important to the thermal-hydraulic phenomena of interest are discussed and analyzed for the integral reactor system and the major components in the system.

The third element of the H2TS method requires performing a top-down (system) scaling analysis. The top down scaling analysis examines the synergistic effects of the system caused by complex 368Iw-1.non.Ib-060897 2

FACILITY SCALING REPORT l

(] interactions between the constituen's which are deemed important by the PPIRT. The top down scaling l

V approach uses the conservation equations at a given scaling level to obtain characteristic time ratios l and similarity criteria. It also identifies important processes to be addressed in the bottom-up scaling l

analysis.

The fourth element of the H2TS method requires performing a bottom-up (process) scaling analysis.

This analysis develops the similarity criteria for specific processes such as flow pattern transitions and geometry and flow-dependent heat transfer. The focus of the bottom-up scaling analysis is to develop similarity criteria to scale individual processes of importance to system behavior as identified by the PPIRT and to develop the design information for the test facility.

Top-down and bottom-up scaling analyses were performed for each of the four phases of the AP600 SBLOCA. This resulted in obtaining the set of similarity criteria that would best preserve the phenomena of importance in the AP600 reactor coolant system (RCS), the passive safety system components (accumulators, core makeup tanks (CMTs), passive residual heat removal (PRHR) heat exchanger, in-containment refueling water storage tank (IRWST), reactor sump, containment compartments sump), and the primary side and secondary side of the steam generators. Similarity criteria were also obtained for the single-phase and two-phase pressure losses for all the vent, balance, and injection lines.

f)

LJ SCALING CHOICES / DESIGN CLOSURE Having obtained the set of similarity criteria for the important phenomena, the scaling study then requires that the analyst make the following choices to obtain the physical dimensions for the test facility:

= Selection of the working fluid j

= Selection of component materials l

= Selection of operating pressure

= Selection of length, diameter, and time scale j The bases for these selections include power requirements, the geometric representation of the facility relative to the prototype, ease of interpreting test results, economics, operability, and instrument accuracy.

There are other scaling considerations, such as the flow regimes in the loop piping. Small diameters distort the flow regime and have different two-phase counterblow behavior compared to the prototype.

Reduced size can also cause manufacturing problems for the core heater simulators. For the APEX facility, studies were performed to examine the relative benefits of a 1/8th ,1/6th , and 1/4th-length scaled facility. After an evaluation,it was decided that a 1/4th-length scale was the most appropriate since it minimized the power re-quirements while maximizing the height. A 1/4th-scaled facility also

(}

had sufficient volume and size to correctly model the prototype pressure drop and possible 368Iw.l.non:Ib-o60897 3

1 FACILITY SCALING REPORT three-dimensional behavior that could occur in the simulated reactor vessel, plenums, and downcomer.

To choose a consistent diamem scale, a simple relationship was derived from the one-dimensional momentum equation to relate the length ratio to the diameter ratio such that frictional pressure drops in the model would not exceed the total pressure drop, on a scaled basis, in the AP600. The choice of the diameter ratio was further verified with a bottom-up scaling approach in which the two-phase flow regimes and transitions between flow regimes were examined. The possible distortions of the flow regimes and their transitions were also examined for the horizontal piping. Flooding studies were performed to verify that the chosen diameter ratio would have minimum surface tension effects if flooding would occur.

Using the above approach, the facility dimensions could then be specified with confidence that the key parameters and phenomena that were identified in the PPIRT would be preserved in the APEX facility.

TEST BOUNDARY AND INITIAL CONDITIONS To model the SBLOCA, the APEX experiments are designed to start in a liquid solid recirculation mode with the simulated reactor coolant pumps (RCPs) operating at full flow and with system pressure at approximately 400 psia. A break is then initiated and the system begins to depressurize, not only due to the break, but also due to the activation of the automatic depressurization system (ADS) system. To preserve the depressurization behavior of the APEX facility, a reference pressure must be selected and a scaling rationale developed to relate the lmer pressure APEX tests to the higher pressure AP600 for the depressurization transients. Property scaling relations have been developed to relate APEX and AP600 pressures and fluid properties. These relations enable the analyst to select the appropriate boundary and initial conditions for APEX testing. This includes initial system pressure, facility break areas, ADS valve areas, accumulator gas pressure and the steam generator secondary-side f safety pressures. The property scaling relations also permit the analyst to relate the test results to full-scale AP600 conditions. ,

I l

For a small break LOCA, the AP600 primary system pressure will stabilize, after the initial subcooled blowdown, to a near constant value, slightly above the safety valve set-point for the steam generator secondary side. The primary pressure will remain at this value for a relatively long period, depending upon the break size and when the ADS activates. This pressure was chosen as the reference pressure for the AP600 and APEX. Subsequent to ADS activation, the primary system will rapidly depressurize to the containment pressure. During this time period, the passive safety systems of the AP600 will be in operation and the phenomena of importance, which was identified in the PPIRT, will be present.

SCALING DISTORTIONS It is impossible to identically preserve all of the similarity criteria in a scale model. Therefore, scaling distortions will arise. To assess the applicability of the data obtained in the test facility, it is necessary to identify and evaluate the impact of scaling distortions on the dominant phenomena. The H2TS method addresses this concern through the use of distortion factors that can be calculated from the 368 t w.l.non:lb-060897 4

FACILITY SCALLNG REPORT O characteristic time ratios obtained for each of the important processes. The distortion factors specify the degree of disparity in rate processes occurring in the AP600 and in the APEX on a scaled basis.

The scaling analysis identifies the dominant phenomena and the scaling distortion associated with each of those phenomena. The results of the scaling distortion analysis indicate that the APEX facility will adequately simulate the important thermal-hydraulic behavior of the AP600.

CONCLUSIONS The H2TS method provided a thorough, traceable, and auditable approach to obtain all of the information needed to design and constmet the APEX facility. The scaling rationale and approach was confirmed by modeling the test facility with the same safety analysis computer code used for the

' AP600 plant analysis. When the scaling parameters were applied to the calculated plant results and compared to the APEX transient system, excellent agreement was achieved on the system behavior, timing, pressure history, levels, mass inventory, and flows. This analysis confirms the scaling basis for the OSU APEX facility and verifies that these experiments will capture the key thermal-hydraulic l phenomena needed for safety analysis computer code verification.

O l

O 3681w 1.non:lb 060897 5

FACILITY SCALING REPORT l

ABSTRACT l

l l I

The Westinghouse Electric Corporation (W) has proposed a research effort the Nuclear Engmeenng Department of Oregon State University (OSU) to provide data to verify the integral system and long- I term cooling behavior of Westinghouse's next generation of nuclear reactor system, the AP600. This research is supported by Westinghouse, OSU, the Portland General Electric Company (PGE), and FLUKE Electronics.

This report presents the scaling analysis for a 1/4-length scale experiment that will simulate the integral system and long-term cooling behavior of the Westinghouse AP600 nuclear steam supply system. Under certain accident conditions, the AP600 is designed to bring the reactor system to low pressure while providing gravity driven, natural circulation cooling to the reactor core. Because the AP600 operates at low pressure during the long-term cooling process, a low pressure test facility is ideally suited to examining the thermal-hydraulic behavior of interest. Furthermore, the scaling analysis is greatly simplified by using the same piping materials and fluids. The report also presents the test conditions under which high pressure AP600 behavior can be simulated in the OSU Advanced Plant Experiment (APEX) facility.

The Hierarchical, Two-Tiered Scaling (H2TS) analysis method recently developed by the U.S. Nuclear Regulatory Commission (USNRC), has been implemented. (N. Zuber, " Appendix D: A Hierarchical, Two-Tiered Scaling Analysis," An integrated Structure and Scaling Methodologyfor Severe Accident TechnicalIssue Resolution, U.S. Nuclear Regulatory Commission, NUREG/CR-5809). This methodology provides a systematic approach to assure that the AP600 processes of interest are properly scaled and to assess the effects of any scaling distortions which may be present. The test facility dimensions were selected based on the results of the scaling analysis. The scaling analysis indicates that a 1/4-length scale, stainless steel model can be constructed to simulate the thermal.

hydraulic phenomena of interest.

O 368Iw 1.non:It>060897 6

FACILITY SCALING REPORT ACKNOWLEDGEMENTS a

Westinghouse would like to recognize the original work of Dr. Jose Reyes for the Scaling Report- in particular, Sections 4.0 and 5.0. In addition to the significant contributions to the state-of-the-art of scaling, the capability to accurately model the AP600 cooling phenomena was instmmental in the AP600 Design Certification Process.

The authors wish to acknowledge Dr. Novak Zuber, Dr. Wolfgang Wulff, Dr. Matti Merrilo, the members of the ACRS thermal-hydraulic subcommittee, and the members of the Utility Steering Committee for their valuable comments. The authors acknowledge Mr. Louis Lau of Westinghouse for his efforts in translating the scaling results into design drawings, and Mr. Moshe Mahlab and Mr. Eugene Piplica for their overall guidance throughout the project. The authors also acknowledge Mr. John Groome of Oregon State University for his practical insights on plant operations.

Special thanks are given to Mrs. Teresa Culver, Ms. Robin Keen, and Mrs. Jennie Smith for their excellent word processing skills, their cooperative spirit, and their dedication to producing a high quality document. Patti Todaro and Becky Jung are recognized for their support in the editing and production of the document.

(O Special thanks are also given to all of our families for their continual support.

N.

I i

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t -%/

l 3681w.l.non:Ib-060897 7

_ _ - - - _ ___---_____ _ a

O O

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FACILITY SCALING REPORT

</

1.0 INTRODUCTION

Advanced reactor designs, which rely on passive safety systems for safety injection and long-term core cooling, are currently bound by the same safety requirements as conventional light water reactors.

Title 10 Past 50.46 of the Code of Federal Regulations specifies the acceptance criteria for emergency core cooling systems (ECCS) for light water nuclear power reactors.W Section (a)(1) of the _

acceptance criteria indicates that analyses of ECCS performance can be done using best-estimate evaluation models provided that the eva'aation model includes:

... sufficient supporting justifiution to show that the analytical technique realistically a describes the behavior of f a reactor system during a loss-of-coolant accident. l

' Comparisons to applicab: experimental data must be made and uncertainties in the analysis method and input must be identified and assessed so that the uncertainty in the calculated results can be estimated.

3 Furthermore, Section (b)(5) of the acceptance criteria requires that analyses of the ECCS show that:

After any calculated successful initial operation of the ECCS, the calculated core 1

temperature shall be maintained at an acceptably low value and decay heat shall be remo'v ed for the extended period of time required by the long lived radioactivity Jr remaining in the core.

In view of these requirements, a correctly scaled experimental facility that can be used to examine passive safety injection and long-term, gravity-driven, natural circulation cooling of the core is of significant value to the next generation of passively safe nuclear reactor systems. The experimental results obtained in the. test facility will provide data to be used to verify safety analysis methods and computer codes for passive plants.

. Because of the development of new passively safe' reactor systems such as the AP600, the Nuclear

. Engineering Department at Oregon State University (OSU), in conjunction with the Westinghouse Electric Corporation (W) and the U.S. Department of Energy, have proposed a research effort to experimentally investigate the integral system and long-term cooling behavior of the Westinghouse AP600 advanced reactor system.

, 1.11 Scaling Objectives -

- The general objective of this scaling study is to obtain the physical dimensions of a test facility that L will simulate the flow and heat transfer behavior of importance to AP600 passive safety system O

M81w.llnon:Ib-060897 11

FACILITY SCALING REPORT operation and long-term cooling. In order to develop a properly scaled test facility, the following specific objectives must be met for each mode of AP600 operation of interest:

. Obtain the similarity groups which should be preserved between the test facility and the full-scale prototype.

• Establish priorities for preserving the similarity groups.

• Assure that important processes have been identified and addressed.

• Provide specifications for test facility design.

. Quantify biases due to scaling distortions.

Different sets of similarity criteria are cbtained for the different modes of operation. The similarity criteria depend on the geometry of the components, the scaling level required to address the transport phenomena of interest, and the initial and boundary conditions for each particular mode of operation.

The experimental data obtained from this test facility will be used to assess the computational methods that will be used to analyze the passive safety system and long-term cooling performance of the AP600.

1.2 General Scaling Methodology Meeting the scaling objectives of the previous section presents a formidable challenge. Therefore, to assure that these objectives are met in an organized and clearly traceable manner, a general scaling methodology (GSM) for the OSU Advanced Plant Experimental (APEX) test facility has been developed. The model for this scaling methodology is partly drawn from the USNRC's Severe Accident Scaling Methodology (SASM) presented in NUREG/CR-5809.(2) A flow diagram describing the GSM is presented in Figure 1-1.

The first task outlined by the GSM is to specify the experimental objectives. This is done in Section 2.0. The experimental objectives define the types of tests that will be performed to respond to specific licensing and design needs. These objectives determine the general modes of operation that should be simulated in the test facility.

The second task outlined by the GSM is the development of Plausible Phenomena Identification Ranking Tables (PPIRTs). The nature of caling forbids exact similitude between the AP600 and the test facility operating conditions. As a result, the design and operation of the test facility will be based on simulating the processes most important to passive safety system performance and long-term cooling. The function of the PIRTs is to identify the key thermal-hydraulic phenomena that should be scaled in the context of loss-of-coolant accident (LOCA) transients. Many of the phenomena of importance to AP600 LOCA behavior have already been identified by existing Phenomena Identification and Ranking Tables (PIRTs),(3) However, some of the AP600 modes of operation have never been verified. Therefore, the first series of tests may help identify additional thermal-hydraulic 3681w.l .non:lts.060897 ].2

FACIIJn' SCALING REPORT (N phenomena of importance. Hence the use of PPIRTs rather than PIRTs. The development of the AP600 LOCA PPIRTs is presented in Section 3.0.

The third step in the GSM is to perform a scaling analysis for each of the modes of operation specified by the experimental ol.jectives and further defined by the PPIRTs. There are several methods that can be used to perform these scaling analyses. The most commonly used methods are evaluated in Section 1.3. Becaur,e a single system is being examined, all of the modes of operation share some common scaling features.

The fourth step is to use the scalirg analysis results to develop a set of characteristic time ratios (dimensionless H groups) and similarity crite:ia for each mode of operation. Because it is impossible to identically satisfy all of the simihuity criteria si'nultaneously, the set will include only those criteria which must be satisfied in order to scale the most important phenomena identified by the PPIRT.

Step five is an evaluation of the scaling criteria to determine if the scale model geometry, boundary conditions, or operating conditions introduce significant scaling distortions. Distortions ne also evaluated relative to other modes of operation.

Upon satisfying the important scaling criteria, step six provides the specific component geometries and

,. operating conditions for each mode of operation in the scaled test facility.

( I V

Step seven is an evaluation of the key thermal-hydraulic PPIRT processes to prioritize the system design specifications.

Step eight serves to integrate all of the design requirements for the OSU APEX test facility.

i The remaining sections of this report present the results obtained by executing each step in the general scaling methodology.

1.3 Evaluation of Scaling Analysis Methods Various scaling analysis methods have been used to design small-scale thennal-hydraulic test facilities.

This section presents a brief overview of three scaling analysis methods: the power-to-volume scaling analysis method; the Ishii-Kataoka scaling analysis method; and the Hierarchical, Two-Tiered Scaling (H2TS) analysis method. It is shown that the similarity criteria used in power-to-volume scaling is a subset of the Ishii-Kataoka similarity criteria. Furthermore, the Ishii-Kataoka similarity criteria can l

also be a subset of the H2TS similarity criteria- depending on the problem being addressed by the I H2TS analysis method.

.- s v

3681w 1.non:lb-060897 13 I

l _

FACILITY SCALING REPORT 1.3.1 Power to-Volume Scaling Analysis Method The traditional approach to designing small-scale thermal-hydraulic test facilities has been to use power-to-fluid volume (P/V) scaling. This scaling approach has been successfully applied in various studies such as the FLECHT SEASET program.W Re optimum condition for this scaling approach occurs when the scale model implements the same working fluid as the full-scale system and when the scale model is built using similar materials, is at full height, and is operated at same pressure. This generally results in constructing a very tall and thin scale model. As a result, the hydrodynamic behavior in the plenum regions are fully represented. Furthermore, if a different working fluid is used in the scaled model or if the scale model is operated at a different pressure or uses materials significantly different than the full-scale system, then additional scaling parameters must be considered.

He power-to-volume scaling approach is straight forward. All of the scale prototype component vol-umes are reduced in size by the ratio of the model power to the prototype power to obtain a reduced-scale model. Because the test facility would still be full length, the design constraints and materials expense are often considerable. Furthermore, the resulting facility design will have a distorted surface area to volume ratio and may have an excessive heat loss rate or stored energy release rate.W The power-to-volume scaling analysis method does not represent a comprehensive scaling approach.

If a wide range of phenomena is being considered, then it is only one of many criteria that must be satisfied.

1.3.2 Ishii Kataoka Scaling Analysis Method Another approach to designing small-scale thermal-hydraulic test facilities is to use the similarity criteria for forced and natural circulation flow dedved by Ishii and Kataoka.0) Ishii's criteria permit a variable power-to-fuel volume (q") ratio while maintaining core exit conditions identical in the model and the prototype. These criteria include the power-to-volume similadty criterion. The advantage of using Ishii's power density scale simulation is that a full-length test facility, which implements the same working fluid the same pressure and structural materials, is not needed to satisfy the scaling criteria. This permits added flexibility in the design choices. Thus, a reduced height scale model, which gives a better representation of multi-dimensional effects in the plenum and downcomer regions, can be designed.(6)

Ishii's similarity criteria were developed for single-phase and two-phase fluid flows within a closed loop under forced or natural convection conditions. The focus of this method is on the scaling criteria for natural circulation conditions. Based on a review of Ishii's natural circulation scaling theory, a general statement regarding its applicability can be made:

Ishii's criteria are not intended to scale all of the thermal-hydraulic behavior produced in each component of the full-scale system. However, it does provide a basic framework for scaling the integral system.

3681w.l.non:Ib-060897 ].4 w

FACILITY SCALING REPORT

' l 4

I 4

Because of this, there are several limitations that must be recognized before applying the criteria to the '

design of a thermal-hydraulic test facility.

Limitation 1: Flow Pattern Transitions If one strictly adheres to Ishii scaling for each component in the system, certain flow pattern transitions may be delayed or missed entirely. One reason for this is that Ishii's Froude number is

' based on length scale. This criterion is suitable for flow pattem transitions dominated by phase separation. However for certain flow pattern transitions, a diameter-based Froude number, such as that proposed by Taitel and Dukler,(7) is most appropriate. .

+

l - Depending on the intent of the test, distortions in flow pattern transitions may or may not be significant. In their paper, Schwartzbeck ,

and Kocamustafaogullari(8) indicate that scaling distortions L are unavolobble in transitions. However, it is almost always possible to achieve similarity in flow

[ upattems. 'Iherefore, the designer must specify a priori the thermal-hydraulic phenomena of interest in L the full-scale. system. Having this information, significant two-phase flow pattem transitions in l

specific components can be modeled using similarity criteria specifically developed for that transition.

This may require modifying some of the geometric dimensions of a specific component. l l

. Limitation 2: Drift Flux Number Criteria

(

In their report, Kocamustafaogullari and Ishii(9) indicate that whenever the local slip is large in

{, comparison to the slip due to the transverse velocity and void profiles, the fluid velocity in the model

- is distorted. However, this distortion introduces only limited changes in the void-quality relation.

They indicate that for most cases, the local slip does not dominate; therefore, the drift flux number L criterion can be satisfied. For the core, satisfying the phase change number criterion and the l subcooling number criterion takes precedence over satisfying the drift flux number criterion.

l.

- Limitation 3: Leak Flow Scaling Is'hii's scaling criteria do not address the scaling of breaks in the system. In his thesis, Moskal(W

! recognizes'this limitation and recommends a scaling approach for breaks. I l Limitation 4: Core Heat Transfer i

1

. With respect to the core, Ishii and Kataoka(S present the basic heat transfer correlations for forced and free convection flow in their development of the single-phase scaling criteria. However, these heat transfer correlations are not directly used in the Biot and Stanton numbers to obtain scaling criteria - i I

that are heat transfer regime specific. Selection of the appropriate heat transfer correlation depends on the flow regime expected in the full-scale system. Therefore, the designer must evaluate the flow regimes expected in'the core for the full scale system.

l l

3681w.l.non:lb-060897 1-5 L_ ___ _ _ _ _ _____ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . . . _ _

FACILITY SCALING REPORT The Biot and Stanton number similarity conditions, coupled with the convective heat transfer correlation, simulate the boundary layer temperature drop. If the heat transfer mechanism is not completely simulated, the scaled system will adjust to a different temperature drop in the boundary layer. This will not have a significant impact on the overall flow and energy distribution in the slow transients typical of those to be simulated in the APEX test facility.

Because of the effectiveness of boiling heat transfer, the resulting temperature drop in the thermal boundary layer under two-phase flow conditions is not as important as simulation of the critical heat flux condition, which signifies a flow regime change.

Limitation 5: Steam Generator Heat Transfer Ishii's scaling criteria do not address the scaling of steam generator heat transfer. The paper by Zvirin and SursockW presents scaling criteria for the secondary side which are dependent on heat transfer regime. These criteria arc quite detailed, and it is apparent that all of the criteria cannot be satisfied for a wide range of flow conditions. For practical purposes, it is necessary to design the facility for the dominant type of flows to be examined. With respect to the OSU APEX tests, which involve ADS operation, the steam generators will serve as both a heat source and a heat sink during the transients. When the steam generators are a heat sink, the two-phase primary fluid will be condensing.

When acting as a heat source, the primary-side fluid will be evaporating inside the tubes.

Limitation 6: Loop Flow Rate Scaling O

Most of the reviews of the Ishii-Kataoka loop-integrated scaling criteria recognize that the criteria are rigorously derived and typically can be used to design and evaluate integral test facilitiet A maior concern, however, is preser*-d in the paper by Wang, Hsu, Almenas, and DiMarzo.M The concern arises from Ishii's scaling of the two-phase loop flow velocity. The loop flow velocity is treated as an independent variable. That is, Ishii and Kataoka do not demonstrate how their velocity scaling ratio I satisfies the nonlinear relationship between flow velocity; buoyancy; and loop resistance, which arises when they implement the two-phase loop-integrated momentum equation. In fact, none of the scaling references available in the literature specifically address this issue. This situation is addressed in Section 4.0 of this report. l In summary, these six limitations identified indicate that the Ishii-Kataoka scaling analysis method was l not intended to be a comprehensive method for scaling all thermal-hydraulic phenomena of interest. It will be shown in the Section 1.3.3 that the Ishii-Kataoka scaling analysis method represents a top-down scaling approach performed at the constituent level. The user is left with the responsibility of identifying important processes and performing scaling analyses for those processes (bottom-up l scaling). As indicated in this section, various authors have already performed scaling analyses for processes such as flow pattern transitions, local heat transfer, and critical flow. Therefore, what is l needed is a systematic scaling approach that assures that the important systems and processes are l properly scaled. The Hierarchical, Two-Tiered Scaling (H2TS) analysis method was developed to 3681w l.non:lb-060897 1-6

FACILITY SCAltNG REPORT respond to that need. Therefore, the H2TS method has been applied to the design of the APEX integral system and long-term cooling test facility, implementation of the H2TS method reveals that the Ishii-Kataoka scaling criteria are completely valid, although incomplete for the problem being considered. Furthermore, application of the H2TS method responds to the issue of flow rate scaling and indicates that the Ishii-Kataoka velocity scaling ratio is sometimes used inappropriately.

1 l.3.3 Hierarchical, Two-Tiered Scaling Analysis Method The H2TS method has been used to develop the similarity criteria necessary to scale the systems and processes of importance to AP600 integral system and long-term cooling. The H2TS method was developed by the U.S. Nuclear Regulatory Commission and is fully described in Appendix D of NUREG/CR-5809.(2)

Figure 1-2 is taken from NUREG/CR-5809.(2) It presents the four basic elements of the H2TS analysis method. The first element consists of system decomposition. Each system can be subdivided into interacting subsystems, which can be further subdivided into interacting modules. The interacting modules can be further subdivided into interacting constituents (materials), which can be further subdivided into interacting phases (liquid, vapor, or solid). Each phase can be characterized by one or more geometrical configurations, and each geometrical configuration can be described by three field equations (mass, energy, and momentum conservation equations). Each field equation can be characterized by several processes. This is depicted in Figure 1-3.

After identifying the system of interest and decomposing it as in Figure 1-3, the next step is to identify the scaling level at which the similarity criteria should be developed. This is determined by examining the phenomena being considered.

For example, if the phenomenon being considered involvus mass, tuotticulum, or energy Innspen between materials such a.; water and solid panicles, then the scaling analysis should be performed at the constituent level. If the phenomenon of interest involves mass, momentum, or energy transport between vapor and liquid, then the scaling analysis should be performed at the phase level.

Therefore, identifying the scaling level will depend on the phenomenon being addressed.

Section 2.0 presents a detailed road map for the AP600 system hierarchy. Sections 4.0 through 8.0 present scaling analysis performed at different levels. Each rdon identifies the thermal-hydraulic phenomena of interest, the system level (that is, control volume) at which the analysis will be performed, the geometrical configuration, the applicable balance equations, and the processes important to the thennal-hydraulic phenomena of interest.

The third element of the H2TS method requires performing a top-down (system) scaling analysis. The top-down scaling analysis examines the synergistic effects on the system caused by complex interactions between the constituents which are deemed important by the PPIRT. Its purpose is to use 3681w.).non:lb-060897 ].7

FACILITY SCALING REPORT the conservation equations at a given scaling level to obtain characteristic time ratios and similarity criteria. It also identifies important processes to be addressed in the bottom-up scaling analysis.

The fourth element of the H2TS method requires performing a bottom-up (process) scaling analysis.

This analysis develops similarity criteria for specific processes such as flow pattern transitions and flow-dependent heat transfer. The focus of the bottom-up scaling analysis is to develop similarity criteria to scale individual processes of importance to system behavior as identified by the PPIRT.

Time Ratios The basic objective of the H2TS method is to develop sets of characteristic time ratios for the transfer processes of interest. This can be done by writing the control volume balan .e equations for each constituent k as follows:

kk " (I'l)

Qk Vk)*IlknA kn Defining A[Qgyg]:

kQ kVkMQ Vk[o k @kVk)out where:

yg = the conserved property; p, pu, or pc (mass, momentum, or energy per unit volume)

Vk = the control volume Qg = the volumetric flow rate ju = the flux of property yg ransferred t from constituent k to n across the transfer area Akn Hence, A[Qgyg] represents the usual mass momentum or energy convection terms; and EjknA go represents transport process terms such as condensation or sources such as decay power.

Equation 1-1 can be put in dimensionless form by specifying the foUowing dimensionless groups in terms of the constant initial and boundary conditions:

Yk + , Yk + ,O k , ikn + A kn V*=

g (1-3)

Qo, Vg,o , yg,o k. Jkn.o A kn,0 l

l l

9 368Iw.1.non Ib-07I897 18

FACILITY SCALING REPORT i l l

/N

--+

Substituting these groups into Equation 1-1 yields the following: l l

l i

dV*y*'

g g l- V k,0Vk,0 k,0Vk,00 ,Q A

. . @kn.0 kn,0Mkn.A. (34) dt kVk. kn l

Dividing both sides of this equation by Qg,o yg,o yields the following:

dV 'yg* ++ i (1-5)

Tk dt -

kVk, dDkn+ A +n k )

)

where the residence time of constituent k is the following:

T" V k.0 k (1-6)

O k.0 l

I and the characteristic time ratio for a transfer process between constituents k and n is given by the following:

.{

s kn,0Akn,0 (1-7) j l 4,)Ok,0Vk,0 1

1 Because each transfer process has a characteristic time ratio, the imponance of each process can be

[

1 ranked by comparing the time ratios.

If a specific transfer process is to have the same effect in the prototype and the model, then the characteristic time ratios must be preserved, j l

l

Process Ranking Using Characteristic Time Ratios j 1 i I l can be defined as the set of time ratios that characterize all of the The term SK), (II,.y),.

'~

individual processes which occur during the evolution of a transient. The subscripts i, j, N,,N _ j identify j the specific process, the hierarchical level, the total number of specific processes, and the total number 4 of hierarchical levels.

)

+0.

1

~

3681w l.non:Ib-060897 19

FACILITY SCALING REPORT Because of differences in geometrical scale and fluid properties, it is impossible to exactly duplicate the time ratio set for the full-scale prototype (Sp ) in a reduced-scale model. That is, exact similitude t for all processes cannot be preserved; therefore:

Sp # S, (1-8) where:

p = the full-scale prototype m= the reduced-scale model It is possible to design a reduced-scale test facility that preserves the similitude of a subset of time ratios T [H ] that characterize the processes of greatest imponance to the transient. This optimizes the i3 model design to investigate the important processes while distorting the less important processes.

To determine which processes govem the overall evolution of a transient, numerical estimates of the characteristic time ratios for the prototype and the model must be obtained for each hierarchical level of interest. Physically, each characteristic time ratio (Hi ) is composed of the fc.llowing: a specific frequency (co), i which is an attribute of the specific process, and the residence time constant for the control volume, tcy. That is:

IT = co,tcy (1-9)

The specific frequency defines the mass, momentum or energy transfer rate for a particular process.

The residence time defines the total time available for the transfer process to occur within the control volume. A numerical value of the following means that only a small amount of the conserved property would be transferred in the limited time available for the specific process to evolve:

U, < < 1 (1-10)

As a result, the specific process would not be important to the overall transient.

Numerical values of the following means that the specific process evolves at a high enough rate to pennit significant amounts of the conserved property to be transferred during the time period (tcy):

Ili > 1 (1-11)

Such processes would be import.mt to the overall transient behavior.

O 368Iw 1.non:Ib-060897 1 10

FACILITY SCALING REPORT l'

l l

l-1 Scaling Criteria Development Scaling criteria are developed by requiring that the characteristic time ratios for a subset of specific processes in the prototype (usually those of greatest importance) are matched in the model at each j L hierarchical level. Hat is: )

l Tp], = Tp]p M  !

' These criteria are satisfied by adjusting the physical geometry, fluid properties, and operating

. conditions of the model; thus optimizing the model design for the specific process of interest. )

I Evaluation of Scale Distortion The effect of a distortion in the model for a specific process can be quantified as follows:

DF = f)P -f)m (1-13) b)p The distortion factor (DF) physically represents the fractional difference in the amount of conserved I property transferred through the evolution of a specific process in the prototype to the amount of l conserved property transferred through the same process in the model during their respective residence l 1 . times. ' A distortion factor of zero indicates that the model ideally simulates the specific process. A

{

distortion factor of +0.05 indicates that the specific process in the model transfers five percent less of  ;

the conserved property (on a scaled basis) than the same process in the prototype. The distortion l factor can also be written as follows:

DF = 1_ -(to;)R C' R or DF = 1 -%)R L

The degree to which a specific uor.2, process could impact a particular class of transients can be

' determined by comparing the maximum characteristic time ratio for each of the transfer processes that

arise during the transient.

' 'For the AP600 modes of passive safety system operation, all of the transfer processes that impact the l ' system and subsystem transient behavior exhibit one of two properties which enables the analyst to l1 determine the maximum characteristic time ratio for a transfer process using initial conditions. They I, --

are as' follows:

l J

9 3681w 1.non:Ib-060897 ' ]-11

l FACILITY SCALING REPORT l

I

1. The maximum transfer rate for the process (that is, specific frequency) uccurs at the start of each mode of operation and decays with time towards an equilibrium condition. The majority of transfer processes fall into this category. For example, at the onset of the depressurization mode, the system pressure, the break and vent valve mass and energy flow rates, the secondary side and passive heat exchanger heat transfer rates, the core power, the metal stored energy release rate, and the two-phase natural circulation flow rates will be at their maximum values. Similarly, subsequent to depressurization, at the onset of long-term cooling, the decay power and hence the recirculation flow rate will be at their maximum values for this mode of operation. For these transfer processes, the initial conditions for each mode of operation can be used to evaluate the maximum characteristic time ratios.

2. The maximum transfer rate for the process, although reaching a peak value later in the transient, is limited by the initial conditions of the system. All of the passive safety injection processes fall into this category. For the pre-pressurized accumulators, the maximum injection flow rate is limited by the initial charging pressure. For a gravity-driven system, the maximum injection flow rate is limited by the initial liquid level in the safety injection tank. For these transfer processes, the limiting injection flow rates as determined from initial system conditions can be used to esaluate the maximum characteristic time ratios.

The reader is encouraged to examine the details of the H2TS analysis method presented in NUREG/CR 5809.m Further insights will be gained as the method is applied to each of the modes of operation of interest to the AP600.

1.4 Rationale for Scaling Choices The H2TS methodology described in the previous section enables the analyst to develop a set of scaling criteria that can be used to determine the dimensions and operating conditions of a reduced-scale test facility. Because these scaling criteria are expressed in terms of ratios of model to prototype fluid properties, material properties, and geometrical properties, the analyst must make some choices in each of these areas to achieve closure in the design process. This section presents the rationale for the j scaling choices implemented in the design of the APEX test facility.

{

1.4.1 Selection of the Working Fluid The working fluid selected for use in the APEX facility is water. The reasons for this selection are as follows:

1 l

1. Because water is the working fluid in the AP600, fluid property similarity is achieved whenever l

the system pressures match. This greatly simplifies the scaling process for phenomena that occur j at low pressure, such as IRWST injection and long-term cooling. i O

368Iw-1.non:Ib-060897 1 12

FACILITY SCALING REPORT I

/3 2. By using water, the property routines used in the benchmark codes need not be modified. Thus, l U the benchmark codes can be used directly to calculate phenomena observed in APEX.

i

3. Water is the most economical choice. ,

1.4.2 Selection of the Component Materials Stainless steel has been selected as the construction material for the APEX reactor coolant system ,

(RCS) and passive safety systems. The primary reasons for this selection are as follows: l

1. Stainless steel is the primary construction material for the AP600 reactor coolant system (RCS) and passive safety systems. Therefore, material property similarity is preserved and the scaling process simplified.

l

2. Although stainless steel is more expensive than carbon steel, it has greater corrosion resistance.

This reduces the potential of blocking DP taps and extends the life of the facility, making it a l more economical choice over the long term.

l l

Other materials are implemented in regions where excessive stored energy might create significant distortions to system behavior. For example, the reflectors in the APEX are filled with a low mass

[]

Q/

ceramic to reduce stored energy in the core.  ;

1.4.3 Selection of the Operating Pressure The maximum operating pressure for APEX is 400 psia. The primary reasons for this choice are as l follows:

1

1. The primary function of APEX shall be to investigate the low-pressure passive safety injection and the long-term recirculation cooling process which occurs near atmospheric pressure.

! 2. Vessel wall thickness requirements increase with system pressure. Similarly, the saturation temperature increases with system pressure. A reduced-scale facility operating at prototypical j l pressure would have a disproportionately large metal mass and, therefore, excessive stored energy.

This would affect the depressurization rate and the transition to long-term cooling.

1.4.4 Selection of the Length, Diameter, and Time Scale ,

l The length ratio for APEX is 1/4. This length scaling ratio is applied to all the piping lengths and component elevations, except in specific components where other more important criteria take precedence.

368Iw.1.non:Ib-060897 1-13

FACILITY SCALING REPORT For natural circulation processes (gravity-dominated), the time scale corresponding with the 1/4 length scale requirement is 1/2. Therefore, to maintain proper event sequence timing, all of the transport processes shall be scaled to satisfy the 1/2 time scale requirement.

The diameter scaling ratio for all vertical components and piping is 0.1443. The diameter scaling ratio for all horizontal or inclined piping containing two-phase fluid is 0.1612. The reasons for these selections are as follows:

1. Using these length and diameter scaling ratios, the fluid volume requirements are not excessive (for example, IRWST volume is approximately 3000 gallons).

2. Core power requirements associated with fluid volume are also reasonable (2 percent decay power is approximately 400 kW).

3. The time scaling ratio of 1/2 makes long-term cooling test durations reasonable.

4. The component length to diameter ratio (IJd) indicates that multidimensional flow effects will scale well when fluid property similitude exists (Ud = 1.73).

5. The 1/4 length scaling ratio assures that there is sufficient elevation difference between the hot and cold legs to permit accurate measurements of differential pressure. '

6. The diameter ratio for vertical piping and components assures that the skin friction pressure drop in the model will not exceed the combined skin friction and form pressure drop in the AP600 (on a scaled basis) for the case of fluid property similarity. Excessive line resistance would distort system loop flow behavior.

7. The diameter ratio for horizontal and inclined lines expected to contain two-phase fluid is based on flooding theory to scale countercurrent flow processes and two-phase flow regime transitions. j

8. The length and diameter scale requirements can be easily met with commercially asallable pipe and drawn tubing.

9. Construction and material costs are reasonable for the geometric scales that have been chosen. l Minimum Diameter Ratio for Frictional Losses It is de, sired that the skin friction pressu. Jrop in APEX not exceed the combined skin friction and fcan pressure drop in the AP600 on a ses ed basis. Of particular interest is the natural circulation process when fluid property similitude exists.

O' 368Iw.1.non:Ib-060897 1 14

l i

FACILITY SCAltNG REPORT I

/~'N Natural circulation flow is governed by the balance between frictional resistance forces and buoyancy V) i forces. Buoyancy forces depend on fluid density differences and the elevation difference between the heat source and the heat sink. The frictional resistance depends on skin friction and form losses. In general, the form loss can be controlled in the test facility through the use of orifices. Skin friction is l controlled by careful selection of pipe diameter and pipe wall material. j 1

It is important that the pipe diameters in the test facility not be so small that the pressure drop in the lines exceeds the desired scaled pressure drop. The following analysis presents a simple method of selecting a minimum pipe diameter ratio.

For steady-state natural circulation flow, the buoyancy-friction balance equation for a section of vertical pipe having a uniform cross-sectional diameter (d) and a length (f) is given by the followirg:

opgf= P" A +K (1-16) 2 d The limiting case would be to assume that the form loss coefficient (K) equals zero. Thus the skin friction factor (f) dominates in this particular line. Therefore, Equation 1-16 would be written as follows:

N Apgf = P" II (1-17) 2 d l This equation can be written in terms of a scaling ratio as follows:

i (Ap)gfg =pg u[ '_gg*

(1-18)

,d,R In general, this equation reflects the fact that the pressure drop ratio should equal the length scaling ratio for natural circulation systems. That is:

fR=AP g (W)

Assuming fluid property similarity in Equation 1-18 yields the following:

dR =ugg f @M In this equation, u g is either the single-phase or two-phase fluid velocity ratio. Similarly,gf is either (G

s I the single-phase or two-phase friction factor ratio. Evaluating the friction factor ratio requires the use v of a Moody diagram. The friction factor is typically a function of the material relative roughness, e/d, 3681w.1.non:1b-060897 ].]$

FACILITY SCALING REPORT and the Reynolds number. For the APEX test facility, the internal pipe surfaces will be polished such that the absolute roughness is equivalent to drawn tubing (that is, e = 5.0 x 104 ft) as opposed to commercial steel (e = 1.5 x 104 ft). This will permit reduced pipe diameters.

By coupling Equation 1-20 with the scaling ratios presented in Table 4-1, the time, velocity, power, friction factor, diameter, and volume scaling ratios can be obtained by specifying a length scale.

Unfortunately, the process is iterative.

For example, by specifying a length scaling ratio (fg) of 0.25, a velocity scaling ratio of ua = 0.5 can be obtained using Table 4-1. The velocity ratio is substituted into Equation 1-20 which is solved iteratively for the diameter ratio becauseg f is a function of both gu and dg .

Figure 1-4 presents the friction factor ratio for a 1/4 length test facility as a function of single- or two-phase Reynolds number. Figure 1-5 presents the diameter ratios required to satisfy Equation 1-20 in a 1/4 length scale test facility.

It is noted that for the 1/4 length scale test facility, typical single- and two-phase Reynolds numbers will range from 8 x 10 4to 3 x 10 . 6The relative roughness typically ranges from 1 x 10 5 to 6 x 10 5, Figure 1-5 indicates that the diameter ratio becomes independent of Reynolds number for Reynolds 7

numbers greater than 1 x 10. For the 1/4 length scale model, the minimum diameter ratio was conservatively selected at e/d = 1 x 10-5 and Re = 1 x 107 The minimum diameter ratio was found to be 0.139.

dmin,g =0.139 (1:7.19) (1-21)

Selecting a diameter ratio less than 0.139 would likely result in model piping that was too restrictive, thus causing distorted flow behavior in natural circulation systems. For the OSU-AP600 model, a minimum diameter ratio of 0.1443 was selected, because it satisfied the criteria while matching many commercial pipe sizes. Thus:

APEX dmin.R = 0.1443 (1:6.913) (1-22)

The process described above was performed for a full range of length scales. The results are presented in Figure 1-6. This figure demonstrates that as length scale increases, all of the other scaling ratios must also increase. For example, going from a length scale of 0.25 to 0.5 requires the volume ratio to increase by a factor of 9 and the power ratio to increase by a factor of 5.

Going to full height would require a minimum diameter scaling ratio of 0.55 (uring polished internal surfaces) to properly scale the skin friction pressure drop effects in piping. However, for systems whose loop pressure drop is dominated by form losses, smaller values of loss coefficient K can be used to compem ate for excessive frictional losses.

3681w.ianon:ib-060897 1 16

FACILITY SCALING REPORT

.1,4.5. Summary of Rationale for Scaling Choices Table 1 1 presents a summary of the scaling choices.

-1.5 References l l

1. Title 10, " Energy," Code of Federal Regulations, Part 50.46, Office of Federal Register, National Archives and Records Administration, Available through Superintendent of Documents,

{

U.S. Government Printing Office, Washington, DC 20402,1987.

L

2. Zuber, N., " Appendix D: A Hierarchical, Two-Tiered Scaling Analysis," An Integrated Structure and Scaling Methodologyfor Severe Accident Technicalissue Resolution, U.S. Nuclear Regulatory Commission, Washington, DC 20555, NUREG/CR-5809, November 1991.

3. Shaw, R.A., et al..' Development of a Phenomena identification and Ranking Table (PIRT)for 1 Thermal-Hydraulic Phenomena During a PWR Large-Break LOCA, U.S. Nuclear Regulatory Commission, Washington, DC 20555, NUREG/CR-5074, November 1985.

l4. Hochreiter, L.E., FLECHT SEASET Program Final Report, NUREG/CR-4167, EPRI NP-4112, ,

Prepared for the U.S. Nuclear Regulatory Commission, November 1985.

5. Ishii, M., and Kataoka, " Scaling Criteria for LWRs Under Single-Phase and Two-Phase Natural l Circulation," Proceedings of the Joint NRC/ANS Meeting on Basic Thermal Hydraulic Mechanisms in LWR Analysis, NUREG/CP-0043, Bethesda, MD, September 14-15, 1982.

6. Condie, K.G., T.K. Larson, and C.B. Davis, " Evaluation of Scaling Concepts for Integral System Test Facilities," pp 13-22.

j. . 7. Taitel, Y., and A.E. Dukler, "A Model for Predicting Flow Regime Transitions in Horizontal and l

Near Horizontal Gas-Liquid Flow." AIChE lournal, Vol. 22, No.1, pp 47-54, January 1976.

8. Schwartzbeck, R.K., and C. Kocamustafaogullari, " Similarity Requirements for Two-Phase Flow l Pattem Transitions," Nuclear Engineering and Design,116, pp 135-147,1989.

9. Kocamustafaogullari, G., and M. Ishii, " Scaling Criteria for Two-Phase Flow Natural and Forced

. Convection Loop and Their Application to Conceptual 2x4 Simulation Loop Design," Argonne

' National Laboratory, ANL-83-61 NUREG/CR-3420, May 1983.

U 10. Moskal, Thomas E., " Examination of Scaling Criteria for Nuclear Reactor Thermal-hydraulic Test g -  : Facilities," Ph.D. Dissertation, Carnegie-Mellon University, Pittsburgh, PA, March 1987.

11.- Zvirin, Yoram, and Jean-Pierre Sursock, " Scaling of Natural Circulation Experiments," pp. 31-39.

36s tw.1.non:ib.060897 1 17 L_______i..___.__1______..__.__ _ . _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . . . _ _ _ _ _ _ . _ _ _ _ _ _ _ _ . _ _ _ . _ _ ._ _.

FACILITY SCALING REPORT

12. Wang, Z., Y.Y. Hsu, K. Almenas, and M. DiMarzo, "On the Applicability of Ishii's Similarity Parameters to Integral System Test Facilities," Nuclear Engineering and Design,117, pp. 317-323, 1989.

i l

l e

I l

l l

l O

3681w 1.non:Ib-060897 }.}g l l

l

FACILITY SCALING REPORT O

'g TABLE l 1 RATIONALE FOR SCALING CHOICES Rationale for Scaling Choices Choices Reasons

. Working Fluid: Water 1. Fluid property similitude at low pressure.

2. Benchmark codes need not be modified.

• Component Material: Stainless Steel 1. Material property similitude.

2. Corrosion resistance.

• Operating Pressure: 400 psia 1. Adequate for studies of IRWST injection and long-term recirculation cooling.

2. Reduced vessel thickness and metal stored energy.

3. Pressure scaling permits adequate simulation of initial conditions.

. Length Scale: 1:4 1. Fluid volume requirements not excessive (e.g.,

IRWST volume -3,000 gals.)

2. Power requirements reasonable (2% decay power

~400 kW).

3. Time scale makes long-term cooling test duration reasonable (tg = 0.5).

4. Ud ratio indicates that multidimensional flow effects will scale well under fluid property similitude (Ud = 1.73).

[}

\_./

5. Elevation is sufficient such that DP measurements between hot and cold legs are well within instrument capability.

6. Minimum diameter ratio to assure that skin friction pressure drop in APEX does not exceed the combined friction and form losses in the AP600 can be met easily with commercially available pipe and drawn tubing.

7. Two-phase countercurrent flow and flow regime transioons are preserved in horizontal piping.

8. Construction and material costs are reasonable.

• Vertical Pipe Diameter Scale: 1:6.931 l

• Horizontal Pipe Diameter Scale: 1:6.2 (two-phase flow)

. Time Scale: 1:2 l

l b)

L 368Iw.l.non:1b060897 ].[9

FACILITY SCALING REPORT l

Specify experimental objectives O

(1)

(2) PPIRT

"" "" '"E M O# I NI" *

(3) 1 through N Operation Mode Operttion Mode Operation Mode

1. 1 #2 #N WI PIIgroups and (4) (4) (4) similarity criteria 4--

Ye:

(5) T- --

(5)

- - - - Significant scahng distortions? (5)

(6) System design specifications (6) (6)

Evaluates of key T/H PPIRT process to g) prioritize system design specification v

APEX Test Facility design specifications and Q/A Critical Attributes O

Figure 11 General Scaling Methodology 3681w.1.non:Ib-060897 1 20

- S SS  :

s E EI e T C S s s A 4 O Y g e D R LA n c t's l.

t P i

o Ap u r

n/ N lar co p Vor EPA sf t Dg emU- G  : ei d s a n N MIN l Ml st Ag n E Rl y r i OL c E l TA Oal t ap Va R enm F I c

TC Dai R S OS E E B P D i'

• +

s MS d e s EI n y sn s T S a h si e c e s S Y sc c ed e e o s 3 Y L r s r y t S A n pi a n/N o ut o ss r e ms e /a pl pn eNA i

t s r iorl i r mWG an g et H: h nd

ad U- a

v o gt a Sg n Y a mg l EON Er Des a i

t

n c r Ei a r e I

Li F e toi n

E DI l Bl I

l Vaam pb I

- L T PA Vno uq I chi Aac Nmo t a oc O C e R Sct T S EI t b s OC R E D

S D TS P E I N s s s e IO n n m s 2 R* o e T O i o i t EA t i t m F, a t a e i nLC Y r r c t

eAI H t

n t n s mC F I EC e n e d e s e

eST DR c ac c n i V A l I N o en s o E e r E O R E c r o Ac R P D RI I

PH

+

N  : s cn i

IO io t i T st s t MI y ia r res nES h et r t s ie eTO  : mcr Yc

t n e s mSP e YM E

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cmce l

E S O C

Ws yi e T a Nn oer cgp E

O Sh R

EC D

D I

! 1

• 292 h 2$CB32TsaE4 E pe4{ b g- SS gu .-

9 oN$$*

• 2 ? siwI=" -

- 3

,I

FACILITY SCALING REPORT f

1 O i

SYSTEM (S)

S SUBSYSTEM (SS) 881 SSk MODULES (M) M1 M2 --

CONSTITUENTS (C) C1 Ck __ h PH ASES (P) 8 / 8 SYNTHESIS ANALYSIS h

GEOMETRICAL 61 02 Ok CONFIGURATIONS (G)

"IELDS (F) M E MM PROCESSES 4 P2 PIc O

Figure 1-3 Decomposition Paradigm and Hierarchy (NUREG/CR 5809) 3681w-1.non:lb-060897 1-22

FACILITY SCALING REPORT 5

U FRICTION FACTOR RATIO vs. REYNOLDS #

(Fluid Property Similitude) 1.2j j

1 o

g..... . . . . . . . . .

. . . .g REYNOLDS NUMBER (MODEL) l Relabwe Roughness (Mode 0 l

-=- ee = .oooot -+- eM = .oooo2 -*- eM = .oooos

-e- ee = .00004 -*- ee = .0000s se a 00006 l

l Figure 1-4 Friction Factor Ratio (t,/f ) pas a Function of Single. Phase or Two. Phase I Reynolds Number for a 1/4 Length Scale Natural Circulation System with Fluid Property Similitude j DIAMETER RATIO vs. REYNOLDS #

o ,. (Fluid Property Similitude)

I o.3:A _ _ _ . _

o.2s

\ \ -- -

' -~~~~~

0.15

~NNV - m ,m .

g . . . .. g. g . . . . . . . .

. . . .f REYNOLDS NUMBER (MODEL) l Relatve Roughness (Mode 0

-e- Wd = .oooot : ee =.oooo2 --- ee = .00003

-e- ee .00004 -*- ee = .0000s ee =.0000s Figure 15 Diameter Ratios Required to Satisfy Equation 120 O

3681w.l.non:Ib 060897 1 23

FACILTIT SCALING REPORT O

SCALING RATIOS vs. LENGTH SCALE (Fluid Property similitude)

._. .. ,... .. . .. . .. ... ... ... . ..,. . .w_

1.. . . . ... . . ....%..._........... . . .

. . . . =._...... .. .. ........... .....

DIAMEER 4... :.: w nME m . . . . . . _ . . .

0.1. .

POWER e ... - . . _ . .... _

E a

VOLUME

~~-

0.01- :n*

_(j.

-[*""*-

--".=".==.n

=.:: 00 . . _**""

t. .

4 _.

0.001 . . . . .

0.8 0.9 0.2 0.3 0.4 0.5 0.6 0.7 1 LENGTH SCAT.E i

l Figure 16 Sceling Ratio Variation as a Function of Length Scale [The Diameter Ratio Represents the Minimum Required to Satisfy Equation (1.20)]

3681w.l.non:lb.071897 jy I

,l

l I

FACILITY SCALING REPORT 2.0 EXPERIMENTAL OBJECTIVES AND GENERAL SCALING METHODOLOGY The general objective of this research program is to design, construct, and operate a test facility that l will simulate the gravity-driven injection and natural convection, passive cooling modes of the AP600.

From a practical standpoint, pnority has been given to testing those modes of AP600 operation l

previously unexamined for standard PWRs. Thus, automatic depressurization system (ADS) operation, core makeup tank (CMT) injection, accumulator (ACC) tank injection, in-containment refueling water storage tar.k (IRWST) injection, passive residual heat removal (PRHR) operation, non-safety system (NSS) residual heat removal, and lower containment sump (LCS) recirculation will be tested separately and in the context of loss-of-coolant accidents (LOCAs).

l Four types of tests will be performed in the OSU AP600 test facility:

l

• Single-Phase Natural Circulation l

• Steady-State Small Break Loss-of-Coolant Accident (SBLOCA) Two-Phase Natural Circulation

= Transient SBLOCA Cold Leg Breaks (q

j Hot Leg Breaks Inadvertent ADS Operation Cold Leg /CMT Balance Line Breaks Direct Vessel Injection (DVI) Line Breaks e

Low-Pressure Long-Term Cooling In addition to examining the effectiveness of the passive safety systems in the context of LOCA transients, the AP600 tests could also be used to investigate the effects on core cooling due to changes in specific design features, such as ADS valve size. The specific program objectives are as follows:

To validate computational methods used to analyze AP600 passive safety system behavior.

Design, construct, and operate a scale model that provides valid thermal-hydraulic code validation for the AP600 long-term cooling mode.

=

Measure single-phase flows, pressure drops, and temperatures in loop flow paths and in the simulated reactor vessel in order to obtain a mass and energy balance on the system. i

= Provide valid thermal-hydraulic data on the core flow behavior on a scaled basis for each of the different injection modes: CMT, ACC, IRWST, and the lower containment sump return.

'\ 1 368Iw.2.non:1b.o60897 2-1

FACILITY SCALING REPORT

• Provide data on the interfacing effect from the CVS makeup pump and non-safety RHR pump on long-term cooling.

• Provide a basis to scale the test result to higher pressure core cooling transients.

j

= Provide the capability to simulate the injection of nitrogen gas from the passive safety system (PXS) accumulators in order to assess the effect of non-condensible gas on PXS injection and ADS venting capabilities.

l

• Investigate the operation of the PRHR.

l

• Investigate integral system behavior, particularly safety system interactions during SBLOCAs.

l 2.1 Test Facility General Modes of Operation To meet the experimental objectives that have been outlined, four general modes of AP600 operation must be simulated in the test facility. They are as follows:

• Closed system natural circulation (single-phase and two-phase)

• Open system depressurization a Venting, draining, and injection (CMT, ACC, IRWST, ADS)

• Recirculation (PRHR, LCS)

Thus, the general scaling methodology for the AP600 test facili'y can be presented schematically as shown in Figure 2-1.

It is recognized that during a LOCA event, these modes of operation are interrelated. In order to properly scale these modes of operation in the context of LOCA transients, the thermal-hydraulic phenomena important to LOCA behavior must be identified. This is done in Section 3.0 using Plausible Phenomena Identification Ranking Tables (PPIRTs).

The PPIRTs will identify the key thermal-hydraulic phenomena that should be preserved for the diffen nt modes of operation for the facility, such as single-phase natural circulation, two-phase natural circulation, depressurization, and long-term cooling. While experience does exist on the first three modes of operation, the long-term cooling mode is a new area of study and is of particular interest for the AP600.

Since the long-term recirculation flow path and conditions are unique to the AP600, the APEX experiments will simulate this post-LOCA period for several different break and single failure configurations. The different modes of operation are discussed in detail in Section 3.0.

3681w.2.non:Ib-o60897 2-2 l

FACILITY SCALING REPORT

'f 2.2 Fundamental Scaling Requirements :

As indicated in the previous sections and further defined by the PPIRTs in Section 3.0, four basic modes of operation encompassing numerous processes will be investigated using the APEX facility.

'Ihe number of constraints applied to the APEX facility design are defined by the number of hierarchical levels and the types of phenomena to be examined using the facility. For example, if one were solely interested in system depressurization, a scaling analysis performed at the system level

. could be used to develop a simple test configuration capable of simulating the salient features of.

AP600 depressurization. However, the details of processes occurring at the subsystem or constitutive levels would not be simulated. The greater the number of questions posed of a single test facility, the greater the number of constraints that must be applied to its design.

It is desired to develop a facility with extensive simulation capability to address the large number of phenomena identified in Section 3.0 to permit the evaluation of unexpected phenomena that might arise during the process of testing. This requires performing scaling analyses at the system, subsystem, and in some cases, the constitutive levels for the different modes of operation. Although such an extensive analysis requires developing and evaluating a significant number of characteristic time ratios, the strength of the H2TS methodology lies in reducing this number to the minimum subset needed to address only the phenomena of importance. This gives the designer enough flexibility to make a monumental task manageable.

i. (_./

-- Priority is given to scaling the gravity-driven injection and natural circulation processes at low pressure

' in the AP600 because these functions have not been previously tested. The natural circulation scaling analysis presented in'Section 4.0 will be used to establish the fundamental scaling ratios for time, relative elevation, and volume. These scaling ratios will be applied to all of the processes and components important to each of the four modes of operation. The fundamental scaling requirements can be stated as follows:

l 1. For each of the important processes, the ratio of the residence time in the model to that in the full-l scale prototype will be fixed to a value of one-half. Thus, for the APEX test facility the residence L time scaling ratio for all of the important processes is set as follows:

(Tey), = M M

2. The volume of each of the components important to a mode of operation will be scaled to satisfy the following volume scaling ratio:

VR = 1:192 (2-2)

' 36asw.2mib-osos97 2-3

FACILITY SCALING REPORT

3. The relative elevation of each of the components important to a mode of operation will be scaled to satisfy the following length scaling ratio:

I l

Lg = h4 (2-3 Having all of the important components and processes satisfy these requirements is a necessary condition for proper simulation of event sequence evolutions which involve multiple modes of operation.

2.3 AP600 System Decomposition and Hierarchy For purposes of clarity, the AP600 integral system can be divided into the reactor coolant system (RCS) and the passive safety system (PXS).

Figures 2-2 and 2-3 present the decomposition and hierarchical levels for each of these systems.

These figures serve as a road map to the scaling analyses presented in Sections 4.0 through 8.0. They identify the hierarchical level at which the scaling analysis was performed (that is, the control volume),

the geometrical configuration of interest, the balance equations implemented in the analyses, and the processes to be scaled. ,

Thermal-hydraulic phenomena involving integral reactor coolant system interactions, such as primary system depressurization or loop natural circulation, are examined at the system level. Thermal-hydraulic phenomena, such as PRHR decay heat removal and steam generator heat transfer, are examined at the subsystem level. Specific interactions between the steam-liquid mixture and the stainless steel structure are examined at the constituent level.

2.4 Initial Conditions for Long-Term Cooling The OSU APEX test facility has, as its primary purpose, the task of obtaining data on low-pressure long-term cooling. The initial conditions for long-term recirculation cooling are established through the process of RCS depressurization coupled with CMT, ACC, and IRWST injection. This presents a significant challenge to the APEX facility, whose initial blowdown pressure of 400 psia is significantly lower than the 2100 psia specified for the AP600.

Two approaches can be used to establish initial conditions for long-term cooling in the APEX facility.

The first approach is to maintain pressure similitude. This requires simulating in the APEX an AP600 depressurization transient already in progress. Initial mass and energy inventories need to be determined through calculation. Another approach is to scale the pressure and fluid properties to simulate the depressurization process on a scaled basis followed by a transition to pressure similitude at low pressure. This latter approach is examined in Section 5.0.

3681w-2.non:lb.072597 24 Rev.I

t i FACILITY SCALING REPORT i

! O O

j Specify Experimental (I) Objectives (2) $BLOCA PPIRT i

(3). Perform Scaling Analysis l~

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System Design

(6) Specifications Specifications Specifications Specifications 1 1 1-1 1 _.

Evaluates of key T/H PPIRT (7) Process to Prioritize System Design Specification APEX Test Facility Design l (t) Specifications and Q/A Critical i Attributes i

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Al I ' - Figure 21 General Scaling Methodology for the AP600 Test Facility l

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Figure 2-2 AP600 Reactor Coolant System Decomposition and Illerarchy (Process) 3681w-2.non:Ib-080497 2-6 Rev.I  ! I l

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I n v Figure 2-3 AP600 Passive Safety System Decomposition and Hierarchy . l l 368!w 2.non:lb-080497 27 Rev.1 l

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FACILITY SCALING REPORT O v 3.0 PHENOMENA IDENTIFICATION AND RANKING This section presents the results of the first step in the general scaling methodology (GSM) presented in Figure 2-1. Development of the proper similarity criteria for the design of the AP600 integral system and long-term cooling test facility requires that the important thermal-hydraulic phenomena be identified. - The development of a Plausible Phenomena Identification Ranking Table (PPIRT) is extremely usefulin this respect. Prior to developing a PPIRT, it is necessary to examine the AP600 emergency core cooling (ECC) process. 3.1 AP600 Design and Emergency Core Cooling The AP600 is a light water nuclear reactor system designed to produce 1940 MWt. The general operating conditions of the AP600 have already been described in the literature.(1,2) pigure 3-1 is a schematic that illustrates the primary system components. The primary loop consists of the reactor vessel, which contains the nuclear fuel assemblies; two hot legs, which connect the reactor vessel to the steam generators; two steam generators; a pressurizer; four canned motor pumps; and four cold legs. l p

 ;                                                                            Normal, full-power operation of the AP600 is straightfonvard. Heat is generated in the reactor fuel.

This heat is transported by forced convection to the water. Since the entire system operates at 2250 psia, bulk boiling of the water does not occur. The heated water is transported through the hot legs to the U-tubes inside the steam generators. The energy of the primary coolant inside the tubes is transferred to the water on the secondary side by conduction through the tube walls and convection boiling on the outside surface of the U-tubes. The cooled primary water inside the tubes is pumped by four canned motor pumps through four cold legs and back into the reactor ves*I where the heating cycle repeats. Primary system pressure is maintained constant by the pressurizer. With respect to thermal-hydraulic phenomena, normal full-power operation of the AP600 is typical of most pressurized water reactor (PWR) systems. The unique feature of the AP600 is the method it uses to remove decay heat from the nuclear fuel in the event of an accident Figure 3-2 presents a simple schematic that describes the operation of the passive safety system. This system implements gravity-driven cooling devices which require no operator interaction. In existing PWRs, the post-accident safety injection flow into the reactor vessel downcomer is essentially independent of the back pressure that exists in the reactor coolant system (RCS). The j pumped flow fills the reactor downcomer, and the gravity head in the downcomer drives the flow } through the core and the loops. The mode of operation for the primary system is a two-phase natural l.g) ( v circulation flow. If the pressure drop in the core or system is excessive, then the pumped flow spills out the break, and only the downcomer gravity head acts to force the flow through the core and primary system. l 3681w.3.non:ll>060897 3-1

FACILITY SCALING REPORT Long-tenn cooling for the AP600 design is generally similar to existing plants; however, there are some unique features of the AP600. The safety injection flow from the IRWST or sump is dependent on the pressure in the reactor coolant system (RCS), since this flow is driven by the elevational heads in either the in-containment refueling water storage tank (IRWST) or sump. Therefore, the reactor coolant system (RCS) would pressurize to a pressure greater than the containment pressure plus the gravity head in either the sump or IRWST; the check valves in the injection lines would close; and the safety injection would diminish. 'Ihis does not happen in a pumped safety injection system. However, to prevent any pressurization of the reactor coolant system (RCS), the fourth-stage automatic depressurization system (ADS) opens several large flow area valves on the reactor hot legs to vent steam and two-phase mixture such that over pressurization is precluded. The fourth-stage ADS also serves another purpose. By creating a large break on the hot leg side of the reactor vessel, before the steam generators, the recirculation path will be from the downcomer, through the core, out the hot legs, and out the ADS 4 vent valves-instead of having to flow through a hot steam generator, which l creates a large two-phase pressure drop as in existing PWRs. Therefore, the venting capability of the ! AP600 is enhanced relative to existing plants such that the gravity-driven flow from the IRWST and sump provides adequate core cooling. The passive safety system includes a 530,000 gallon in-containment refueling water storage tank (IRWST), two 2000 ft 3core makeup tanks (CMTs), and two accumulators containing 1700 3ft of I water and 300 ft3of nitrogen pressurized to 700 psig. The water supply is borated to provide reactivity control in addition to core cooling. The passive safety system includes two tiers of three sets of valves located on the pressurizer and two independent fourth-stage valves located on the hot l legs. These valves compdse the automatic depressurization system (ADS). Each set of valves is staged to provide a controlled depressurization of the primary system. 3.1.1 AP600 Large Break Loss-of Coolant Accident i This section describes how the passive safety system is expected to operate during a large break loss-j of-coolant accident (LBLOCA). The LBLOCA provides the system initial conditions necessary for long-term cooling simulation. Westinghouse has performed a series of calculations to predict LOCA behavior in the AP600 as described in the AP600 Standard Safety Analysis Report (SSAR).W It was found that most of the LBLOCA thermal-hydraulic phenomena of interest to the AP600 is typical of that expected for standard light water PWRs. A detailed description of the sequence of events for a standard PWR is provided in NUREG/CR-1230 W and NUREG-5074,W and is specifically discussed for the AP600 in Reference 3. Westinghouse performed a calculation of the worst case LBLOCA for the AP600 design.W They found that the double-ended guillotine break of the cold leg was the most severe. As with standard PWRs, the AP600 LBLOCA sequence includes a blowdown period, a lower plenum refill period, and a core reflood period. Figure 3-3 illustrates the different periods of an AP600 LBLOCA. It provides a sample plot of the peak fuel cladding temperature as a function of time for the duration of the LBLOCA. l 368Iw.3.non:Ib-060897 3-2 l 1

i l J FACILITY SCALING REPORT For the AP600 these three periods are defined as follows: Blowdown: This period begins with the break initiation and ends when the accumulator injection initiates in the downcomer; a period of approximately 35 seconds. Lower Plenum Refill: This period begins with the accumulator injection and ends when the accumulator mixture level in the lower plenum reaches the core inlet; a period of approximately 15 seconds. Core Reflood: This period begins when the liquid mass in the core starts to increase and ends when the whole core is quenched and submerged again. For the average fuel rod in the AP600, this period is approximately 200 seconds. The primary source of cooling during this period is from the accumulators, which have lower injection flow rates to prolong the injection periodi q

During these periods, the AP600 LBLOCA exhibited the same behavior as the standard PWR with the following exceptions:

l

                     * - Blowdown -

i

1. Because the cold leg is smaller in the AP600 than in the standard PWR, the break size is smaller, and therefom, depressurization of the system is longer.

2. The AP600 canned motor pumps coast down faster than the standard reactor coolant pumps (RCPs). Therefore, the negative core flow rates are greater, which increased fuel rod cooling during blowdown.' The flow path from the upper head to the upper plenum

                                                   . aids the downflow and permits the head to completely drain.

Lower Plenum Refill - The accumulators are main sources of water during refill. Accumulator coolant is injected . directly into the downcomer. Thus, the amount of accumulator water bypassed out of the break is reduced. Core Reflood I. Accumulator coolant injection is the main source of water during reflood. I i

2. ; Because of the additional fuel cooling during blowdown, the peak clad temperature is f, more likely to occur during reflood.

3 368is.t non:1b-060897 3-3

FACILITY SCALING REPORT  !

                               -                                                                                                                                                                                                                 i

3. The AP600 operates at a lower linear power so more core quench occurs during blowdown, and the fuel heats up slower during reflood than in the standard PWR. ,

4. Cooler fuel at the end of blowdown results in higher reflood rates.  ;

5. The accumulators are designed to continue injection up to 120 seconds for a large break, as compared to 45 to 50 seconds for existing PWRs.

l The entire process of blowdown, lower plenum refill, and core reflood occurs in less than four minutes. The ADS does not activate within this time frame because, up to this point, CMT injection has been limited. The accumulator flow, which can last up to 120 seconds, prevents CMT injection. After the accumulators are empty, the CMTs begin to inject and continue until they empty. At this time, injection begins from the IRWST. Because of the large break, system pressure rapidly approaches containment pressure. This permits IRWST gravity-driven injection. Unlike a standard PWR, subsequent to core reflood in the AP600, the CMTs and the IRWST provide core cooling on a long-term basis by gravity-driven injection and natural circulation heat transfer. Therefore, in addition to the three periods just described, an IRWST injection cooling period and a lower containment sump (LCS) recirculation cooling period is also considered. For the AP600, these two periods are defined as follows:

• IRWST Injection Cooling: This period begins when the downcomer pressure is low enough to O

permit gravity-driven injection from the IRWST and ends when positive flow is established from the containment sump to the reactor vessel. During this per.iod, the lower containment compartments are filling with water up to the design flood-up elevation.

• LCS Recirculation Cooling: This period begins when positive flow is established from the containment sump to the reactor vessel and continues until all core decay heat is removed.

Section 3.2 identifies the specific thermal-hydraulic phenomena that occur during the five periods of an AP600 LBLOCA. 3.1.2 AP600 Small Break Loss-of-Coolant Accident This section describes how the passive safety system is expected to operate during a small break loss-of-coolant accident (SBLOCA). Westinghouse presents a series of SBLOCA calculations to examine the effectiveness of the AP600 passive safety system in the AP600 Standard Safety Analysis Report (SSAR).(3) The initial portion of the SBLOCA is typical of that expected for standard light water PWRs. However, subsequent to ADS initiation, the AP600 SBLOCA scenario differs from what i might be expected in the standard PWR. 3681w.3.non:1b-060897 3-4

l FACILITY SCALING REPORT ,O An SBLOCA in an AP600 can be divided into five periods. These periods are illustrated in U Figure 3-4, which is a plot of system pressure as a function of time during an AP600 SBLOCA. The onset and duration of each period will vary depending on the break size. l For the AP600, these five periods are defined as follows: Blowdown: This period begins with the occurrence of the break and ends when the system pressure reaches a quasi-equilibrium state at the steam generator safety valve setpoint. Saturation Natural Circulation: This period begins subsequent to reactor coolant pump coastdown. For larger breaks, on the order of eight inches, this period may be very short or bypassed completely. This period ends with the activation of the ADS.

                                                    . ADS Operation: This period begins with the opening of the ADS 1 valve and ends when the system pressure reaches the IRWST injection pressure.

l l IRWST Injection Cooling: This period begins when the downcomer pressure is low enough to permit gravity driven injection from the IRWST and ends when positive flow is established from the containment sump to the reactor vessel. During this period, the lower containment compartments are filling with water to the design flood-up elevation. t 1 LCS Recirculation Cooling: This period begins when positive flow is established from the containment sump to the reactor vessel and continues until all core decay heat is removed. The first two periods of the AP600 SBLOCA are similar to those in standard PWRs. There are two major differences in overall SBLOCA behavior in the standard PWR and the AP600:

1. The AP600 does not have pump loop seals. Hence, the core uncovery caused by clearing of the loop seals does not occur in the AP600.

2. Operation of the ADS changes the SBLOCA scenario from that in the standard PWR. Rapid depressurization permits all of the high capacity injection systems to operate relatively early in the scenario.

Section 3.2.2 identifies the specific thermal-hydraulic phenomena that occur during the five periods of an AP600 SBLOCA. 3.2 AP600 Plausible Phenomena Identification Ranking Table  ! i

                                                                                                                                                                                                              }

This section is aimed at identifying the important thermal-hydraulic phenomena that would occur  : during LBLOCAs and SBLOCAs in the AP600 system. The development of a comprehensive LBLOCA or SBLOCA PPIRT requires significant effort. It is possible that the initial test program l 368Iw-3.non:ll>.060897 3-5 I

FACILITY SCALING REPORT may reveal additional phenomena of importance. Therefore, rather than instituting a formal AP600 PPIRT development process at this time, the methodology presented in Figure 3-5 will be used to develop a set of Plausible Phenomena Identification Ranking Tables (PPIRTs) for AP600 LOCAs. Because the AP600 is a light water PWR, advantage can be taken of the extensive LOCA research performed over the past twenty years. Much of this research is summarized in NUREG-1230.W The imponant thermal-hydraulic phenomena for standard PWR LDLOCAs and SBLOCAs have already been identified and ranked in order of imponance.(5,6) Therefore, a wealth of information is available for the development of the AP600 PPIRTs. The methodology presented in Figure 3-5 recognizes the fact that the existing PIRTs were developed for PWRs having somewhat different geometries and operating conditions. By comparing AP600 1 LOCA calculations with the existing PIRTs and applying engineering judgement, it is possible to determine if differences between standard PWR and AP600 geometry or operating conditions are important with respect to the occurrence of certain thermal-hydraulic phenomena. If the phenomena is found to be applicable to the AP600, it is incorporated into the AP600 LOCA PPIRT along with the expert's ranking on the importance of the phenomena. The PPIRT development methodology also requires that for the purposes of test facility scaling a ranking of high importance be assigned to phenomena associated with all modes of AP600 operation not previously tested by experiment. This is done for two reasons:

1. To require the development of an experimental data base for code assessment and future PIRT development for similar systems.

2. To assure that unexpected thermal-hydraulic phenomena of importance are identified through the process of experimentation.

By applying this methodology, a set of AP600 LOCA PPIRTs can be developed. The AP600 LOCA PPIRTs will be conservative with respect to previously untested modes of operation. However, upon completion of testing, these PPIRTs can be revised to include the new findings and support the formal PPIRT development process. Section 3.2.1 presents the AP600 LBLOCA PPIRT, and Section 3.2.2 presents the AP600 SBLOCA PPIRT. 3.2.1 AP600 Large Break Loss-of Coolant Accident Plausible Phenomena Identification Ranking Table A'though the proposed research focuses on SBLOCA and long-term cooling experirnents, this section has been included as a matter of completeness and because it provides an additional comparison between the AP600 and standard PWRs. 3681w 3.non:Ib.060897 3-6

FACILITY SCALING REPORT O A large break LOCA PIRT for standard PWRs has been developed by a team of twenty-one experts U with extensive backgrounds in nuclear system thermal hydraulics.(5) The LBLOCA PIRT identifies and ranks the importance (relative to peak fuel cladding temperature) of the thermal-hydraulic phenomena that occurs during a LBLOCA in a standard PWR. These results were independently confirmed using the analytic hierarchy process (AHP) described in NUREG/CR 5074. Therefore, as stated in that report, "the LBLOCA PIRT can be used for any task requiring a knowledge of LBLOCA phenomena."(5) A comparison of the thermal-hydraulic behavior resulting from a LBLOCA in an AP600 to that produced in a standard PWR is presented in Section 3.1.1 of this repon. This comparison, coupled with engineering judgement, indicates that the AP600 LBLOCA would exhibit the same types of thermal-hydraulic phenomena as the standard PWR LBLOCA. For the LBLOCA, the major

              .,                    differences in thermal-hydraulic beha'vior would occur subsequent to core reflood. This would involve IRWST and CMT injection and lower containment compartment (LCC) recirculation. Therefore, with respect to AP600 blowdown, lower plenum refill and core reflood, the existing PWR LBLOCA PIRT is directly applicable with minor modifications. The thermal-hydraulic phenomena for these three -

periods, along with the expert ranking, have been incorporated into the AP600 LBLOCA PIRT _ pr:sented in Table 3-1, The thermal-hydraulic phenomena associated with the IRWST injection cooling period and the LCS recirculation cooling period are also presented in Table 3-1. h Table 3-1 separates the LBLOCA into five periods. "Ihe left-hand column lists all of the plant components of importance. Beneath each component heading is listed the set of thermal-hydraulic phenomena that involve that specific component. Each thermal-hydraulic phenomenon is ranked as follows: (H) A high rank indicates that the phenomenon in question significantly impacts the peak fuel cladding temperature during the LBLOCA. (M) A medium rank indicates that the phenomenon in question moderately impacts the peak - fuel cladding temperature during the LBLOCA. (L) A low rank indicates that the phenomenon in question has little impact on the peak fuel cladding temperature during a LBLOCA. (P)- This rank indicates that it is plausible that the phenomenon in question may impact the peak fuel cladding temperature during a LBLOCA. However, an expert panel ranking has

                                               ,   not been performed. For purposes of facility scaling, high importance has been assigned to this phenomenon.

(-) This rank indicates that the phenomenon is either negligible or does not occur during the O g period being considered. I 1  ; 3681w.3.non:1b-060897 '3-7 l l \ t i

FACILITY SCALING REPORT The thennal-hydraulic phenomena for the first three periods of the LBLOCA have been described in NUREG/CR-5074 and NUREG-1230 for standard PWRs. The remaining thermal-hydraulic phenomena, occurring during the IRWST injection cooling and LCS recirculation cooling periods, are specific to the AP600. Because of ADS operation during the SBLOCA, the thermal-hydraulic phenomena of importance during the upper plenum refill period and the LCS recirculation are similar j for the LBLOCA and the SBLOCA. This is particularly tme for LCS recirculation. Thus to avoid l repetition, and in view of the experimental objectives which focus on the investigation of SBLOCA transients and long-term cooling, a detailed description of the thermal-hydraulic phenomena of interest during these tw - 'ods is presented in the next Section 3.2.2 in the context of SBLOCAs. 3.2.2 AP600 Small Break Loss-of Coolant Accident Plausible Phenomena Initial Ranking Table Section 2.0 presented the experimental objectives for the integral system test facility. SBLOCA and long-term cooling tests will be the primary emphasis of the experimental investigation. Therefore, the PPIRT developed in this section serves to identify and rank the thermal-hydraulic phenomena of interest. The scaling analyses presented in the following sections address the phenomena identified by the SBLOCA PPIRT. A SBLOCA PPIRT for a Babcock and Wilcox (B&W) PWR has been developed by a team of fifteen experts. The two safety criteria examined by the group were the following:

• Peak fuel cladding temperature O'

a Core liquid inventory The SBLOCA phenomena that would significantly impact these two safety criteria were identified and ranked in order of importance. The ranking was assessed by staff from the Idaho National Engineering Laboratory (INEL) using the analytical hierarchy process.W Although the B&W PWR has some design features that are significantly different than the standard Westinghouse PWR, the first two periods of the SBLOCA are very similar. For the AP600 SBLOCA, the major differences in thermal-hydraulic behavior occur upon ADS activation. Therefore, with respect to the AP600 blowdown and saturation natural circulation periods, the existing PWR SBLOCA PPIRT is directly applicable with minor modifications. The thermal-hydraulic phenomena for these two periods, along with the expert ranking have been incorporated into the AP600 SBLOCA PPIRT presented in Table 3-2. Table 3-2 separates the SBLOCA into five periods. The left-hand column lists all of the plant components of importance. Beneath each component heading is listed the set of thermal-hydraulic phenomena that involve that specific component. O 3681w.3.non:1bo60897 3-8

FACILITY SCALING REPORT p Each thermal-hydraulic phenomenon is ranked as follows: d (H) A high rank indicates that the phenomenon in question significantly impacts the peak fuel cladding temperature or core liquid inventory during an SBLOCA. (M) A medium rank indicates that the phenomenon in question moderately impacts the peak fuel cladding temperature or core liquid inventory during the SBLOCA. (L) A low rank indicates that the phenomenon in question has little impact on the peak fuel cladding temperature or core liquid inventory during an SBLOCA. (P) This rank indicates that it is plausible that the phenomenon in question may impact the peak fuel cladding temperature or core liquid inventory during an SBLOCA. Although, an expert panel ranking has not been performed, SSAR small break calculations have been

         ..I performed for a wide range of conditions. For purposes of facility scaling, high importance has been assigned to this phenomenon.

(-) This rank indicates that the phenomenon is either negligible or does not occur during the period being considered. *

l. L Table 3-3 provides a description of the AP600 SBLOCA,and long-term cooling phenomena that have been ranked (H) or (P) in the SBLOCA PPIRT. These thennal-hydraulic phenomena are evaluated further in Sections 4.0 through 7.0 for the purpose of scaling the AP600 integral system and long-term

       ,                      . cooling test facility. In addition to a description of the phenomena, Table 3-3 identifies the components of interest, the periods during which the phenomenon is active, and the section of the

report that addresses the scaling of the phenomena. The property scaled equations include fluid property ratios which have been pressure scaled. Fluid property similitude is assumed for the low-

_ pressure phenomena that would occur during IRWST injection cooling and LCS recirculation cooling. In summary, this section has identified the plausible thermal-hyliraulic phenomena for LBLOCA and SBLOCA events. Because the focus of the proposed experimental effort is SBLOCA and long-term cooling, the SBLOCA PPIRT has been used to guide the scaling analyses presented in the next four sections. l 0J. 1 3681w.3.non:lb 060897 3-9 4 L

FACILITY SCALING REPORT 3.3 References

1. Vertes, C.M., " Passive Safeguards Design Optimization Studies for the Westinghouse AP600,"

Fifth Proceedings of Nuclear Thermal Hydraulics,1989 Winter Meeting of the American Nuclear Society, San Francisco, California, November 26-30,1989.

2. Westinghouse AP600 Plant Description Report, U.S. Department of Energy, Technology Programs in Support of Advanced Light Water Reactors, DE-AC03-86SF16038, Jaisuary 1989.

(Westinghouse Proprietary, Class 2)

3. AP600 Standard Safety Analysis Report (SSAR).

4. Compendium of ECCS Research for Realistic LOCA Analysis, U.S. Nuclear Regulatory Commission, Washington, DC 20555, NUREG-1230, December 1988.

5. Shaw, R. A., et al., Development of a Phenomena identification and Ranking Table (PIRT)for Thermal-Hydraulic Phenomena During a PWR Large-Break LOCA, U.S. Nuclear Regulatory Commission, Washington, DC 20555, NUREG/CR-5074, November 1985.

6. Ortiz, M. G., L.S. Ghan, and G.E. Wilson, " Development of a Phenomena Identification and Ranking Table (PIRT) for Thermal-Hydraulic Phenomena During a PWR Small Break LOCA,"

infonnal report, EGG-EAST-9080, INEL, May 1990. O 3681w-3.non:1b-060897 3-10

o it g Sl a i n H I I Culc o - - / - - - - - I

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s n a a t c qe e d L e m h Tdaidr r e r n v Hveib u P e,iPcaR yr u O o e gf no t t n O n  : e c eh r e n s id a o i d a e T. ri it n C h e ta n e I e s u ah n e hS t Wai tu n sinte o P R ioh e. s iq weh r e s n v C M t wa n al t t o l loi d s ugR pdi in I o ( R e F mn cd t o vi r r o e a 3- E e m o yr h e ui u Fl ee s owr l 3 -T o c le ar h sd t t s l f b t ed e ei q EG nd sf ioiul o S nd pac i t s mnd ois iul inRle n ri LN t q e pi niR yo . r q , BO AL lula i mDa r Ani m d tai n se sndTI ot c l I r c T D c e o ir h f eid o ov c od e eS Rt P ne p a N c t e yntai d gt c u l o n o mn taw c e v A Rb a sf c Nr cle Io oRh n m r t ii A C O d L S eS Ct B S n L aD r s o n tuA a 0 io Sl e d it 0 o a t b 6 i l u c ni s P r e c jieno A P i c r n t e n F e I la d w O R Tu n o S cr od c w N S Wi c n O I C L R RNB 1 e olo T P I R C S t E n e D n o p m R o S H C CL R P

4. 5 2 2 b r4l 3 ,6*

FACILITY SCALING REPORT , O A v 5-ModelF steam k generator s, v latograted w v

                                                                                                          -              y                                 e s

Coldles pipe (22in10) oj .J safety w

                                                              -                                  evection s..,.                           noate
                                                                 'w                          Re3C10f M

10) conneo v motor pumps Figure 31 Diagram of the Westinghouse AP600 System e 368Iw-3.non:Ib 060897 3-26

FACILITV SCALING REPORT o

                                                                /p-j CONT.

g@g " i oEeREss {l

, @Xs '~ i 7REFUR XI VALVES O/2) W g @h 02:

g IRWST CAvlTY Sf* ACCUW. (1 0F'2)

                                                                                       ..l.. pagg CORE WAKEUP p TANK (1/2) g HX
                        'T                                       PRESSURt2ER
                                                                                     '      ' 0/0 (Q                                                                                    3.
                                                                                                              -t +-

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                                                                      ,,                           REACTOR Futt

( $ vEssa CONT. T l O Figure 3-2 Flow Diagram for AP600 Passive Safety System Operation l l 3681w 3.non:lb 060897 3-27

FACILITY SCALING REPORT O l l AP600 LBLOCA SCENARIO LCS Nl BLOWDOWN - l U.P. REFILL l RECIRC.

               -              i                   :                                                                  LP. REFILL                     i
                                               !                                                                                                     i             .

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                              -                                                                                  -                                  j j                                                                       REFLOOC l                     i i                i                                                                                                     i i

i  : i  ! i g g , g g 5 I I TIME Figure 3-3 AP600 Large Break LOCA Scenario l 3681w-3.non:1b-oso897 3-28

FACILITY SCALING REPORT i l l l

   \x l

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Figure 3-4 AP600 Small Break LOCA Scenario l 3681w-3.non:Ib-060897 3 29

FACILTIT SCALLNG REPORT O Westinghouse Existing PWR Engineering AP600 LOCA LOCA Jedgment Calculations PIRTS

                                                         <r Assess Effects of Differences

Between Standard PWR and AP600 Geometry and Operating Conditions Select Phenomena Applicable To AP600 0

For Purposes of Facility Scaling Assign High Importance to Plausible Phenomena Associated with Modes of AP600 Operation Previously Untested 1P Prepare AP600 LOCA PPIRTs Figure 3 5 AP600 LOCA PPIRT Development Methodology 368t w-3.non:Im897 3-30 m

l l l FACILITY SCALING REPORT l 4.0 CLOSED LOOP NATURAL CIRCULATION SCALING V This section develops the similarity criteria that must be satisfied to simulate single-phase and two-phase natural circulation behavior in closed loop geometries. The criteria developed in this section will also be implemented in Section 8.0 for the analysis of lower containment sump (LCS) recirculation. The analysis is divided into two sections. Section 4.1 presents the single-phase natural circulation I scaling analysis, and Section 4.2 presents the two-phase natural circulation scaling analysis. Each of these analyses implements the Hierarchical, Two-Tiered Scaling (H2TS) method to obtain similarity criteria for the primaiy loop (system level) and the core (constituent level). De parameters that have been scaled to address the phenomena of interest to single-phase and two-phase natural circulation are as follows:

• Time scale

                                                . Length scale

• Primary system volume
• Primary-side friction and form losses

                                                . Reactor vessel geometry
  \/O                                                Core flow area e
   "/                                           . Core length

• Core geometty a Core power

                                                . Heater rod diameter a    Heater rod length a    Heater red power
                                                . Ilot leg diameters

• Hot leg length
• Cold leg diameters
• Cold leg lengths

! = Pressurizer surge line diameter

                                                . Pressurizer surge line length

• All component relative elevations l l

4.1 ' Single Phase Natural Circulation Scaling Analysis l l De single-phase liquid natural circulation is the simplest operating mode. The phenomena of interest are as follows: {

 -[\                                            . The stability of the natural circulation mass flow rate for various heat inputs.
                                                =    Single-phase natural circulation heat transfer in the core.                                   I l                                                                                                                                                   l 3681w-4a.non:lb-060897                                41                                                l

FACILITY SCALLNG REPORT Figure 4-1 is a flow diagram for the single-phase natural circulation scaling analysis. First, a top-down analysis is performed. This includes an analysis at the system level (loop) for steady-state conditions and at the constituent level for transient conditions in the core. The objectives of the top-down analysis are to obtain a closed form solution for the steady-state mass flow rate and to obtain a set of characteristic time ratios and specific frequencies for the core transport processes. Following the top-down analysis, a bottom-up/ process analysis is performed to obtain process specific correlations for convective heat transfer and pressure drop in the core. The results of the top-down and bottom-up analyses are then used to develop a set of similarity criteria for single-phase natural circulation. These similarity criteria will be integrated with the two-phase natural circulation similarity criteria presented in Section 4.2. 4.1.1 Top-Down System Level Analysis His section presents an analytical expression for the singleph5 ase natural circulation flow rate (fluid velocity and mass flow rate) for the loop geometry depicted in Figure 4-2. The nomenclature for the equations presented in this section is as follows; ac cross-sectional flow area of core = na,c O a, cross-sectional flow area of i* component a,c cross-sectional flow area of the heated subchannel ! A, total rod heat transfer surface area in core = nA,e A,e surface area of a single fuel rod = ndt Ax total rod solid stnicture cross-sectional area in core = nAxe A xe cross-sectional area of a single rod = nR2 h coefficient of volume expansion Cp liquid specific heat at constant pressure Cy liquid specific heat at constant volume Cy solid specific heat d fuel rod diameter dh hydraulic diameter of the heated subchannel (d h = 4a,c/() 6 conduction depth (6 = A xc /() AT axial fluid temperature difference across the length of the core f Darcy friction factor g gravitational acceleration l H average convective heat transfer coefficient over the heated length k solid thermal conductivity K loss coefficient for the heated subchannel i length of the heated subchannel l 368Iw-4tnon:1b060897 4-2

k 1; FACILITY SCALING REPORT j i l i O Ig elevation difference between the thermal centers's of the core and steam generator 4 - wetted perimeter of the heated subchannel J n . number of rods or subchannels I q, core heat generation rate Q, volumetric flow rate p, liquid density p, solid density l R fuel rod radius l ' T, liquid temperature L <T,> . volume averaged fuel rod temperature l T, heater rod / fuel surface temperature ! T,-T, average rod surface to liquid temperature difference u, liquid velocity V, liquid control volume l

          <$>        volume averaged parameter = 1           $dv V

U gy; average liquid velocity in the DVI

         .Uct        average liquid velocity in the cold leg Uggt average liquid velocity in the balance line

k ci cold leg nozzle loss coefficient k g. cold leg-to-cold leg balance line to loss coefficient k emi, CMT inlet nozzle loss coefficient kcmt , CMT outlet nozzle loss coefficient ~

kgcy . discharge line check valve loss coefficient k ny DVI nozzle loss coefficient L-f frictional length The following subscripts are used: C cold liquid o reference constant H hot liquid s solid i liquid W . wall surface The following simplifying assumptions are made to obtain the mass, momentum, and energy balance equations presented in Table 4-1:

1. Steady-state flow.

2. One-dimensional flow along the loop axis.

O 3. Uniform properties at every cross-section.

l. - 4. Two' constituents: water and metal structure, l

3681w-4a.non:l' b-072197 4-3

FACILITY SCALING REPORT

5. Single-phase liquid.

, 6. Convective acceleration is negligible in momentum balance equation. I

7. Core heat loss is negligible.

8. Viscous dissipation is negligible.

9. Boussinesq approximation is applicable.

As shown in Figure 4-2, the primary loop is divided into two control volumes: a hot fluid side having an average temperaturegT and a cold fluid side having an average temperature c T . By implementing the Boussinesq approximation, all of the fluid densities in the loop are assumed equal except for those comprising the buoyancy term. The entire loop is divided into compor.ents. Properties for each component are denoted by the subscript i. Substituting the energy conservation equation (Equation 4-3) into the momentum equation (Equation 4-2) yields a solution for the steady-state velocity at the core inlet. This is given by Equation 4-4. The steady-state fluid velocity within each component having a cross-sectional flow area, aiis obtained using the mass conservation Equation 41 and is given by Equation 4-5. In Equations 4-1 through 4-5 (found in Table 4-1), F 7 is the total loop friction number given by the following:

                                                                                                                                                           ,       ,,    ,2
                                                                                                                                                                       *                                      (4-6)

F= T +K i-l , dh ,s ai, Equation 4-5 is the standard equation for single-phase natural circulation flow. It predicts the fluid velocity through any of the loop components for different core powers (q,) and loop resistances characterized by F7. 4.1.2 Top-Down Constituent Level Analysis for the Core The transfer processes between two constituents are examined in this analysis: single-phase liquid and the metal fuel / heater rods. Heat transfer to the metal walls and to the ambient is examined in Section 5.0. The H2TS approach described in Section 1.3.3 is implemented. Table 4-2 presents the control volume balance equations (Equations 4-7 through 4-10) for a core consisting of fuel rods arranged in a square array of heated subchannels. Using the procedure outlined in Section 1.3.3, all of these terms are non-dimensionalized by dividing them by their respective initial, or boundary, conditions. The liquid and solid temperatures are ' non-dimensionalized using the initial steady-state temperature rise across the core (ATo ). 3681w4a.non:1M60897 4-4 l

FACILITY SCALING REPORT The non-dimensionalized' liquid mass balance equation is written as follows: t,,, p ,V ,* = A p,*Q,' (4-11)

                                                         ~ where the residence time is given by te,o, which is defined as follows:

V (4-12) T,,o = Ot. o The non-dimensionalized liquid momentum balance equation is written as follows: Tt.o Peu , V,*)= A p,u,*Q,' +Hgi f *p,*AT

                                                                                                                                                  ,           ,                                (4-13)
                                                                                                                                                      . . .     <         s.

r p,u, , gg l -Up +K j , 2 , dh 1 s a  ; i i g The non-dimensionalized liquid momentum balance equation includes the Richardson number (Hg ;) and the Friction number (U p), which are defined as follows: OTo8ATo l URi " 2 (4~l4) l ug and Up = E +K (4-15) dh

                                                                                                                                                        .o i

The non-dimensionalized liquid energy balance equation is written as follows: P, Cy*, T,' V,* = A p,*Cp*,T,*Q,' I. t H,p, -T 8 Boundary

                                                   . 3681w-4 anon:lb-060897                                                                '45 e_-___________-____-_-______-_-____                                                            _ _ _ - _ _ _ _ _ _ _ _ _ _ _

FACILITY SCALING REPORT This equation includes the specific heat ratio (yo) and the Stanton number (D,t), which are defined as follows: CP'- (4-17) y ,,, = vi,o and HA 3o 5 n,, = (4-18) pgCpeQg The energy balance equation for the fuel rods is written in non-dimensionalized form as follows: Ts.o T, = H,*(r, -T,)* Boundary *M5 where t,,o .s a heat transfer time constant, the inverse of which is the heat transfer process-specific frequenci. That is: 1 H,oA 5 co

                                                                                                               =         =_                                      (4-20)

Ts,o P Cy ,V, The characteristic time ratio (Hq,) is expressed as follows: 95 g5 = (4-21) H,oA,6To Note that the liquid balance equations (Equations 4-11,413, and 4-16) implement the same time scale (t,,o). This permits a direct comparison of the specific frequencies of each of the liquid transport processes. The fuel rod energy balance equation (Equation 4-19) implements a different time scale (t,,o). Using Equations 4-12 and 4-20, the ratio of the liquid residence to fuel rod heat transfer time scales can be written as follows: e =h= lo"80 (4-22) I s,o O f.o@s ysCb l 3681w-4a.non:Ib-060897 46

 . -                 - __ _ - - _- - -_ - _____ -___-_-__- -_ _ - - - -_-_ - - _ -_ _ -__                                                      _-_ _ _ _ _   _- - _ _ ___ - _ _ _- _ m

FACILITY SCALING REPORT l' [

  ,q)

This equation can be rewritten as follows: I' tH ' t (4 23) E.t = u,op,Cy,6 , where S is the conduction depth given by the cross-sectional area of the solid heater divided by the wetted perimeter (Axc@).

                            -When c,j si much less than 1, the liquid variables change more rapidly than those of the solid.

When y,,is much greater than 1, the solid variables change more rapidly than those of the liquid. Thus, the transport time ratios presented in Equations 4-22 and 4-23 can be used to indicate the degree of coupling between the liquid transport phenomena and the solid transport phenomena. (See NUREG/CR-5809, Appendix D).0) 1 The solid energy balance equatica can be rescaled in terms of the liquid residence time scale by

                           - multiplying both sides of Equation 4-19 by E t.t. This yields the following:

d V'O 5t.o T, = E,,, H,*(r,-T,)* + 14g,+ (4-24) where the liquid heat source number (Hq ,) is given by the following equation: (4-25) 14, = PsC y ,ugAT o V, In this form, U ,q can be compared to the liquid H groups having the same time scale. Table 4-3 lists the non-dimensionalized balance equations for the core. Table 4-4 lists the residence

,                           times and characteristic time ratios for the core transport processes. The U groups presented in
                          - Table 4-4 are not expressed in terms of control parameters. Substituting the core energy balance l

equation (Equation 4 3) into Equations 4-14 and 4-25 eliminates the dependence on ATo and yields the i following: l OTo8 9soI Mgi = (4 26) 3 p,oCp,,o u a,

   . O.

36siw-4a.non: baos97. 4-7 i LL _ __ - ._ __

FACILITY SCALLNG REPORT I

                                                                                                                                             . PbC pgV6 e

(4-27) l p,Cy ,V, Equations 4-26 and 4-18 can be fully expressed in terms of control parameters by substituting the steady-state core liquid velocity (u g ) given by Equation 4-4. His yields the following expressions: Fi ( l 4.T2Le (4-28) i and I 1/3 H,3, A ,3 F7 (4 29) m-2 7,q,o lgg(pgCpgc)2 a I { Because the operating conditions and core power can be controlled by the investigator, the fluid ( velocity and the core AT can be adjusted in the model to match the important H groups of the prototype. To evaluate the Friction number (Equation 4-15), the Stanton number (Equation 4-29), the Richardson number (Equation 4-28), and the Heat Source number (Equation 4-27), the designer must specify the following: e fluid properties (Ero, Cp g, pg) ! e material propenies (Cy ,, p,) e rod geometry (A,, d, Ax)

              =     core heated length                                                                  4 (f)
              =     core geometry                                                                                                            (ff/dh , K)
              =     heat transfer regime                                                                                                     (H,o)
              =     core power                                                                                                               (q,,)

A review of the four H groups reveals that by matching the Stanton number and maintaining core l' geometrical and fluid property similarity, core heat transfer can be simulated in the model. He relative importance of these different processes can be determined by writing and evaluating the specific frequency for each process. He specific frequencies (Equations 4-21 through 4-24) are presented in Table 4-5. I To evajuate these specific frequencies, detailed information on the convective heat transfer coefficient l (H,) mid the friction factor are needed. This information is provided through the bottom-up/ptocess , f analysis. l l \ 3681w-4a.non:Ib-060897 48

                                                                                                                                                     - - _ _ _ _            _ - _ - - - _ _ _ - - _ _ _ _ _ - _ - - _ _ _ -                        n

FACILITY SCALING REPORT 4.1.3 Bottom Up/ Process Analysis for the Core V]

 /

The additional information required to evaluate the specific frequencies presented in Table 4-5 is developed in this section. The first is the convective heat transfer coefficient (H3 ). Convective Heat Transfer Natural circulation in the primary loop generates significant flow through the core. To select an appropriate heat transfer model for H, it is necessary to determine if free-convection or forced convection best describes core heat transfer during primary loop natural circulation. The criterion typically used to weight the relative importance of free-convection heat transfer to forced-convection heat transfer is given by the following:(2) Gr

                                                                                                                                              > 1                    (4-34)

Re 2 For the case of rod bundles, the Grashoff number and Reynolds number are defined in terms of the hydraulic diameter (3) as follows:

 .a Orashoff Number.
 /( ~

Gr = 80T fs 4ddh (4-35)

                                                                                                                                                /

(Mt Pt) Reynolds Number: Re = pg g dh (4-36) 14e For values of Gr!Re2much greater than one, free-convection heat transfer would dominate. For values of Gr/Re2much less than one, forced-convection heat transfer would dominate. For the AP600 core, the fluid velocities in the heated subchannel would be very low (0.15-0.3 m/s) under single-phase natural circulation conditions for decay powers ranging from one percent to [ () i 3681w-4tnon:1b-060897 4-9  !

FACILITY SCALING REPORT five percent and for pressures exceeding 1080 psia. However, the Reynolds number would still be quite large. For this range of conditions, the criterion given by Equation 4-34 would vary as follows: l Gr 0.004 s s 0.17 (4-37) (Re)2 Thus, for the problem of interest, a constitutive equation for H, based on turbulent forced convection heat transfer would be appropriate. The standard equation of Dittus-Boelter can be implemented.W k H,, = 0.023 _.". Re E8 Pr E4 (4-38) dh where: Pr = P6M6 = Prandtl Number (4-39) kg Substituting Equation 4-38 into Equation 4-29 yields the expression for the Stanton number: 1/3 N o ^s T 0 (4-40) 11;7 = 0.023 Re nc Pr .4 2dhOTo9soLth8(PmCpg ac)2 where: To9solth8P dbh (44}) Re,c = 3 Cpgp gc7 aF Equation 4-40 represents the characteristic time ratio that should be matched in the model facility to simulate the single-phase natural circulation heat transfer in the full-scale prototype. Friction Factors and Loss Coefficients l The friction number defined by Equation 4-15 includes the component friction factor and the form lo s coefficient. The friction number similarity group will be matched in the model by carefully selecting pipe diameters and using drawn tubing and orifices where appropriate. Standard friction factor 368Iw-4 Anon:Ib-060897 4 10

f i FACILITY SCALING REPORT correlations (Moody diagram) can be implemented. However, it is expected that the pressure drop in

      \     the AP600 loop is dominated by form loss.

4.1.4 Steady-State Loop Similarity Criteria Having performed top-down and bottom-up analyses for single-phase natural circulation, a set of similarity criteria can now be developed for the loop. The component fluid velocity scaling ratio can be found using Equation 4-5. t ' l/3 r ' OTo %Leh ac (4-42) U ti.R

                                                            ,acCpg pFg T ,R s"isR 1

l l where: p = average loop fluid density R = model to prototype ratio 7% h Complete kinematic similarity requires that similitude between flow area ratios be maintained. That is: G!

                                                                 ."1 -1                                              (4-43)                            !
                                                                ,a sR Thus, (ac )g = (a i )g = ag. Furthermore, requiring a fixed length ratio throughout the loop results in the following:

(Lg)g=fg (M) l where the subscript i has been dropped to indicate its validity for the entire loop. It will also be required that the total loop friction number ratio equal one. I (F7)g =l (W f% O 3681w4a.non Ib-060897 4 11

FACILITY SCALING REPORT Applying the requirements of Equations 4-43 through 4-45 to the velocity scaling ratio yields the loop fluid velocity scaling ratio: r ' 1/3 r ' l/3 UR *

                                                                                                                                            , Et oCPlo, R    s C sR
                                                    'Ihe fluid residence time scaling ratio can be obtained by dividing the length scaling ratio (fg) by the loop fluid velocity scaling ratio. That is:

f T g=l M) l UR l Or: 1 1/3 TR" ( OTo9so .R Table 4-6 summarizes the steady-state, single-phase natural circulation loop scaling ratios (Equations 4-49 and 4-50 are included in Table 4-6). j An important case will now be examined- that of imposing the requirement of isochronicity. For the case of isochronicity, the core power is scaled such that fluid velocity ratio equals the length scale j ratio. While real-time natural circulation tests could be performed in a reduced height facility, the convective heat transfer coefficient and thermal boundary temperature difference would be distorted. i Isochronicity Imposing the requirement of isochronicity me2ns that the OSU/AP600 test facility would be capable of simulating single-phase natural circulation flow on a real-time basis. That is, the time required for a fluid element to travel around the model loop would be the same as in the full-scale prototype. This is assured by requirir.g that: Tg =1 (4-51) O 368Iw-4a.non 1t6897 4 12, s

i FACILITY SCALING REPORT j i Substituting this requirement into Equation 4-47 yields the following: uR *IR (4-52) I Substituting this result into equation (4-46) and rearranging the equation yields the following: j r , gg.

• ag[g H-@

To sR l Table 4-7 summarizes the steady-state single-phase natural circulation scaling criteria that satisfy the I requirement of isochronicity, Note that all of the scaling ratios listed in Table 4-7 are expressed in terms of fluid and geometric l parameters. It is apparent that various test geometries will satisfy the requirement of isochronicity. 4.1.5 Single-Phase Natural Circulation Summary .(  ; L Section 4.1 presents the scaling criteria for single-phase natural circulation flow. Specific similarity - l criteria are presented in Tables 4-6 and 4-7 for steady-state natural circulation conditions.

                                   'Ihe similarity criteria developed in this section are integrated with the two-phase natural circulation criteria developed in Section 4.2. After examining both the single-phase and two-phase similarity criteria as a set, numerical values are assigned to the geometric parameters.

4.2 Two-Phase Natural Circulation Scaling Analysis

                                                                                                                           ~
                                  . Under small break loss-of-coolant accide'nt (SBLOCA) or ADS blowdown conditions, the primary system pumps will be tripped. Therefore, two-phase natural circulation will exist in the primary loop (depending on whether the steam generators function as heat sources or heat sinks). To assure that l                                  .this behavior can be simulated in the AP600 test facility, a detailed scaling analysis of loop natural circulation is performed.

o y 3681w-4tnon:it> 060897 4-13

The geometry of interest is the entire AP600 primary coolant loop. Therefore, the scaling analysis l l presented in this section provides a basis for sizing most of the primary loop components. The ( specific phenomena of interest identified by the SBLOCA PPIRT and addressed in this section are as follows:

      . The stability of two-phase natural circulation mass flow rates for various reduced primary-side fluid inventories.

• The transition from stable two-phase natural circulation flow to intermittent or oscillatory flow.
• Hot leg counter-current flow limitations (CCFL).
• Core heat transfer and critical heat flux under two-phase natural circulation conditions.

The similarity criteria required to scale these phenomena are developed in this section. Figure 4-3 presents the loop natural circulation scaling analysis flow diagram. First, a top-down/ system scaling analysis is performed. This includes an analysis at the constituent level for transient conditions in the core and an analysis at the system level (loop) for steady-state conditions. The objective of the top-down scaling analysis are following.

• To scale the steady-state two-phase fluid mass flow rates.
• To scale the two-phase core heat transfer.

Following the top-down scaling analysis, a bottom-up/ process scaling analysis is performed to develop similarity criteria to scale:

• Two-phase friction and loss coefficients a Flow regime transitions
• Hot leg counter-current flow limitations (CCFL)
• Core heat transfer regimes e Critical heat flux The similarity criteria developed through these analyses are compiled and integrated with the single-phase liquid scaling criteria to produce a set of scaling ratios for the AP600 test facility.

4.2.1 Top-Down Two-Phase Constituent Level Scaling Analysis for the Core The transfer processes between two constituents are examined in this analysis: a twephase mixture and the metal fuel / heater rods. Heat transfer to metal walls and to the ambient are examined in Section 5.0. The H2TS approach described in Section 1.3.3 is implemented. Table 4-8 presents the control volume balance equations for the core having a constant cross-sectional flow area (ac ). The 368Iw-4a.non:Ib-060897 4-14

derivation of the constituent level balance equations for a two-phase fluid mixture is presented in Appendix A. The nomenclature for these equations and for the equations in the tables that follow is given in Section 4.1 and supplemented as follows: pf two-phase fluid mixture density Vf fluid mixture control volume Qf fluid mixture volumetric flow rate uf fluid mixture velocity Ap density difference op = (pb - Pp) pp saturated vapor density p3 saturated liquid density a vapor volume fraction (1-a) liquid volume fraction h latent heat of vaporization is hf fluid mixture enthalpy vg drift flux velocity, which is related to the relative velocity between the gas and liquid as vg = (1-a)(ug-u,) Tf is the fluid mixture ternperature ( & is the wetted perimeter 6 is the conduction depth The remaining variables are as previously defined in Section 4.1.1. The two-phase mixture equations presented in Table 4-8 (Equations 4-56 through 4-59) require less information than would be required for an analysis performed at the phase level. The mixture equations differ from the standard drift flux model in that the equations only require information at the constituent level. Thus, a separate mass conservation equation for the liquid or vapor phase is not included. Using the procedure outlined in Section 1.3.3, all of these terms can be non-dimensionalized by dividing them by their respective initial or boundary conditions. The energy transport terms are non-dimensionalized using the initial steady-state enthalpy rise across the core (Aho ). The non-dimensionalized two-phase mixture mass balance equation is written as follows: Tf,o f f =Apf*Q f -4'd PV a Ap+v (4-60)

 \

3681w-4a.non:lb-060897 4 15

FACILITY SCALING REPORT where the residence time is given by fi,,, which is defined as follows: V f-(4-61) T,o = Ot.o f and the Drift Flux number (UNd) is defined as follows:

                                                                                            -"                                          (4-62)

UNd = ug,o pg,o The non-dimensionalized two-phase mixture momentum balance equation is written as follows:

                                                                                     = A pg*ug*Qf     +       Ap *a *V f*

T t.o Pi uf*Vf g+

                                                                                     +1)l(d       A                                     @^O (1 -<x)pg f         i+

U

                                                                                     -Up Pt t+ +Q   t   +it +K 2     ,d h      ,

In this equation, the Froude number (U Fr ) is given by the following: 2 ug,o pf o g, (4-64) Ob ko The Density number is given by the following: Pss,oPh. g, (4-65) Uo(I N )o (OPo) O l 3681w4tnon:Ib-060897 4.] 6 L

FACILITY SCALING REPORT l t l' ! The Friction number (Up )is given by the following: 1 U=p +K (4-66) I

                                                                                                            ,dh       so I

i

                     ' The non-dimensionalized two-phase mixture enthalpic energy equation is written as follows:
                                                                                                                                                                             .l 1

l Tf.o Pr h*V* f f

                                                                                                  = Al pf f*h *Qp* +  Q H,](r,-T f)             soundary

{ (4-67) 1 "PgsPishvgg 1

                                                                                                   +f( Q A                 Pt I

In this eguation, the Stanton number (H,t) is given by the following: H sfoAs (4-68) q , Pf.o C vsuf,o ac I l-and the Enthalpy number (Hh ) is given by the following: l l' bg,o (1 -ot)oaA a po _ ~ (4-69) Aho pf,o i The energy balance equation for the fuel rods is non-dimensionalized by letting: l T*= C,p,) y (4-70) 3 Ah, l i i i l

      . \ /~

/ 36stw-4a.non:ib 060897 4-17 Y______-_=________________-___

l FACILITY SCALING REPORT and 1 C y,(r,-T f) f,-T )*f = (4-71) l I Using these definitions, the non-dimensionalized fuel rod energy balance equation is written as . follows: l Ts,o T, = H,* (r,-T r)*

• II gr9s (4 72) 1 where t,,, is a heat transfer time constant, the inverse of which is the heat transfer process-specific frequency. That is:

l ( I SI 5 c),,, = _ _ = (4-73) l Ts,o PsCy ,V,  ; The characteristic time ratio (TIqf) is expressed as follows: q, Cy, (4,74) gf = H,f,oA,Aho Note that the non-dimensionalized two-phase fluid mixture equations (Equations 4-60,4-63, and 4-67) implement the same time scale (tf ,o). This permits a direct comparison of the specific frequencies of each of the fluid mixture transport processes. The fuel rod energy balance equation (Equation 4-72) implements a different time scale (t,,o). Using Equations 4-61 and 4-73, the ratio of the liquid to fuel rod transport process time scales can be written as follows: Tf,o foH,f,oA,

                                                                                                                =                                                        (4-75) eg = _Is ,o                                                                 O f.oEsC        vs V, O

3681w-4a.nottib-060897 4 18

FACILITY SCALING REPORT [ ) This equation can be rewritten as follows: LJ IH 5-(4-76) ed = uf,op,Cy ,6 where 6 is the conduction depth given by the cross-sectional area of the fuel rod divided by the wetted perimeter (Axc/(). When eg is much less than 1, the fluid mixture variables change more rapidly than those of the solid. When ey s imuch greater than 1, the solid variables change more rapidly than  ; those of the fluid mixture. Thus, the time scale ratios presented in Equations 4-75 and 4-76 can be used to indicate the degree of coupling between the fluid mixture transport phenomena and the solid transport phenomena. (See NUREG/CR-5809, Appendix D.m) The non-dimensionalized fuel rod energy balance equation can be rescaled in terms of the residence time of the fluid mixture by multiplying both sides of Equation 4-72 by eg. This yields the following: T, =c yr H,} (r, -T f)] D (%- Tf,o + Upch9s v) I s , where the phase change number (Upch)is given by the following: 95 (4-78) I g = Es xUf.o o The characteristic time ratios developed in this section are the same as those developed by Ishii and Kataoka.W Table 4-9 su nmarizes the two-phase fluid mixture non-dimensionalized balance equations for the core. Table 4-10 lists the residence times and characteristic time ratios for the core transport processes. d i The H groups presented in Table 4-10 can be expressed in terms of control parameters and fluid properties by substituting the steady-state core energy balance equation given by the following: 7.- q,o = pro fo cu a s, (4-79) l

 '(

V) 368Iw.4a.non:Ib-060897 4.I9

FACILrrY SCALING REPORT Thus, the enthalpy number becomes the following: hfgo(I -"o)"o0Po"f.o"c _ (4-80) 950 and the phase change number reduces to the following: Pfo#c g PsA x (4-81) To evaluate the characteristic time ratios presented in Table 4-10, the following information is needed:

• fluid properties (h f,,, pf,,, higo' Pgso' Piso' 090)
• material properties (p,, Cy,)
• rod geometry (A,, d, Ax) e core heated length (f) a core geometry (fFdh, K)
• heat transfer regime (H sr.o) l
• core power (q,o)

{ The core fluid velocity (u r,o) the core fluid drift-velocity (vgo) and the core fluid void fraction (a o) are functions of global system parameters. Because the operating conditions and the core power can be controlled by the experimenter, the fluid velocity and the enthalpic energy rise across the core can be adjusted in the model to match the important U groups of the prototype. The relative importance of each of these processes can be determined by writing and evaluating the specific frequency for each process. The specific frequencies are presented in Table 4-11. To evaluate these specific frequencies, a detailed bottom-up/ process scaling analysis is performed in Section 4.2.6. 4.2.2 Top Down Integral System Steady State Two-Phase Scaling Analysis Section 4.2.1 presents a top-down/ constituent level scaling analysis for two-phase natural ci rc ulation. The SBLOCA experiments proposed for the AP600 test facility will be initiated from steady-state single-phase forced flow conditions. The initial depressurization leads immediately to two-phase natural circulation conditions. By simplifying the mass, momentum, and energy equations presented in Table 4-8 and integrating the equations over the entire loop, it is possible to obtain a steady-state solution for the loop mass flow rate. As with the single-phase scaling analysis, the integration process 368Iw-4a.non:1b460897 4-20

FACILITY SCALING REPORT [] makes the facility design more flexible. Global system phenomena will be simulated properly, but V some local phenomena may not be simulated. Two-Phase Natural Circulation Loop Flow Rate This section presents an analysis that describes how two-phase natural circulation flow rate (two-phase fluid velocity and mass flow rate) in a thermal-hydraulic model is expected to scale with facility size. This analysis results in an analytical expression for the two-phase fluid velocity. Figure 4-4 depicts the loop geometry considered for this analysis. The loop is divided into two regions: a two-phase region with a fluid density (pf) and a single-phase region with a fluid density (p,). The simplifying integral analysis assumptions are as follows:

1. Steady-state flow.

2. One dimensional flow along the loop axis.

l 1 1

3. Uniform fluid properties at every cross-section. l

4. Homogeneous flow.

p') 5. Chemical equilibrium -- no chemical reactions. l LJ

6. Thermal equilibrium -- both phases at the same temperature.

7. Mechanical equilibrium -- velocity and pressure of both phases are equal (vg = 0).

8. The sum of convective accelerations due to vaporization and condensation are negligible.

9. Viscous dissipation is negligible.

Because the loop contains both a single-phase liquid region and a two-phase fluid region, the assumptions listed above are applied to the mass, momentum, and energy equations given in Table 4-2 and Table 4-8. The simplified set of equations are then integrated over their respective single-phase and two-phase regions to obtain the loop balance equations. These equations can be used to obtain an analytical expression for the fluid velocity at the core entrance. It is shown that the coefficients of the fluid velocity equation can be scaled in a meaningful manner to obtain a two-phase fluid velocity scaling ratio, a two-phase power density scaling ratio, and a two-phase time scaling ratio. A complete derivation is presented in Appendix B. i (] !

  'w/

368Iw.4.tnon:1b.060897 4-21

FACILITY SCALING REPORT Loon Mass Conservation Eauation For steady-state, one-dimensional flow, the mass conservation equation at every flow cross section along the loop is written as follows: rh =p;u;a; wonstant (4-89) where: th = constant mass flow rate for the system p; = fluid density and fluid velocity within the i* component u3

           =    fluid density and fluid velocity within the i* component a3
           =    flow cross-sectional area of the i* component Steady-state conditions require that:

p;uia;=peua ce (4-90) where: pc = fluid density at the core entrance ue = fluid velocity at the core entrance ac = flow cross-sectional area at the core entrance l Looo Momentum Conservation Eauation l Applying the assumptions stated previously, the steady-state momentum balance equation is integrated over the single-phase and two-phase regions to obtain a general force balance relating frictional and f buoyant forces. That is: N p u'.2 'fg +K =(p,-pf)L g E e (4-91) i =1 2 ,d h si l O 3681w-4a.non:Ib-060897 4 22 __________d

FACILITY SCALING REPORT where: p[ = - average fluid density in the single-phase region p[ = average fluid density in the two-phase region Le = distance between the thermal centers of the heat source (core) and the heat sink (steam generator) f = single-phase or two-phase friction factor

                                                                 ~fi        = ~ length of the i* component dhi .    =   i component hydraulic diameter Ki       =   i* component loss coefficient                                                                   .

1 Using Equation 4-90, Equation 4-91 can be written in terms of single-phase and two-phase regions as .)

                                                                 ' follows:

s > i

                                                                                         ] ]        f            f i]          f    i f i]

dE f1+K ** +d E f1 +K f %ff)L ag (4-92) 2 pg 7p d h , ; , ai , p,SP, dh ,j,a, For this analysis, p[is the average two-phase fluid density given by the following definition for a two-phase fluid under homogeneous equilibrium conditions.

     ,                                                                                                                        Pes
                                                                                                                                                                    - (4-93)

AP 1 +x ,

                                                                                                                                , pgs, where:

p. = saturated liquid density p,, .= saturated vapor density (Ap = ps - pgs) 1 x; = equilibrium vapor quality

                                   .j
                                                                 ' 36siwanon: leis 97                                      4-23

l FACILTIT SCALD.G REPORT Loop Enercy Conservation Eauatiqn Under steady-state conditions, the energy balance for the core, which governs the rate of energy transport into the system is given by the following: pcua c c(h f-h,) =q, (4-94) where q, is the core heat generation rate less heat losses from the core. The steady-state energy removal rate from the system is given by the following: 9s "9so + 9 toss (4-95) where: qso = steam generator heat removal rate qLoss = heat loss from the system components other than the core The equilibrium vapor quality used in Equation 4-93 can be defined as follows: hf-hh x* = (4-96) h,, Substituting the energy equation (Equation 4-94) into Equation 4-96 yields the following: 9 h x = ' W ' ' c sub (4 97) Peuah c c ts where: h sub = subcooling enthalpy; hsub = h,, - hl

                                 =

h,g latent heat of vaporization; h,g = hg , - h 3 The steady-state loop balance equations are summarized in Table 4-12. O 3681wAnon:tb-071897 4-24

FACILITY SCALING REPORT O Two-Phase Mixture Density Eauation (b Substituting Equation 4-97 into Equation 4-93 yields the following:

g. P85

[q,-pc cuah c sub

                                                                             ,    Peuah c c ts   n Pss s Substituting Equation 4-98 into Equation 4-92, applying the Boussinesq approximation such that pc = p, = pg,, and performing a significant amount of algebra, yields the cubic equation for fluid velocity listed in Table 4-13.

Equations 4 99,4-99a through 499c,4-100, and 4-102 are found in Table 4-13. Steady-State, Two-Phase, Loop Similarity Criteria In order to simulate the same fluid velocity behavior in the model as in the full-scale prototype, the

       ._                       coefficients of Equation 4-99 (found in Table 4-13) must be scaled properly. It is interesting to note

[ that a mathematical approach for the scaling of Equation 4-99 is inferred in a recent text by Schroeder,

                              . entitled Fractals, Chaos, Power Laws.M A set of two-phase, natural circulation similarity criteria can be obtained by scaling the coefficients $,,
                                $ b' &c of Equation 4-99 such that the following transformation is possible:

3 3 2 3 (un, + $,u 2n+&bm um - $cm) = u p + $,pu p + pp u -$cp H-M where:

                                         =    a constant factor m        =    values for the model p        =    values for the prototype l

The transformation laws for scaling polynomials are derived in Appendix C. A general method of L scaling single-state variable catastrophe functions is by Reyes.O By satisfying the requirement of Equation 4-103, the basic equation remains unaltered. The only change is a change of scale in the

    . ,m                                                                                                                                    j e         \

k 368iwAwn;1b-060897 4-25

I FACILITY SCALING REPORT l solution. The condition stated in Equation 4-103 can be satisfied mathematically by scaling the velocity and the equation coefficients as follows: um =u p (4-104a)

                                                                                           $&am=&ap                                                                                                                               (41Nb) p2@bm=&bp                                                                                                                              (4- W c) 3 0&cm=&cp                                                                                                                               (4-104d)

These equations can be rearranged in terms of scaling ratios as follows: UR = 1/ (4-105a)

                                                                                         $aR=1/                                                                                                                                   (4-105b)
                                                                                         &bR = 1/                                                                                                                                 (4-105c)
                                                                                         $cR = 1/                                                                                                                                 (4-105d) where the subscript R represents a ratio of the model parameters to the prototype parameters.

The physical scaling implications imposed by Equations 4-105a through 4-105d can now be determined for two special cases: the case of saturated conditions and the case of property similitude. Two-Phase Similarity Criteria at Saturated Conditions During a SBLOCA, the primary system fluid is typically at saturated conditions. As a result, the value of h subi s zem. Thus, Equations 4-99a through 4-99c can be simplified as follows: Ap F TP 4, = 9 s 3 (4-106a) ac , p,3 pg shig, , FT,

                                                                                                                                                                                     ,2 e            *2       '

9s Ap <FTP 4g (4-106b) a e, p,3 pg,h,g , ,F,T 2 Ap 4 , gq,L th (4-106c) ac p,3pssh ig, FT; O 3681w-4a.non:lb-060897 4-26

FACILITY SCALLNG REPORT

              ' To achieve kinematic similarity, similitude between flow areas will be maintained. That is:

r 3 a _.f =1 (4-107) 6 IsR It will also be required that the single-phase and two-phase friction numbers (FSP and FTP) be the same in the model and the prototype.

      '       - Thus:

(FSP)R .= 1 (4-108) l (FTP)R = 1 (4-109) (F7)g =1 (4-110) L. Applying the requirements presented in Equations 4-107 through 4-110 to the coefficients of l

             ' Equations 4-106a through 4-106c and taking the ratio of the model values to the prototype values         j
              . yields the following:

e , r , 4aR . (4-111a) a ( cR6 P85P s 8 ,h ig , R

                                                                      $2    r          '2 kbR "                                             (~

aC Rs PISP8,his s R. l r , e , 4cR = (41lic) g E c j g [isPgshig, g i i 3681w-4a.non:Ib-060897 4 27 b. 1 . .

FACILITY SCALING REPORT Specific similarity criteria can now be developed by substituting Equations 4-llla through 4-llic into Equations 4-105b through 4-105d. Substituting Equation 4-llla into Equation 4-105b yields the following: e , r , I 95 Ap

                                                                                                                                       .=                                                                                                                                (4-112)

O ac Pf5Ps,hig s R s s R Setting Equation 4-112 equal to Equation 4-105a yields the following:

                                                                                                                               -              s e                          s 95               AP                          3                                                                                (4 113) ua   gCsR, fs gs fgs s                                                R The requirement of Equation 4-113 has implications on the vapor quality. For saturation conditions, the vapor quality Equation 4-97 can be simplified as follows:

8 x, = (4-114) Pisuah c c ig Substituting Equation 4-114 into 4-113 yields the following: xR -.P =1 (4-115)

                                                                                                                                                  , Pgs, g Another scaling criteria can be obtained by substituting Equations 4112 and 4-llic into Equation 4-105d. His yields the following:

fs Pgsh ig (4-116) qq ( ) s AP ,R O 3681w-4a.non:lb 060897 4-28

FACILITY SCALING REPORT

 -'s./  (']     Combining Equations 4-113 and 4116 yields the fluid velocity scaling ratio:

ug4)th N )

                   ~

The time scaling ratio is obtained by dividing the system length scaling ratio by the velocity scaling ratio given by Equation 4-117. Thus: IR Ta" in (4-118) (L )R th 1 The mass flow rate scaling ratio is obtained by the following: ' tha =(p,,)g(Lth ) ag WW

 .;         i Table 4-14 summarizes the steady-state,' two-phase natural circulation loop scaling ratios for the special case of saturated conditions. For the case of a fixed length scaling ratio and geometric similarityi (L th)R "IR                    (4-120)

The scaling ratios then become identical to those developed by Ishii and Kataoka.W Two-Phase Similarity Criteria with Property Similitude By requiring fluid property similarity, it is assumed that the operating conditions in the model are identical to those in the full-scale prototype. Therefore, the ratio of all the fluid and material properties equals one (for example, pg = ha = 1). In addition, complete kinematic similarity requires that similitude between flow 1rea ratios be maintained. That is: f i "C

                                                                                                                   =]                  (4-121)

L i sR

      \

3681w-4a.non:lb-060897 , 4 29 l L

FACILITY SCALING REPORT It will also be required that the single-phase and two-phase friction numbers (Fsp and FTP) be the same in the model and the prototype. Thus: (FSP)R =1 (4-122) (FTP)R =1 (4-123) (F7)g =l (4-124) Applying the requirements of Equations 4-121 through 4-124 to the coefficients of Equations 4-99a through 4-99c and taking the ratio of the model values to prototype values yields the following: haR

• a

                                                                               ,    CsR 15

• kbR= (4-1256)

($3q, +$2L g),

                                                                                                                                                                             ~

DcR

• s "C sR where $3 and $2 are fluid property groups that would be identical in tne model and prototype.

Specific similarity criteria can now be developed by substituting Equations 4-125a through 4-125c into Equations 4-105b through 4-105d. Substituting Equation 4-125a into Equation 4-105b yields the following: I 95 __. _ (4-126) 6 CsR O 3681w-4a.non:Ib-060897 4-30

FACILITY SCALING REPORT I - OV Setting Equation 4-126 equal to Equation 4-105a yields: 95 1 (4-127) ua cC s R

                                                                                        , The requirement of Equation 4-126 has implications on the vapor quality. Using the definition of                                                                  f vapor quality provided in Equation 4-97, the vapor quality scaling ratio can be written as follows:

f i f i 9s h sub [ XR*f i ( 8) l 9: h sub I

                                                                                                                                                             &i -

l s ua,p cc , h g, j By requiring fluid property similarity, the following subcooling ratio is obtained: V, 'h sub

                                                                                                                                                                 'h sub (4-129)
                                                                                                                                         ,hg,                          h g, This requirement, coupled with Equation 4-127, means that:

XR -1 (4 130) i - Thus, the vapor quality at the core exit would be identical in the model and the prototype, J 1. l Another scaling criteria can be obtained by substituting Equations 4-126 and 4-125c into Equation 4-105d. This yields the following: p l 1/2' 9R9 R(I'th)R -(4 131) \ . lbl i . L 3681w.4a.non:1b-060897 .4 31 I l, ,

FACILITY SCALING REPORT Combining Equations 4-105a,4-126, and 4-131 yields the fluid velocity scaling ratio: urb) The time scaling ratio is obtained by dividing the system length scaling ratio by the velocity scaling ratio given by Equation 4-132. Thus: TR* (4-133) in (Le)g The mass flow rate scaling ratio is obtained by the following: thg 4 th) aR

                                                                                                                                                                              ~

Table 4-15 summarizes the steady-state, two-phase natural circulation loop scaling ratios for the special case of fluid property similitude. For the case of a fixed length scaling ratio, it noted that: (Lg)gq g (4-135) The scaling ratios would then become identical to those developed by Ishii and Kataoka.W 4.2.3 Integral System Scaling Ratios for Single Phase and Two-Phase Regions As shown in Figure 4-4, both a single-phase region and a two-phase region can exist during two-phase natural circulation. As a result, it is important that the transition from single-phase flow to two-phase flow be scaled ptoperly. This section addresses the transition phase and presents a unified set of scaling criteria for the single-phase and two-phase regions. l l The simplest method of assuring that the single-phase to two-phase flow transitions are scaled properly is to require that the two-phase power scaling ratio equals the single-phase power scaling ratio.W By substituting the two-phase power scaling ratio (Equation 4-116) into Equation 4-46 (assuming I geometric similarity), a new single-phase velocity scaling ratio is obtained: 368iw-4 anon:tb-060897 4-32 l

FACILITY SCALING REPORT t b e yn in TPashig uR.sp =fR (4-136) 3pgPI sR 1 Similarly, the single-phase residence time scaling ratio would become: In h ' -In TPgs ig 7R.sp ,IR (4-137)

                                                               $ ApCPD sR Therefore, with respect to integral system flow behavior and single-phase to two-phase flow transitions requiring the same power scale results in unifying the single-phase and two-phase scaling ratios (assuming material property and geometric similarity).

Table 4-16 presents the combined single-phase and two-phase loop scaling ratios for natural circulation j with single-phase and two-phase flow regions. He scaling ratios presented in Table 4-16 are identical ! to those developed by Ishii for the case of material property similitude.W {A} ~ The scaling ratios presented in Table 4-16 are functions of pressure. This suggests that high pressure ! natural circulation can be simulated in a low-pressure test facility. However, when the same working i fluid is used in the model and the prototype, time scaling and velocity scaling' distortions arise in the f l single phase region because of fluid property differences. These scaling ratio distortions are presented graphically in Figures 4-5 through 4-8 for four different model pmssures and a wide ras.ge of

       . prototypic pressures.

Figures 4-9 through 4-12 present the model powers required to simulate steady-state, natura!

! circulation in a high pressure prototype at five decay heat levels. This analysis was performed for four L different model pressures. 1 i For the special case of fluid and material property similitude, the scaling ratios in Table 4-15 simplify to obtain the scaling ratios presented in Table 4-17. (Equations 4-130 through 4-134 are found in j Table 4-15; and Equations 4-145 through 4-152 are found in Table 4-17.) l

                                                                                                                                                               \

j 4.2.4 Evaluating System Scaling Ratios for a 1/4 Length Scale Model l By specifying a length scaling ratio and the diameter scaling ratio for the primary loop, the system scaling ratios (power, velocity, residence time, and mass flow rate scaling ratios) given in Table 417 can be evaluated for both the single-phase and two-phase regions of the loop. Multiplying the.

 ~(q
   ,/-  geometric scaling ratios by the dimensions of the full-scale AP600 nuclear steam supply system yields the dimensions for similar components in the APEX facility.

3681w-4a.non:Ib-060897 4-33 l

FACILITY SCALING REPORT The detailed rationale for selecting the length and diameter scaling ratios is presented in Section 1.0. For the APEX facility, the length scaling ratio has been set to 0.25 and the diameter scaling ratio has been set to 0.14428 (that is, a ratio of 1:6.931). The reason for this selection is based on minimizing power requirements, while maximizing height and maintaining sufficient system volume to properly model loop pressure drop and three-dimensional behavior in the downcomer, core, and plenum regions. Using these choices, the loop scaling ratios are determined to be as follows: U R,sp W R.tp =0.5 (i.e.,1:2) (4-153) Tg,,p =tg,tp=0.5 (i.e.,1:2) (4-154) siig,,p =sg,tp =0.0104 (i.e.,1:96) (4-155) qg,,p mg,tp4.0104 (i.e.,1:96) (4 156) In addition to the scaling latios presented in Equations 4-153 through 4-156, the area, volume, and power density scaling ratios are also defined. System Area Scaling Ratio The area scaling ratio for the primary loop is defm' ed as follows: a ag =. m. (4-157)

                                                            "P

, j l The flow area ratio is related to the diameter scaling ratio as follows: I (aR) *

  'lhe numerical value for ag si 0.0208 (that is, a ratio of 1:48.04).

Volume Scaling Ratio Tbc volume scaling ratio is the area scaling ratio multiplied by the length scaling ratio. That is: 3681w-4a.non:lt>.060897 4 34

FACILITY SCALING REPORT A) .

 \-'                                                                                                                                    V Vg=     =aR IR                   (4-159) i Ris yields a volume scaling ratio of 1:192.                                                                        ;

1 Power Density Scaling Ratio The power density scaling ratio is defined as follows: m 9R 1 9R (4-160)

                                                                                                                                       "7 R"-(75 R

This yields a power density scaling ratio of 2.0. Table 4-18 summarizes the scaling results for the 1/4 length scale model. L f~3 'C 4.2.5 Summary of Top Down System Scaling Analysis ne top-down system scaling analysis implemented a homogeneous equilibrium model to obtain the integral system two-phase scaling ratios for a natural circulation system. The primary focus of the analysis was to examine how the global loop flow rate is affected by scale. t A major difference between this analysis and the analysis of Ishii s)is the method in which the

                                                      - velocity scaling ratio was obtained. The analysis presented herein resulted in a cubic, mixture flow velocity equation (Equation 4-99). This implies that the mass flow rate in the system may achieve the most stable of three flow rate solutions or possibly oscillate between solutions. By using the similarity transformation requirement (Equation 4-103), a set of mathematical requirements were imposed on the coefficients of the cubic velocity equation.

These mathematical requirements must be satisfied to assure that the model will be capable of achieving a similar family of scaled flow rate solutions. Implementing these mathematical requirements permits one to obtain the fundamental scaling constant . This scaling constant is then used to obtain the important two phase scaling ratios. An important result of this analysis is the observation that the family of scaled flow rate solutions can (3 only be obtained when the two phase scaling criteria are satisfied as a set. Scaled loop flow rate () behavior cannot be obtained when each criterion is treated independently. [ ' 368Iw.4a.non.Ib-060897 4-35 i

FACILITY SCALING REPORT In summary, the following observations are made for the AP600 test facility which uses the same working fluid and long-term cooling operating conditions as the full-scale system: l

                                     . The set of two-phase scaling criteria presented in Table 4-15 will be satisfied as a set in the AP600 model. 'Ihus, it will be capable of producing a family of scaled flow rates that would be characteristic of the full-scale prototype.

• An examination of the two-phase scaling ratios obtained in this analysis reveals that these ratios are identical to those obtained by Ishii when the thermal center length scaling ratio equals the component length scaling ratio.
• Because fluid property similarity will be maintained operationally and the thermal center length scaling ratio will be fixed by system geometry, the greatest challenge to model performance will be maintaining the single-phase and two-phase region frictional similarity as stated by Equations 4-150 and 4-151. The careful selection of pipe diameter and the use of orifices adequately respond to this requirement.

4.2.6 Bottom Up Scaling of Two-Phase Natural Circulation Processes Sections 4.2.1 through 4.2.4 presented top-down scaling analyses at two levels: the constituent level (which provides specific similarity criteria for each component) and the integral system level (which provides similarity criteria for the loop as a whole). As shown in Figure 4-3, bottom-up scaling analyses are needed to scale important processes in the core, the hot leg, and other components. This section presents scaling analyses for the following:

• Two-phase friction and loss coefficients a Flow regime transitions
• Core two-phase flow processes Scaling Two-Phase Friction and Loss Coefficients As identified in the top-down scaling analysis, it is important that the friction number for the two-phase region of the model be the same as that of the full-scale AP600. This requires a closer examination of two-phase friction factors and loss coefficients.

Dividing the twc,-phase friction number as defined by Equation 4-101 into a friction factor component and a loss coefficie.-t component, a two-phase friction multiplier and loss coefficient multiplier can be incorporated as follows:M O 3681w4tnon:lb-060897 4 36

r; FACILJTY SCALING REPORT l k FTP = Nn +E NU TP TP (4 161) I- where:

                                                                                                                   .< s2 L

ft 1 +(Apx/pg,) _ ac (4-162) Nn=_d h (1 +Apx/p ,)u , aj , g and

                                                                                                   ,           ,,    ,2 M     ac 1+Apx                                          (4-163)

Nu =K s Pgs .t a3 , V l l where:

f. = Darcy friction factor for liquid i i K- = loss coefficient ')

/~ . '

l6 Because of the low fluid velocities encountered in natural circulation flows, the dominant losses will

                                                                                                                                                               )

be due to fonn losses rather than friction. Thus the loss coefficients will be carefully modeled.  ! Because proper flow scaling requires scaling the total friction number in the two-phase region rather than in each individual component, it is possible to satisfy this criteria through the careful use of flow orifices and fluid property similitude. It is noted that Equations 4-162 and 4-163 are identical to the friction number similarity criteria and I orifice number similarity criteria developed by Ishii and Kataoka. Scaling Two-Phase Flow Pattern Transitions and Minimizing Surface Tension Effects Transitions in two-phase flow patterns may significantly impact the integral behavior of the system. Therefore, an assessment must be made for each component in the system to determine which

                                                 -two-phase flow pattern transitions may be delayed or entirely missed. In their paper, Schwartzbeck
                                                . and KocamustafaogullariN catalogued the applicable flow pattern transition criteria. These criteria can be used to determine if two-phase flow pattern transitions would be properly scaled in the horizontal and vertical sections of the model.

O

(O l

                                                . 3681w-4a.non:lN97                                    4-37 l

I FACILITY SCALING REPORT For two-phase flow in horizontal pipes, the following flow transition scaling criteria can be used:

                                                                  .              Stratified-Smooth to Stratified-Wavy (10) e

• l/2 r '

l/2 g/4 =0 (I~E)R OP Vf pq R R (g _X)R s 6 p p8sR e Stratified to Intermittent or Annular-Dispersed Liquid (II) t/2

                                                                                                                                                                                                                         /

U R=DR (4-165)

                                                                                                                                                                                                                                    , Pf , g
                                                                .            Intermittent or Dispersed Bubbly to Annular-Dispersed (30)

Pf (4-166) G - R 'MP sR

                                                                =            Intermittent to Dispersed Bubbly (to) 0.3 foA5 ,                                      IT ("R) aR (4-167)

U VfR (Pf/Ap)U5 g (1-x)0# g For vertical pipes, the following can be used:

                                                                =           Bubbly to Slug Flow (30) f             %

a R " Pf (4-168) APsR O 3681w-4a.non.lb-060897 4 38

i l- FACILITY SCALING REPORT 1 l

        'T              =    Slug to Churn FlowCO)

i. :[d

}, n  ! fg. = a[ fT(x,p) (4 169) I L ' L

• Slug / Chum to Annular FlowUD)

L r 2 'I/4 l- - Pg ug _ ,1 (4-170) SAPsR l l l where: i o = surface tension v = kinematic viscosity l f7= fluid property function as defined in Reference 10 x = vapor quality , p-h An examination of these flow pattern transition scaling criteria reveals that it is impossible to satisfy all of these criteria in a scale model. Derefore, emphasis is given to the important flow pattern transitions identified for various components. i 1 The Idaho National Engineering Laboratory (INEL) recently performed an evaluation of scaled integral test facility concepts for the AP600.02) As part ofINEL's study, they used the flooding review of  ;

                 = Bankoff and LeeU3) to select piping diameters that minimize surface tension effect , As long as the dimensionless diameter (D*) exceeds a value of approximately 32, the geometry can be considered large enough to minimize surface tension effects. The dimensionless diameter is given by:

L n 2 D*= gd 3p (4171) G which is the square root of the Bond number. ! De following sections examine two-phase flow pattern transitions and surface tension effects for the hot and cold leg piping and the pressurizer surge line. A t i

_'%f 3681w-4a.non

Ib 060897 4-39 i

FACILITY SCALING REPORT Hot and Cold Leg Scaling Under two-phase natural circulation conditions, the flow behavior in the hot leg can undergo transitions from stratified flow to either intermittent-slug flow or annular-dispersed flow. This was an important feature of the FLECHT-SEASET experiments in which the hot leg diameter was increased to a maximum in an attempt to model these transitions as closely as possible.(14) The type of flow pattern in the hot leg greatly influences the loop flow rate. A transition from an intermittent-slug flow to a stratified flow in the hot leg will cause a significant change in the mode of heat transfer removal in the steam generators. Therefore it is important to assure that this flow p.attern transition occurs at the same relative liquid level (h/d) in the model hot leg as in the prototype. This type of flow transition can occur in any of the horizontal piping that may be in a drain-down situation. Therefore, special attention is now given to the reactor coolant loop hot and cold legs. Based on Taitel and Dukler's work,(") appropriate scaling of the hot legs requires satisfying the modified Froude number ratio. (Ris dimensionless group has also been used by Wallis to scale flooding behavior in vertical tubes.) He modified Froude number is given by the following: 8 Fr = ' 8 gM (4-172) HL Pgs, i l where j, is the superficial vapor velocity given by ug a. Exact scaling requires that flow regime transitions occur at the same critical Froude number and dimensionless liquid level in the model and in the AP600 hot legs. For identkal fluid properties, with a g= 1, the ratio of the model to the AP600 critical Froude number can be written for the hot leg and set equal to one as follows: HL.R (Fr)g = =1 (4-173) (DHL,R) l l l l~ 3681w-4a.non:lb-0M897 4-40 l

FACILITY SCALLNG REPORT p where it has been recognized that the hot leg vapor velocity ratio equals the hot leg two-phase velocity V ratio (uHL R)- His criteria requires that the local velocity ratio equal the square root of the hot leg diameter ratio. That is: uHL,R = (Dgt,g)W @W Assuming that the loop velocity ratio is an independent parameter, the approach to determining the hot leg diameter that satisfies the modified Froude number would be to rearrange Equation 4-174 as , follows: l 1 r Dat,g = (ugt,g/ @- W I

                                                                                                                                                                                                                              .j
                            - Substituting the loop velocity. scaling ratio into Equation 4-175 yields a hot leg diameter scaling ratio.

This approach is incorrect for a closed loop because it violates conservation of mass requirements.

             ).            . Because the mass flow rate ratio is fixed by global loop parameters, and must be constant throughout the entire loop, increasing a component diameter must necessarily decrease the local velocity ratio.

Therefore, the correct approach is to solve Equation 4-175 simultaneously with the mass conservation

                           . equation for the hot leg. De mass conservation equation for the hot leg is as follows:

ri1HL.R

• UML,R aHL,R (4-176)

, Substituting Equation 4-175 into Equation 4-176 and expressing the hot leg area ratio ii terms of the hot leg diameter ratio squared yields the following: I: thHL,R = (DHL,R)

                           - Rearranging Equation 4-177 to solve for the hot leg diameter ratio in terms of the mass flow rate ratio yields the' following:

Dat,g = (riint,gf @- @ r)

                           - 3681wAnon:Ib-060897                                                                                                                      4 41 L                             _ _ _ _                                  - - - - - - . _ _ - - - - - - - - _ _ - - - - - - - _ - - - - - . - - - _ - - - - - _ . - _ - - - - - _ - - - - _ - - - - - _ -

FACILITY SCALING REPORT l Substituting the numerical value for the constant mass flow rate ratio (1:96) into Equation 4-178 yields a hot leg diameter ratio: DiiL.R = 0.1613 (4-179) Thus for the 31-inch (78.74-cm) hot leg used in the full-scale AP600, the model would be required to implement a 5-inch (12.7-cm) diameter hot leg to satisfy the scaling criteria. Substituting the numerical value of D 3tt,g into Equation 4-176 yields the value of the fluid velocity scaling ratio in the hot leg: U HL.R = 0.40 (4-180) Equations 4-179 and 4-180 satisfy both the Froude number scaling requirement and the mass conservation equation. In addition, it is desirable to maintain the time scaling and volume scaing ratios constant for the entire loop. His requires that the length scaling ratio for the hot leg be reduced accordingly. Multiplying the time scaling ratio (0.5) by the hot leg velocity ratio yields the required hot leg length scaling ratio. That is: O f34t,g = 0.2 (4 181) It should be noted at this point that the volume scaling ratio for the model was selected through an iterative process which assured that this important flow pattem transition would be scaled properly while maintaining a constant volume and time scaling ratio. This is tested by calculating the hot leg volume scaling ratio: Vat,g = f 3;t,g(D;;t,g)2 H- @ Substituting the aumerical values: Viin,g = 1:192 (4.i83) Thus the volume scaling ratio is preserved. 1 0 3681w-4a.non:lt460897 4 42

FACILITY SCALING REPORT [ - The AP600 design implements cold legs that are elevated above the hot legs. During drain-down situations, two-phase flow pattern transitions may become important. The analysis performed for the cold legs is the same as that performed for the hot legs. Thus the scaling ratios are identical. Figure 4-13 shows the flow regime transition boundaries for the AP600 and OSU model hot legs. The 5-inch diameter model hot leg overlays the AP600 curve exactly. By satisfying the modified Froude ratio given in Equation 4-172, counter-current flooding behavior (CCFL) is also approximated.M The flow pattem transitions for the hot and cold legs can be evaluated by substituting the diameter, velocity, and length scaling ratios presented in Equations 4-179 through 4-181, respectively, into the scaling criteria presented in Equations 4-164 through 4-167. The I results are presented in Table 4-19. With respect to minimizing surface tension effects in the hot and cold leg piping, Figures 4-14 and 4-15 illustrate that the diameter ratio of 1:6.2 selected for the hot and cold leg piping results in r values of D*, which exceed the critical value of 32 over the full range of pressures of interest. l Pressurizer Surge Line Scaling Flow pattern transitions in the pressurizer surge line may be important to the operation of the ADS. Examining the flo v pattem transition criteria presented in Equations 4-168 through 4-170 indicates l %.J . (9 that an appropriate scaling criteria would be the following: i

                                                                                                                                                   ~

Isurge.R "Usurge.R [ This requirement, coupled with that of fluid property similarity, would satisfy the bubbly to slug flow l transition and slug flow to chum flow transition criteria. It is difficult to determine if the transition from slug flow to annular flow would be scaled in this geometry since the velocity ratio that must be considered is the surge velocity. Unfortunately, the surge velocity is a fu.iction of many system operation conditions. Oscillations in the surge line will be a function of the system compliance. This is discussed in Section 5.0. In general, by requiring the t/D ratio to be the same as in the plant, the friction number under fluid property similitude conditions is satisfied. ' The effects of sizing the pressurizer surge line using Equation 4-184 are presented in Table 4-19. With respect to minimizing surface tension effects in the pressurizer surge line, Figure 4-16 illustrates that the diameter ratio of 1:4 selected for the pressurizer surge line results in values of D*, which exceed the critical value of 32 over the full range of pressures of interest.

         /

U l i 3681w-4a.non:lb-060897 4-43

FACILITY SCALING REPORT Scaling Core Two-Phase Flow Processes (Fluid Property Similitude) The core simulator in the APEX test facility provides the energy that ddves single-phase and two phase natural circulation in the loop. The important two-phase flow processes for the AP600 core l occur after ADS 4 operation when fluid property similitude exists. The effects of core vapor generation on IRWST injection and long-term recirculation cooling are of particular interest. The l following aspects of the core design have already been fixed at the system level, as indicated in l Table 4-18: 1

• Total core power  !
• Core flow area
• Core height The number of rods, the rod spacing, and the rod diameter can be determined by considering the core heat transfer processes of interest. This is done by examining the similarity groups developed at the constituent level and presented in Table 4-10. Inserting the numerical values of the scaling ratios

                                                                                                                            ]

presented in Table 4-18 into the definitions provided in Table 4-10 and requiring property similitude, it is found that the following constituent level, two-phase similarity criteria have already been satisfied for the core: (UFr)R

• I (Up)R
• I (Uh )R
• I (Upch)R
• I

'Ihe remaining criteria will be examined more closely.

Core Heat Transfer The core heat transfer similarity criterion is given by the Stanton number (UST). Because boiling is an extremely effective heat transfer mechanism, the ternparature difference between the heater surface and the fluid saturation temperature is usually quite smal!. Therefore, requiring exact similitude of the boundary layer temperature difference by forcing (USr)R = 1 is not necessary. However, a heat transfer phenomena which should be examined more closely is the critical heat flux. O 3681w-4a.non:lt>20897 4-44

FACILITY SCALING REPORT i l f T The following scaling criterion, developed by Ishii,M. will be implemented: {

                                                                                                                                                                                      .J l
                                                                                                                                 'a          '

! E

                                                                                                                                                 =l                        (4-185)
                                                                                                                                                                                         )

kHF.R = J

                                                                                                                                 ,9c, ROD, R Various models for critical heat flux are available in the literature. For transients involving pool                              ;

boiling or reverse flow, the modified Zuber correlation for low flows may be applicable /IO For slow transients at the low flow rate range, such as the injection and recirculation processes of the AP600,

                                                  . the following critical heat flux correlation, developed by Katto for low flow two-phase natural                                       I
                                                   - circulation is recommended 37F i-so.043 l                                                                                                                 #      Gegh h d .c      Upas
                                                                                                                                                     .h sub                (4-186)  H t                                                                                                               9 CRIT ,

4(c h G,2f,e 4 1 where Ge is the mass flux through the core, given by the following.  ; ii rh ,. G e =_.C. (4-187) ! a, , - where: l K

o. = - surface tension dh.e

                                                                                    =     subchannel hydraulic diameter i

For a square array of heated rods, the hydraulic diameter is given by the following: 1 2 2 dh .c"4[sxd-zd f43' (4-188) where: r' s = ' pitch (distance between the center lines of adjacent rods) g , I: 3681w-4a sion:Ib o71897 4-45 r

          -_.-.___-_-____.._l_.-_-______.-----

FACILITY SCALING REPORT i d = heater rod diameter The heater rod heat flux is given by the following:

                                                                   #          959 (4-189)
                                                                 .9c,RW =Nxdi c

where: Fq= hot channel peaking factor q, = total core power N= total number of rods Substituting Equations 4-186 and 4-189 into the similarity criterion given by Equation 4-185 yields the following:

                                                                      '       '0.043 Ngdg(d .c)RG h     g    o p,,                                                h sub
                                                                                                                                           ,3                                              (4 190)

(Fq)gqR hg G[fc' .R Equation 4-190 offers some flexibility in selecting the number of rods, the rod diameter, and the rod spacing (hydraulic diameter). Table 4-20 presents the core configuration selected to meet this criterion. In addition to meeting the CHF criterion, core symmetry, geometric similarity, and economic factors (that is, rod number and standard rod sizes) were also considered in the selection. O 368Iw-4a.nortIb-060897 4 46

                                 - - _ _ _ _ _ = _ - _ _ _ _ - _ _ - _ - _ = _ _ _ _ _ _ _ _ - _ - _ _ _ _ _ _ _ _ - _ _ - _ _ - _ . _ _

FACILITY SCALING REPORT JO Comparing the dimensions presented in Table 4-20 with their counterparts in the full-scale AP600 O yields the following ratios: N g4.00125 (i.e.,1:797.5) (4-191) { 1 dR=2.78 (4-192) (dh.c)R =4.28 ' (4-193) 1 l Also, taking the ratio of Equation 4-187 and substituting the numerical values from Table 4-18 yields the following-  ! l G= g - * . 4.5 (4-194) A I 6 cR s The heat flux hot channel factor (Fq) for the model is 1.932. A factor of 2.4 was used to represent a l full-scale model. Thus: i .(- (F)g q 4.805 (4-195) l , Substituting the numerical values into Equations 4-186 and 4-189 and solving Equation 4185 for the I-range of fluid and flow conditions to be tested in the model yields the graphs presented in

Figures 4-17 through 4-20. - It is observed that in general, L licHF,R= 1.0 (4-196)

, Ihus, the critical heat flux similarity criteria is satisfied. This indicates that if q"cRrr is exceeded during the AP600 tests, it would also be exceeded in the full-scale AP600 hot channel. It should be noted, however, that this similarity is applicable to the specific CHF mechanism modeled by Katto.(17) Core Power Distribution Having developed a core geometry, which satisfies the core heat transfer similarity criteria presented in the previous section, a detailed description of the axial and radial power profiles can now be presented. 36siw-4tnan:n-060897 4 FACILITY SCALING REPORT ne hot channel power peaking factor (F q) was obtained by multiplying the radial power peaking factor (FRadial) by the axial power peaking factor (F3xig). Dat is: Fq=FRadial Fgxi,j (4-197) For the AP600, FRadial = Faxig = 1.55. Therefore, a full-scale PWR Fqof 2.4 was used. Model Axial Power Profile The axial power profile for the model heater rods has been selected so that two-thirds of the core power is generated by the top half of the rods. This approach provides added conservatism with respect to upper core heat-up and the approach to core uncovery. The following general equation was ! used to develop the power profile:

                                                             ' '                       r                                         'a q '(z) =C *_                                 1                   _Z._                                                                       (4-198) t tC jg                                          tC            j where:
 ~

q'(z) = linear power as a function of axial elevation, z i C = normalization constant fc = heated length f a = distribution exponent f The normalization constant (C) is found by requiring the following: le r 'r 'a l I f 9 ROD o ,I

                                                       .z,_

cj s g__z-I,c dz = I C (4-199) where qgon is the heater rod power. Performing this integration and solving for C yields the following: C= ROD 1 _ l ~ -I (4-200)  ; fe .(a +1) (a +2) . O 368Iw-4tnon:Ib-060897 4 4g

FACILTTY SCALING REPORT

 ;       Expressing the average linear power (qly ) as qROD c/I further simplifies Equation 4-200:
                                                                                                                    ~

I - I C = q[y, (4-201)

                                                             ,(a+1)                           (a +2),

The exponent a is found by requiring that the peak linear power occur at the position z = zygx. Thus: 9 (*) =0 at z =zygx (4-202)

       ' Applying the requirement of Equation 4-202 to Equation 4-198 yields an equation for the exponent a -                                                                    !

in terms of known parameters. f 1 f  % MAX c (4-203) a= 1 Ic , ZMAX, ( < , For the model, i, is 36 inches (91.44 cm); and zygx is 27 inches (68.58 cm), which is the center of j the top half of the rods. Thus, a = 1/3, and the normalization constant is given by the following: C =3.11 q[y, . (4-204) l ! Substituting the values of a and C into Equation 4-198 yields the final expression for the linear power  ; L profile: i

                                                                   ' '                                                1/3 q /(z) =3.11 ggy,       1                              1           .z                                                       (4-205) g  Cs                                         C V[h 3681waa.=:itso60897 -                                 4 49

FACILITY SCALING REPORT This axial power profile is presented in Figure 4-21. Table 4-21 presents the fraction of the total rod power generated within different segments of the heater rod, assuming six equal segments. The axial peaking factor (Fgxi,3) is found by evaluating Equation 4 205 at z33rx and dividing by q'gy,. His yields the following: Fgxi,3 =[ ]a.c (4-206) Radial Power Profile Having established an axial power peaking factor (Fgx;,3 = [ ]*C) and a total hot channel peaking factor (Fq = [ ]a.c; as required to satisfy IlCHF,R = 1), the radial power peaking facter (FRadial) IS found to be 131. He hot channel rod power is given by the following: 9HC. rod =FRadial 9 Ave (4-207) He model will have two radial power zones, with each zone containing 24 rods. Figure 4-22 shows the radial power zones for the APEX core. Two-Phase Drift Velocity To evaluate the drift flux number presented in Table 4-10, an equation for the drift velocity (vg) is required. He following equation, developed by Ishii will be used:(38) l/4

                                                                                                                                                                               ' 08 P '

vg =0.2[1 -{pg ,/ph)i 3II +X(AP/Pgs)]u c+1.4 (4-208) Ph, Substituting this equation into the drift flux number yields the following: Ap

• ITNd =0.2[1 -(pp /ph) 3[I +X(AP Pgs)3 / + (4-209) c

                                                                                                                                                                                , Pu ,

when the local slip (the second term on the right side of Equation 4-208) is small relative to the distributed slip (the first term on the right side of Equation 4-208), the Drift number criterion is automatically satisfied by invoking fluid property similitude. Kocamustafaogullari and Ishii point out 3681w-4a.non:lb-071897 4 50

l I FACILITY SCALING REPORT l ^ !/T that even if the local slip dominates the slip behavior, the distortion of the velocity will introduce O limited changes in the void-quality relation."(19) To examine the relative effects of local and distributed drift velocity, Equation 4-209 shall be divided into two parts as follows:

                                                                                               .dist = 0.2 1yps/g pts)% 1 +x(Ap/pg ,)

II Nd (4-210) l and 1.4 ' ogAp * (4-211) l t- 4'd. local " ,PsT*2t , l Core Two-Phase Friction and Loss Coefficients Appropriate orifices will be used in the core to assure that the pressure drop across the core is i A simulated under two-phase conditions. Thus, the friction number criteria will be satisfied. That is, for

  .k                                         the core:

Ilp,=1 (4-212) 4.3 Primary Loop Design Specifications I:

                                           . Sections 4.1 and 4.2 present detailed scaling analyses for single-phase and two-phase natural circulation in closed loops. These analyses resulted in specific scaling ratios that must be satisfied in order to simulate the transfer phenomena of interest. For steady-state conditions with property similitude, the loop scaling ratios presented in Table 4-22 have been implemented for single-phase and two-phase natural circulation flow. The primary loop design specifications are presented in Table 4-23.

l

4.4 Evaluat!on of the Core Processes Specific Frequencies, Characteristic Time Ratios, and Scaling Distortions To assess the relative importance of the various natural circulation processes occurring in the core during single-phase or two-phase natural circulation, the specific frequencies and characteristic time ratios presented in Tables 4-4,4 5,4-10, and 4-11 must be evaluated for both the APEX model and 3681w.4a.non

ll>060897 4-51 .

t

FACILITY SCALING REPORT the full-scale AP600. A comparison of these values reveals which of the processes are dominant and to what extent the proposed APEX geometry and operating conditions introduce scaling distortions. 4.4.1 Evaluation of Single-Phase Natural Circulation Core Transport Processes (Pressure Scaled) Table 4-24 presents the numerical values for the core fluid residence time, the specific frequencies, the characteristic time ratios, and the distortion factors for the core transport processes that occur under single-phase natural circulation conditions. The calculations imposed the requirement of isochronicity. For purposes of calculation, the APEX operating conditions were assurned to be 385 psia with an average fluid temperature of 420 degrees Fahrenheit and the corresponding AP600 conditions were 2250 psia with an average fluid temperature of 600 degrees Fahrenheit. For these operating conditions and geometry, the residence time of the fluid in the core is approximately 13 seconds in both the APEX and the AP600; thus preserving the requirement of isochronicity. The Reynolds number and Prandtl number are also preserved under these conditions. The specific frequencies listed in Table 4-24 indicate that the dominant processes are characterized by the Richardson number and the Friction number. The least important processes are characterized by the Heat Source Number and the Stanton Number. "Iherefore, matching the exact core heat transfer regime is not important to the primary driving mechanism of single-phase natural circulation. Because scaling distortions in the less important processes do not impact natural circulation, their distortion factors need not be evaluated. Table 4-24 indicates that the Richardson number is identically satisfied for the geometry and operating conditions described. The core Friction number can be adjusted through the use of orifices to match the AP600 value. 4.4.2 Evaluation of Single Phase Natural Circulation Core Transport Processes (Fluid Property Similitude) Table 4-25 presents the numerical values for the core fluid residence time, the specific frequencies, the characteristic time ratios, and the distortion factor for the core transport processes that occur under single-phase natural circulation conditions assuming fluid property similitude is preserved. The calculations imposed the requirement of isochronicity. For purposes of calculation, the APEX and AP600 operating conditions were assumed to be 385 psia with an average fluid temperature of 420 degrees Fahrenheit. For these operating conditions and geometry, the residence time of the fluid in the core is approximately 13 seconds in both the APEX and the AP600 thus preserving the requirement of isochronicity. The Reynolds number and Prandtl number are also presened under these conditions. Yhe specific frequencies listed in Table 4-25 , indicate that the dominant processes are characterized by the Richardson number and the Friction number. The least important processes are characterized by the Heat Source number and the Stanton number. As with the previous case, matching the exact core heat transfer regime is not important to single-phase natural circulation. 368i Amn:Ib-060897 4-$2

FACILITY SCALING REPORT l I (\ Because scaling distonions in the less important processes do not impact natural circulation, their distortion factors need not be evaluated. As with the pressure scaled case, Table 4-25 indicates that the Richardson number is identically satisfied for t'.e geometry and operating conditions described. The core Friction number can be adjusted through the use of odfices to match the AP600 value. 4.4.3 Evaluation of Two-Phase Natural Circulation Core Transport Processes (Pressure Scaled) Table 4-26 presents the numerical values for the core fluid residence time, the specific frequencies, the characteristic time ratios, and the distortion factor for the core transport processes that occur under pressure-scaled, two-phase natural circulation conditions. The fluid velocity at the core inlet was calculated using Equation 4-99. He calculations assume saturated conditions in the primary loop with an APEX reference pressure of 320 psia and a corresponding AP600 reference pressure of 1080 psia. For these operating conditions and geometry, the residence time of the fluid in the APEX core is i approximately 4.4 seconds. For the AP600 core, the fluid residence time is approximately 8.7 seconds, thus maintaining a one-half time scale. Void fraction similitude at the core exit is preserved. j The specific frequencies listed in Table 4-26 indicate that the dominant processes are characterized by the Friction number, the Enthalpy number and the Drift Velocity number. The least important processes are characterized by the Phase Change number, the Density number, and the Froude number. Table 4-26 indicates that Friction number is preserved because it is an adjustable parameter. The Drift Velocity number is divided into a local and distributed drift velocity. The distributed velocity is l reasonably modeled in the facility having a distortion factor of 4 percent. The local drift velocity is not preserved. However, this will not significantly affect the void-quality distribution. The Froude number is also preserved. The largest distortion occurs in the Enthalpy number. This is expected because it is a ratio of fluid property groups. This distortion is necessary to preserve the proper void distribution in the core. l Although the core transport processes are not fully preserved in a pressure scaled facility, the void I fraction profile and loop timing are preserved; thus providing the proper fluid compliance in the core, which is important to overall system behavior. 4.4.4 Evaluation of Two Phase Natural Circulation Core Transport Processes (Fluid Property Similarity) Table 4-27 presents the numerical values for the core fluid residence time, the specific frequencies, the characteristic time ratios, and the distortion factors for the core transport processes that occur under two-phase natural circulation conditions at low AP600 pressure when fluid property similarity exists. t The fluid velocity at the core inlet was calculated using Equation 4-99. The calculations assume saturated conditions in the primary loop with an APEX and AP600 pressure of 320 psia. For these f 3681w-4a.non:lb.060897 4-53

FACILITY SCALING REPORT operating conditions and oamnehy, the residence time of the fluid in the APEX core is 3.75 seconds. For the AP600 se the fluid residence dme is 7.5 seconds, thus maintaining a one-half time scale. Void fraction similitude at the core exit is preserved. The specific frequencies listed in Table 4-27 indicate that the dominant processes are characterized by the Friction number, the Enthalpy number, and the Drift Flux number. The least important processes  ! are characterized by the Phase Charge number, the Density number, and the Froude number. Table 4-27 indicates that the Friction number is preserved because it is an adjustable parameter. The Enthalpy number is preserved because of fluid property similarity. The Drift Velocity number is divided into a local and distributed drift velocity. The distributed velocity is preserved in APEX. The local drift velocity is not preserved. However, this will not significantly affect the void-quality distribution. The Froude number is also preserved. 4.5 Conclusions This section implements the H2TS methodology to obtain the primary loop and core scaling criteria for single-phase and two-phase closed loop natural circulation. The important core transport processes are identified and scaling distortions evaluated. For single-phase natural circulation, the dominant transport processes are characterized by the Richardson number and the Friction number. An assessment of the scaling distortions indicates that the dominant processes can be adequately simulated in the APEX test facility. For two-phase natural circulation, the dominant transport processes are characterized by the Enthalpy number, the Friction number and the Drift Velocity number. An assessment of the scaling distortions indicates that the core transport processes are not fully preserved under pressure-scaled conditions. In particular, under pressure-scaled conditions, the enthalpy rise across the core is distorted in order to preserve the vapor void fraction at the core exit. However, because the void fraction profile and loop timing are preserved, the proper fluid compliance in the core is achieved, which is important to overall system behavior. When fluid property similitude exists, the dominant core transport processes are preserved with the exception of the local drift velocity. However, this will not significantly affect the void-quality distribution. In conclusion, the scaling analysis presented in this section indicates that single-phase and two-phase  ; natural circulation data of sufficient quality to benchmark computer codes can be obtained with the  ; APEX test facility. l l O 36siw-4a.non:ib-060897 4-54

FACILITY SCALING REPORT ( 4.6 References i 1. Zuber, N., " Appendix D: A Hierarchical, Two-Tiered Scaling Analysis," An Integrated Structure

l. and Scaling Methodologyfor Severe Accident TechnicalIssue Resolution, U.S. Nuclear l Regulatory Commission, Washington, DC 20555, NUREG/CR-5809, November 1991.

2. Incropera, F.P., and D.P. DeWitt, Fundamentals of Heat and Mass Transfer, 3rd Edition, Joha l

l Wiley and Sons Publishing, New York,1990. j 3. Metais, B., and E.R.G. Eckert, " Forced, Mixed and Free Convection Regimes," Journal of Heat Transfer, ASME, New York, May 1964. [ 4. Welty, J.R., C.E. Wicks, and R.E. Wilson, Fundamentals of Momentum, Heat and Mass Transfer, ! John Wiley and Sons, New York,1984. l [ 5. Ishii, M., and I. Kataoka, " Scaling Criteria for LWRs Under Single-Phase and Two-Phase Natural )- Circulation," Proceedings of the Joint NRC/ANS Meeting on Basic Thermal-Hydraulic Mechanism ! in LWR Analysis, NUREG/CP-0043, Bethesda, MD, September 14-15, 1982.

6. Schroeder, M., Fractals, Chaos, Power Iows, W.H. Freeman and Company,1991.

7. Reyes, J.N., Scaling Single State Variable Catastrophe Functions: An Application to Two-Phase Natural Circulation, Nuclear Engineering and Design 151, pp. 41-48,1994.

( 8.' Kocamustafaogullari, G., and M. Ishii, " Reduced Pressure and Fluid Scaling Laws for Two-Phase l Flow Loop," Argonne National Laboratory, ANL-86-19, NUREG/CR-4585, April 1986.

9. Ishii, M., and I. Kataoka, " Similarity and Scaling Criteria for LWR's Under Single-Phase and Two-Phase Natural Circulation," Argonne National Laboratory, ANL-83 32, NUREG/CR 3267, March 1983.

10. Schwartzbeck, R.K., and G. Kocamustafaogullari, " Similarity Requirements for Two-Phase Flow Pattern Transitions," Nuclear Engineering and Design,116, pp 135-147,1989.

11. Taitel, Y., and A.E. Dukler, "A Model for Predicting Flow Regime Transitions in Horizontal and Near Horizontal Gas-Liquid Flow," AIChE Journal, Vol. 22, No.1, pp. 47-54, January 1976.

12. Modro, S.M., et al., Evaluation of Scaled Integra! Test Facility Conceptsfor the AP600, Idaho National Engineering Laboratory, SMM-27-91, Transmitted to USNRC, July 9,1991.

13. Bankoff, S.G., and S.C. Lee, A Critical Review of the Flooding Literature, Northwestern University, NUREG/CR-3060, Febmary 1983.

l 36stw-4a.non:ib-060897 4-55

                                                                                                                                                                                  --____-.________J

FACILITY SCALING REPORT

14. Hochreiter, L.E., FLECHT SEASET Program Final Report, NUREG/CR-4167, EPRI NP-4112, Prepared for the U.S. Nuclear Regulator Commission, November 1985.

15. Ohnuki, A., H. Adachi, and Y. Murao, " Scaling Effects on Countercurrent Gas-Liquid Flow in Horizontal Tube Connected to Inclined Riser," Japan Atomic Energy Research Institute, Tokai-mura, Ibaraki Pref., 319-11, Japan, (0292) 82-5275.

16. Griffth, P., J.F. Pearson, and R.J. Lepkowski, " Critical Heat Flux During a Loss-of-Coolant Accident," Nuclear Safety, Vol 18 (3), p. 298,1977.

17. Mishima, K., and M. Ishii, " Critical Heat Flux Experiments Under Low Flow Conditions in a Vertical Annulus," NUREG/CR-2647, ANL-82-6,1982.

18. Ishii, M., "One-Dimensional Drift-Flux Model and Constitutive Equations for Relative Motion Between Phases in Various Two-Phase Flow Regimes," Argonne National Laboratory Report, ANL-77-47,1977.

19. Kocamustafaogullari, G., and M. Ishii, " Scaling Criteria for Two-Phase Flow Natural and Forced Convection Loop and Their Application to Conceptual 2x4 Simulation Loop Design," Argonne National Laboratory, ANL-83-61, NUREG/CR-3420, May 1983.

O O 368Iw-4a non:Ib-060897 4 56

FACILITY SCALING REPORT G TABLE 41 STEADY STATE LOOP BALANCE EQUATIONS FOR SINGLE-PHASE NATURAL CIRCULATION FLOW j l Mass (at every cross-section): l til " pg j jua 8pggc u3 Constant (4l) l 1 Momentum: ' l i 2 r i r 39

                                                                                     " " ,4               f    +K        .1            =

7ogp,o fg -Tc ),L ih N

                                                                                          '        1
                                                                                                        ,    h     ,js ag, Energy:

Pso uga,Cp ,ofg -T c), = q,o (43) l O 1 Core Inlet Fluid Velocity: i e ,10 PT9s 8 um " (4~4)

                                                                                                                     ,ca Cp,p,FT, o Component Fluid Velocity:

P ' 10 < ' 2 T9slthg ac i u6* - (4-5) s ac Cp ,p, FT . , ( ";, (~ ( 3681w-4non:lb-060897 4-57 L. -- _ _ _ _ _ _ _ _ _ _ _ - - _ _ _ _ _ _ _ _ _ - _ - - _ _ _ _ - _ - _ - . _ _ _ _ _ - _ _ _ _ _ .

FACILITY SCALING REPORT TABLE 4 2 SINGLE PHASE CONSTITUENT LEVEL SCALING ANALYSIS: CONTROL VOLUME BALANCE EQUATIONS FOR TIIE CORE Mass: (4~7) (P:V )"A (P Q] Momentum: P'" 8

                                                                                                                                                                                                          +K                       (4-8)

(Pa u ,V,)=A [p,ug,]+pTgpgria, Energy: (P Cy ,(T,)V,)=A [p,C ,Tg,]+H,A,(r,-T,)l Boundary p Solid Energy Equation: PsCy ,V, (r,) = H, A,(r,-T,) ( x1, (4-10) O 3681w-4 anon:lb-060897 4 58 )

FACILITY SCALING REPORT lA

                                             =======-

TABLE 4-3 SINGLE PHASE CONSTITUENT LEVEL SCALING ANALYSIS: NON DIMENSIONALIZED BALANCE EQUATIONS FOR THE CORE f Mass: i

                                                                                                $,0     kl I    *6 kQ+  l i (4-it) l                                                  Momentum:                                                                                                  ,

t to{(piu,N,i - A piu,*Qi +%(sipiaTI (4-13)

-I Pt "t. *Q t . H

                                                                                                                               +K l                                                                                                            t 2    ,  ,d h    ,

Energy: I r , PtCy *, T,' V,* = A p,*C,*,T,*Q,'

To dl ' + 11, H,'(r,-T,)* l Boundary(4-16) l Fuel Rod Energy Equation

T Lo Is if ,H,*(r,-T,)} Boundary bl98 l i I t l I l f

  'J                                                                                                                                                          j I

a 3681w-4a.non:lb-060897 4-59 1

l FACILITY SCALING REPORT TABLE 4-4 SINGLE PIIASE CONSTITUENT LEVEL SCALING ANALYSIS: RESIDENCE TIMES AND CllARACTERISTIC TIME RATIOS FOR TIIE CORE j

                                                                                                                                                                                                                                                                                                                                                                )

Liquid Residence Time: V t,2 g (4-12) Q,,, Solid Heat Transfer Specific Frequency:

                                                                                                                                                                                                           =1=                                ' ^5                                                                                       (4-20) to

t,,, p,Cy ,V, Characteristic Time Ratios: o Tig ; = OT,ogAT t Richardson Number (4-14) ug fi flp = +K Friction Number (4-15) dh

                                  ,o H,',A*

Il Stanton Number (4-18) sT = pg CpgQg Liquid Heat Source Number (4-25) I\, = P,C y,ugAT,V, Specific Heat Ratio: CPt.o (4,;7) 7to " Cg , O 3681w.4uoirib-060897 4 60

FACILrrY SCALING REPORT I t-t l TABLE 4-5 I SINGLE PHASE CONSTITUENT LEVEL SCALING ANALYSIS: PROCESS SPECIFIC FREQUENCIES FOR THE CORE i i Process Specific Frequencies: i

                                                                                                                                                         ,     1/3 ETo89soFT (4 30)     l 4pgCpgac Lj, r         ,

c)p = V E +K (4-31) b ,d h ,, w,, . 5 ^' (4-32) l PbCpgV g l 1/3 l h t "qi , 2 To9so48(PbC pe ae)2 (4-33) L F(PsC,., V ,)3 T l l l l 1 1

           \.

l O , 3681w-4a.non:Ib-060897 4 61 l L__.___.-____-__-___-_____. - . _ _ _ _ _ _ _ _ _ .

FACILITY SCALING REPORT TABLE 4-6 STEADY-STATE, SINGLE PIIASE NATURAL CIRCULATION LOOP SCALING RATIOS Flow Area Scaling Ratio: 8e

                                                                                           ,;                                                                                           (4 43)

• IsR Thermal Center Length Scaling Ratio:

(L g)g =fg (M) Loop Friction Number Ratio: (U7)g =1 (4-45) Loop Fluid Velocity Scaling Ratio: r 'IS ' ' 1/3

                                                                                                  '                                                                                     I4'40) uk*

pgCpg g a C sR Fluid Residence Time Scaling Ratio: gg. Y M ( In WO OTo9so .R Heat Source Ratio:

                                                                                                         *I                                                                             (4'49) htR

• PsCV n 8 R Stanton Number Ratio:

1/3 3 3 kA (4-50) Il ST,R = I (Re,c)08 f (Pr)0 #f =1 dh@To9so (PbCpga c)2 0 3681w-da.non:lb-060897 4 62

FACILITY SCALING REPORT 5 TABLE 4 7 STEADY STATE, SINGLE PHASE NATURAL CIRCULATION LOOP SCALING RATIOS: ISOCHRONICITY Fluid Residence Time Scaling Ratio: Tg =1 (4-51) Fluid Velocity Scaling Ratio: UR "IR (4-52) Power Scaling Ratio: l gg-

• ag[g (W) s To sR Heat Source Ratio:

s g,,g = Pb mb ,3 (4,49) PsCy ,V, .R Stanton Number Scaling Ratio:

r. ,

r , IIST ,R " - (Re)08 R (Pr)U[ = 1 (4-54)

                                              ,Id ahCe R            fbCphsR Reynold's Number Ratio:

Eb h (4-55) (Re)R.  ! s Eb ,R I I 1 l 368Iw-4amrtIt4)60897 4 63 I L

FACILITY SCALING REPORT TABLE 4-8 CONSTITUENT LEVEL SCALING ANALYSIS: TWO PHASE MIXTURE CONTROL VOLUME BALANCE EQUATIONS FOR Tile CORL'. AS DERIVED IN APPENDIX A Mixture Mass: (PrVr) " A (Pf9f) - A (nacApvg) (4-56) Mixture Momentum: r , 2 dt (PruV g r) "^ (prugr) + AP8"V t +6 (1-c)pf

                                                                                                          ,                  ,             (4-57)

_ Pr ugQg rg 2 3+K , Mixture Enthalpic Energy Equation: (Pr(hg)Vg )=A(pgh gQg) +Hsf A,(T,-T g) 3,

                                                                                                         +A              ahv cg              (4-58)

Solid Energy Equation: PsCy ,V, ,) 41,gA,(T, -T g) g,

                                                                                                              +q,                          (4-59) l O'
           %8Iwaa.non:1b-060897                                                         4-(>4

i t l FACILITY SCALING REPORT f ,

                                                                                                                                                                         }
                                   ~

(t \' TABLE 4 9 CONSTITUENT LEVEL SCALING ANALYSIS: TWO PHASE MIXTURE NON.DIMENSIONALIZED BALANCE EQUATIONS FOR THE CORE l Mixture Mass: tr,o (pgYd = A fp*Q[ -Ihd A "*A P +v (4-60) Mixture Momentum: s. 2 Tg,, (p[ugY g) =6(p*u[Q[) g + gat [+% .d 0 yp ff P[ugQ['f, 2 Mixture Enthalpic Energy Equation: A [ Vd=A(pg1[Qd+TIST H,*g,-T f)* h r

                                                             'T,oE(Pr f
                                                                                                           '               l Boundary
                                                                                                    ,                ,.                                           (4-67) 8'
                                                                                       +Illhd pII A          " hgvg r    Pf          ,

Fuel Rod Energy Equation: Tf ,o EIs tf -H,*g,-Tg)* , l Boundary +%ch9s (4-77) l t

 .\ f 3681w-4a.non:lN97                                             4-65

FACILITY SCALING REPORT TABLE 410 CONSTITUENT LEVEL SCALING ANALYSIS: TWO.PIIASE MIXTURE RESIDENCE TIMES AND CIIARACTERISTIC TIME RATIOS Mixture Residence Time: V T,,=_fd f (4-61) O f.o Solid Structure Specific Frequency: I "5f * ^ 5 a3,,, . = (4-73) T,,, p,Cy ,V, Characteristic Time Ratios: th = 8'uf,,pf,, Drift Flux Number (4-62) 2 I- I* I7p, = Froude Number (4-64) 08 fAbo 85- ** I = Density Number (4-65) a,(1 -<ro) (Apo) f  % Il p= E +K Friction Number (4-66) dh Hsf o^ 5 Il Stanton Number (4-68) sT = Ef.o CUysaf o g h Q = g,(1-ct),a,Ap Enthalpy Number (4-69) OhoPl.o Phase Change Number (4-78) hh = p, A xuf,, Ah, O 368Iw-4 anon:Ib-060897 4 66

FACILITY SCALING REPORT 5 \ V TABLE 411 TWO PHASE CONSTITUENT LEVEL SCALING ANALYSIS: PROCESS SPECIFIC FREQUENCIES FOR THE CORE gjo o o "Nd " (4-82) Pf.of uto Pr.o e,= p (4-83) UoE OPo PasoPtsou r,, e. Co(1 -a,)(Apo)2 g I' 0)p = N+K (4 85) I Ah ,, H,f,, A, U O ST = Pf.oCy , Vf ,, (4-86)

                                                                                   ,   froh     o)Uo OPou f a C                         (4-87) 9soI f

U pch = (4-88)

                                                                                                                                                     )

I O 368IwAmn:it>460897 4 67 L

FACILITY SCALING REPORT TABLE 412 STEADY STATE LOOP BALANCE AND STATE EQUATIONS FOR TWO PIIASE NATURAL CIRCULATION FLOW Loop Mass (at every cross-section): piu;a;rpeua c, (4-90) Loop Momentum:

                                                    '2          r       '  '
                                                                                 '2         r          '2
                                                    **                                                 *                        (4-92) hE       f1 +K       .1    +d E f1+K        .

W,-pf)Lgg 2 pg TP , d h , , a,i pgSP, d b , j a; , Loop Energy: peu ,ac(hf -h,) 9, (4-94) Equilibrium Vapor Quality: hf-h6 x* = (4-96) l hg Average Two-Phase Fluid Density; b pg= ' ' (4-93) AP 1 +x,

                                                                                          , Pgs,

) ( O 368Iw-4-1m897 4 68

FACILITY SCALING REPORT TABLE 413 EQUATION FOR CORE INLET FLUID VELOCITY UNDER TWO-PIIASE NATURAL CIRCULATION CONDITIONS 3 uc 4,u,2y ueye=0 (4-M where: q,Ap (F T+FTP)Pgsh ig-2ApF TPhsub (499a)

                                          "cPis FTP s h,28 (F +F7p)p ,h hig rubop +F7p(Ap)2h,2
                                                           - T          g                         ub,
                                        , Ap              ApFTP9s +2Legp,2ahc      2 subPgshig (4 99b) 22      2 a, p  p p ss h[-(F +FTP)Pgs T           h ,,h A sub p +FTP(Ap) hsub, 1                                       h
                                        ,                        2Lih84sAPPss e acpe, p  2 g

pg hd-(F 7+FTP)Pgshh ,g subo p +FTP(Ap) hsub, and

                                                                 ,        ,,,2 FSP "     Il *N         c (4'l00)

SP , d h s ; , aj ,

                                                                                      ,2 Fp=E 7               +K (4-101)

F7=FSP + FTP (4-102) O 3681w-4a.non:lb-060897 4 69

FACILITY SCALING REPORT TABLE 414 STEADY STATE, TWO-PIIASE NATURAL CIRCULATION LOOP SCALING RATIOS FOR SATURATED CONDITIONS Vapor Quality Scaling Ratio: xp =l (4-115) iPgs, R Power Scaling Ratio: In PtsP shig (4-116) 9 R "UR(Lth)R i AP ,R Fluid Velocity Scaling Ratio: UR dbth) Fluid Residence Time Scaling Ratio: Ta= (4-118) (Leh)R Mass Flow Rate Scaling Ratio: REQUIREMENTS

• Kinematic Similarity:

                                                               "C
                                                                       =1                          (4-107)
                                                             $   ieR

• Friction Number Ratio:

(FSP)R = 1 (4-108) (F 7p)g = 1 (4-109) (F T)R = 1 (4-110) O' 36stwaa non:ltawos97 4-70

FACILITY SCALING REPORT ( TABLE 415 STEADY. STATE, TWO-PHASE NATURAL CIRCULATION LOOP SCALING RATIOS: (WITH PROPERTY SIMILITUDE) l l Vapor Quality Scaling Ratio: xg =l (4-130) Power Scaling Ratio: t/2 9R =aR(b)R (4-131) Fluid Velocity Scaling Ratio: uRilt (4-132) Fluid Residence Time Scaling Ratio: IR (Ig)f (4-133) Mass Flow Rate Scaling Ratio: rngqLth ) ag (4-134) , v REQUIREMENTS

• Property Similitude
• Flow Area Ratio:

                                                                                                                                                  =l                       (4 121) 6   isR

• Friction Number Ratio:

(FTP)R "I ("I ) (P T)R =1 (4-124)

 ,r x.

( l

  %/

I 3681w-4a.non:11,-060897 4-71

FACILITY SCALING REPORT TABLE 4-16 SYSTEM SCALING RATIOS FOR STEADY-STATE NATURAL CIRCULATION WITH SINGLE PHASE AND TWO PHASE FLOW REGIONS (MATERIAL PROPERTY SIMILITUDE AND FIXED LENGTH RATIO) Power Scaling Ratio: In P6Pgshtg (4-138) 9R,sp 9R,tp=aRR I g Fluid Velocity Scaling Ratio: r ' l/3 UR.sp *R,tp } Fluid Residence Time Scaling Ratio:

                                                                                                                                           ?          '
                                                                                                                                                          -in
                                                                                                                                                                                                                                                                       ~

TR,sp "R#p O REQUIREMENTS

• Flow Area Ratio:

                                                                                                                                       .l         =1                                                                                                            (4-141)
                                                                                                                                      ,a,gc
  = Friction Number Ratio:

FSP,R =1 (4-142) FTP R =1 (4-143) FT,R "I I4~I44) O 368Iw-4a.non:1b-060897 4 72

                                                                                                                                                                   .i

FACILITY SCALING REPORT O

TABLE 417 SYSTEM SCALI';O RATIOS FOR STEADY-STATE NATURAL CIRCULATION WITH SINGLE PHASE AND TWO-PHASE FLOW REGIONS (PROPERTY SIMILITUDE AND FIXED LENGTH RATIO) Power Scaling Ratio: la 9R,sp 9R.tp=aR IR (4-145) l Fluid Velocity Scaling Ratio: IC UR.sp N R,tp*I R (4-146) Fluid Residence Time Scaling Ratio: 1/2 TR,sp "R,tp*I R I l l (4-147) { Mass Flow Rate Scaling Ratio: NRsp *b R.tp *I RaR (4-148) l REQUIREMENTS

                                                                                                                                                                    \

1 i

                   = Property Similitude
                   = Flow Area Ratio:
                                                                                                                               =l                       (4-149) a s   CeR t

e Friction Number Ratio: FSP.R =1 (4-150) l FTP.R =1 (4-151) l F.R=1 T (4-152) O V 3681w-4tnon:IM)60897 4 73

I FACILITY SCALING REPORT TABLE 418

SUMMARY

OF SYSTEM SCALING RESULTS FOR THE 1/4 LENGTH SCALE MODEL PRIMARY LOOP (PROPERTY SIMILITUDE) Geometry Length Scaling Ratio: 1:4 System Diameter Scaling Ratio: 1:6.931 Area Scaling Ratio: 1:48.04 Volume Scaling Ratio: 1:192.16 Flow Velocity Scaling Ratio: 1:2 Mass Flow Rate Scaling Ratio: 1:96.08 Residence Time Time Scaling Ratio: 1:2 Power Power Scaling Ratio: 1:96.08 Power Density Scaling Ratio: 2:1 MODEL POWER REQUIREMENTS Percent of Total Power 5% 4% 3% 2% AP600 Decay Power (MW): 97.00 77.60 58.20 38.80 APEX Power (kW): 1009.58 807.66 605.75 403.83 O M8Iw-4a.non:Ib-060897 4 74

FACILITY SCALING REPORT Q TABLE 419 TWO PHASE FLOW TRANSITIONS IN THE LOOP LEGS (FLUID PROPERTY SIMILITUDE) AND PRESSURIZER SURGE LINE Hot and Cold Lees: Stratified-Smooth / Stratified Wavy Distortion

                                                                          - Stratified / Intermittent or Annular-Dispersed                                                                    Scaled CCFL                                                                                                              Scaled Intermittent or Dispersed Bubbly / Annular-Dispersed                                                              Scaled Intermittent / Dispersed Bubbly                                                                                Distortion l

Pressurizer Suree Line: Bubbly / Slug Flow Scaled

Slug / Churn Flow Scaled Slug / Annular Flow Possible Distortion TABLE 4 APEX CORE HFPER BUNDLE DLMENSIONS AND POWER l Number of Heater Rods '48 Maximum Power per Rod 15. kW Rod Diameter 2.54 cm 1. in Rod Length 91.44 cm 36.in Rod Surface Area 729.66 cm2 113.1 n2 2

Rod Cross-Sectional Area 5.07 cm 0.785 in 2 Rod Volume 463.6 cm3 28.27 in 3 Total Heater Surface Area 35,023.7 cm2 5,428.8 in 2 Total Heater Cross-Sectional Area 243.36 cm3 37.68 in2 Total Heater Volume 22,252.8 cm 3 1,356.96 in 3 Heater Rod Pitch - 4.01 cm 1.58 in

Pitch / Diameter Ratio 1.58 1.58 Subchannel Flow Area 11.01 cm 2 ' l.711 in2 Hydraulic Diameter 5.52 cm 2.18 in Average Rod Heat Flux 11.54 W/cm2 Maximum Core Power 720. kW Note

Numerical values may not reflect as-built conditions. Refer to Facility Design Report.

i t. . l 36siw-4a.non:1b-060897 4-75

FACILITY SCALING REPORT TABLE 4 21 AXIAL POWER FRACTIONS FOR THE APEX CORE Fraction of Total Integration Band (f0 Rod Power 2.5-3.0 0.19363 2.0-2.5 0.24293 1.5-2.0 0.22485 1.0 - 1.5 0.17981 0.5 - 1.0 0.I1724 . 0-0.5 0.04154 Total 1.00000 O l I O 1 l l 3681w-4a.non:1b-060897 4 76 l

FACILTIY SCALING REPORT j l

                                                                                                                                                                                                               'i I

TABLE 4-22 OSU APEX PRL\lARY LOOP AND CORE SCALING RATIOS l Core: Length Scaling Ratio 1:4 l Rod Diameter Ratio 2.78:1 Hydraulic Diameter Ratio 4.28:1 l Flow Area Ratio 1:48.04 Power Ratio 1:96 l l Velocity Ratio 1:2 Residence Time Scaling Ratio 1:2 Mass Flux Ratio 1:4 I Mass Flow Rate Ratio 1:192.16 Hot and Cold Lees: Length Scaling Ratio 1:5 Pipe Diameter Ratio 1:6.2 n f\ Pipe Flow Area Ratio 1:38.44 Pipe Volume Scaling Ratio 1:192.16 Velocity Ratio 1:2.5 Residence Time Scaling Ratio 1:2 Mass Flow Rate Ratio 1:96.08 Pressurizer Suree Line: t Length Scaling Ratio 1:4 Surge Line Diameter Ratio 1:4

                                        *Other Primary Loon Components:

Length Scaling Ratio 1:4 Diameter Ratio 1:6.931 Flow Area Ratio 1:48.04 Volume Scaling Ratio 1:192.16

                                        *(Not including safety systems and reactor vessel downcomer.)

i

              \j 368Iwaa.non:Ib-060897                                                                                                        4-77

a I l FACILITY SCALING REPORT ) TABI.E 4-23 OSU APEX PRIMARY LOOP AND CORE DESIGN SPECIFICATIONS Core: SI Units En_nlish Units Number of Heater Rods 48 Maximum Power per Rod 15. kW Rod Diameter 2.54 cm 1. in Rod Length 91.44 cm 36. in Rod Surface Area 729.66 cm2 113.1 in 2 Rod Cross-Sectional Area 5.07 cm 2 0.785 in 2 Rod Volume 463.6 cm 3 28.27 in 3 Total Heater Surface Area 35,023.7 cm 2 5,428.8 in 2 Total Heater Cross- 243.36 cm2 37.68 in2 Sectional Area Total Heater Volume 22,252.8 cm 3 1,357.95 in 3 Heater Rod Pitch 4.01 cm 1.58 in Pitch / Diameter Ratio 1.58 1.58 Subchannel Flow Area 11.01 cm2 1.711 in2 Hydraulic Diameter 5.52 cm 2.18 in f 2% Decay Heat Core 403.8 kW Power Radial Power Peaking 1.31 Factor Axial Power Peaking 1.47 Factor Hot Channel Factor 1.93 Hot and Cold Lees: Hot Leg Inside Diameter 12.70 cm 5.00 in Cold Leg Inside Diameter 9.013 cm 3.548 in Pressurizer Surce Line: Surge Line Inside 9.168 cm 3.610 in Diameter Pressurizer Inside 28.30 cm 11.14 in Diameter l Note: Numerical values may not reflect as-built conditions. Refer to Facility Description Report. l O l 1 l l 3681w.4tnon:ib.060897 4-78

FACILITY SCALING REPORT l t l (O*j TABLE 4 24  ; EVALUATION OF SINGLE PHASE ' NATURAL CIRCULATION RESIDENCE TI51ES, l CHARACTERISTIC TIA1E RATIOS, AND SPECIFIC l FREQUENCIES (ISOCHRONICITY, PRESSURE SCALED) APEX AP600 System Pressure: 385 psia 2250 psia i Core Fluid Residence Time: ) Tg,o 13.2 s 13.2 s l l l Core Process Specific Frequencies: l

                                                                                                *mgi                                             3.6 s~3                 3.6 s'l
                                                                                                *m p                                            (Il l.1 s'l              1.3 s'I e,q                                           0.120 s'l               0.%3s'I esT                                           0.012 s'l               0.26 s

Characteristic Time Ratios: l

                                                                                                *Dg;                                           47.5                    47.5

, m-

                                                                                                *Dp                                             (3)l7.2                 17.2 l

H,q 1.58 0.83 HST 0.16 3.43 Distortion Factors (DF): Richardson Number (Hg;) 0% (ITriction, Number (Up): 0%

• Dominant transport process (3) Adjustable parameter l

l i i i

 /O N

I i 3681w-4a.non:Ib4)60897 4-79

FACILITY SCALING REPORT l TABLE 4 25 EVALUATION OF SINGLE PIIASE NATURAL CIRCULATION RESIDENCE TIMES, CHARACTERISTIC TIME RATIOS, AND SPECIFIC FREQUENCIES (ISOCHRONICITY, PROPERTY SIMILITUDE) APEX AP600 System Pressure: 385 psia 385 psia Core Fluid Residence Time: t oo 13.2 s 13.2 s Core Process Specific Frequencies:

                                                           *ogi                                                                                                                                          3.6 s-I                    3.6 s'l
                                                          *my                                                                                                                                             1.3(3)s'I                 1.3 s'l m, q                                                                                                                                         0.120 s~l                  0.06 s'l eT s                                                                                                                                        0.012 s"I                   0.29 s'l Characteristic Time Ratios:
                                                          *fl gi                                                                                                                                    47.5                           47.5
                                                          *fI p                                                                                                                                      17.2(1)                       17.2 U, q                                                                                                                                         1.58                       0.79 FIsT                                                                                                                                        0.16                        3.83 Distortion Factors (DFh Richardson Number (ITg ;)                                                                                                          0%

(IYriction Number (Tip): 0%

• Dominant transport process (I) Adjustable parameter I

l l l l 9 3681w-4a.non:Ib-060897 4-80

l FACILITY SCALING REPORT 1 1 i ( TABLE 4 26 l EVALUATION OF TWO-PHASE NATURAL I CIRCULATION RESIDENCE TIMES, CHARACTERISTIC TIME RATIOS, AND SPECIFIC FREQUENCIES (PRESSURE SCALED) APEX AP600 System Pressure 320 psia 1080 psia Core Fluid Residence Time: tr,o 4.4 s 8.7s Core Process Specific Frequencies: l

                                                          *mh                                         5.7 s'I                                 0.63 s 1
                                                          *mp                                         3.06(3) s                             3.06 s'l

, *mNd 0.64 s'l 0.18 s'I mp 0.14 s'l 0.032 s 3 mpch 0.028 s-3 0.009 s'l l 5.7x10 4 s'I 4 -l mp, 2.97x10 s Characteristic Time Ratios: !(] *U h 25.2 5.5 U *H p 26.7(3) 26.7 l *H Nd 2.80 1.60 U Nd.1 d 1.42 0.76 D Nd. &n 0.81 0.84 Up 0.06 0.28 Upch 0.12 0.07 l U,p 2.5x10-3 2.5x10-3 Distortion Factors (DF): l Enthalpy Number (Gh ): (Distorted to preserve core void fraction) (3Yriction Number (Up): 0% Local Drift Velocity: 86.4 % Distributed Drift Velocity Number: 3.9 %

• Dominant transport process (3) Adjustable Parameter

{ l t 3681w4a.non:Ib-071897 4-81 l l

FACILITY SCALING REPORT l { TABLE 4-27 l EVALUATION OF TWO-PHASE NATURAL l CIRCULATION RESIDENCE TIA1ES, CHARACTERISTIC TIN 1E RATIOS, AND SPECIFIC FREQUENCIES (FLUID PROPERTY SIAIILITUDE) APEX AP600 System Pressure: 320 psia 320 psia Core Fluid Residence Time: t f,, 3.75 s 7.50 s Core Process Sticcific Frequencies:

           *mh                                              9.46 s'l                                                   4.73 s'I
           *mp                                              2.48(l) s~l                                                2.48 s'l
           *%d                                              0.34 s'l                                                   0.11 s'l mp                                            0.14 s'I                                                   0.007 s~l eh pc                                         0.033 s"I                                                  3.15 s'I m, y                                          1.6x10-3s~l                                                8.0x10 4  s'I Characteristic Time Ratios:
           *H h                                            35.56                                                      35.56
           *U p                                            18.62(1)                                                   18.62
           *D Nd                                            1.28                                                       0.84 H Nd. local                                   0.88                                                       0.44 U Nd. dist                                    0.40                                                       0.40               i Up                                            0.05                                                       0.05 Upch                                          0.12                                                      23.68 U, p                                          6.07x10~3                                                  6.07x10-3 Distortion Factors (DF):

Enthalpy Number (Dh): 0% (lYriction Number (Up): 0% Local Drift Velocity: -100 % Distributed Drift Velocity Number: 0%

• Dominant transport process (IJAdjustable Parameter O

3681w-4a.non:lt>.o60897 4-82

FACILITY SCALLNG REPORT l i SINGLE-PHASE NATURAL CIRCULATION PHENOMENA l i l 1r 1r Top Down\ System Bottom-Up/ Process Scaling Scaling LooFMass Flow Rate . Friction and Loss

• Core / Steam Generator LOOP Mass Flow Rate . Heat Transfer Regimes l Heat Transfer O

i 1r 1r l Single Phase Natural Circulation Similarity Criteria l 1r Integrate with Two-Phase Natural Circulation l Similarity Criteria Figure 41 Scaling Analysis Flow Diagram for Single Phase Natural Circulation 3681w4b.non:lb-060897 4-83

FACILITY SCALING REPORT l lI O a el SGB SG A a e n 5  ! 5 5

                                                                  ~                                                                                                                                                               -
                                                                  ! a C    i                           <

O ~~ V ozzIE -

                                                                                                                                                                                                                     . n            -
                                                            .         ,_~                                                                                                                                          _____y 4__g__

j- __ RING ~ PUt1PJ "e ~.~ _~ s

                                                                                                                                                                                                 ~~~               5
                                                                                                                                                                                                   ~-                                       COLD LEG HOT LEG                                                                                                                        g g gg RODS REACTOR CORE s

HOT L10Ul0 l: ;l COLD LIQUID l l l Figure 4 2 Hot and Cold Leg Regions of Single-Phase Natural Circulation Flow Within a PWR O l 368Iw 4b.non:1b 060897 4,g4

s FACILITY SCALING REPORT

                                                                                                                                              .l O

TWO-PHASE NATURAL CIRCULATION l PHENOMENA l l-l' I 1r II Top Down/ Systems Bottom Up/ Process l Scaling Scaling

• Friction & Loss
• Loop Mass Flow Coeffielents I Rates
• Flow Regime Transitions
• Core Heat Transfer . Hot Leg CCFL
• CHF l

ir 1r Two-Phase Natoral Circulation Similarity Criteria f l I ir Single Phase l Natural Integrate i Ciresistion  : Single and Two-Phase  ; Similarity Natural Circulation Criteria ir (. l r Evaluate Scaling l' Distortions l ir Specify Design Specifications for Primary Loop Components m ~ ~ ~ Figure 4-3 Scaling Analysis Flow Diagram for Single Phase Natural Circulation t 3681w-4b.non:lb 060897 4-85

l FACILITY SCALING REPORT O SG A h hSGB e i  ! A: E 2  :  : - 2  :  : E E  :  :  :

:  : 5 i ! 3 C E i

                       !    !                                                 i 5
                                                                                  ~~
                     .                                                      *e e' 5%
                    - _ _      _ _o    _ -       hozzEi _        -

b

                                                                                - -                    4 w ...; . . . aSM.

f *E PUt1P / .L. *. *., *. 3 f ' a .a.*

• COLD LEG HOT LEG iiEITER

                                                           ~
                                                           ~
                                                ~d'o6s-

_ ~::__-_ - k~

                                                   ,                  REACTOR CORE 20         o o oo O 00 L10Ul0     ~ _ ~ ~.

Figure 4-4 Regions of Single-Phase and Two-Phase Natural Circulation Within a PWR O 3681w-4b.non:lb-060897 4 86

l FACILITY SCALING REPORT l

    /
                     ^)

s SCAUNG RATIOSvs. AP600 PRESSURE (Model Pressure = 375 psia)

                                                                                                                                                                                                             ,, o  ..-.                                         ,

1

                                                                                                                                                                                    ----                           SPTime                                       !

l 0.9 SP Velodty

                                                                                                                                                                                -, , * , , , * " ,                 -m-i o OE                                                                     y                                            TP Tr' ne
                                                                                                                                                                                                                   +

0.7 ,',W'- TP Velodty 1 0.6 ' O.5 m,,  :  :  :  :  :  : 0.4 ,'* 03 ... = - ..----. -- o b 400 660 860 10'00 1500 1400 1000 18'00 2000 AP900 PRESSURE (psia) Figure 4-5 Scaling Ratios for Steady State Natural Circulation vs. AP600 System Pressure (Model Pressure = 375 psia) SCAUNG RATIOS vs. AP600 PRESSURE (ModelPressure = 300 psia) 12 ,,, o .. 1.1

                                                                                                                                                                                           . , ,, , e' '        SPTrne l
                                                                                                                                                                         /,,, *~

1 SPvelodty 90 0

                                                                                                                                                     ,',."                                                      TPTrne g OE O.7 TPVM
                                                                                                                             ',, 7" 0.8                   .-

O.5 - ,:  :  :  :  :  :  :  : 0.4 p 03 w.,----*----...... ...m, - 0, 1 )0 400 e60 ado 1000 1200 1400 1EiOO 1800 2000 AP900 PRESSURE (psia) l Figure 4-6 Scaling Ratios for Steady-State Natural Circulation vs. AP600 System Pressure (Model Pressure = 300 psia) (3

    'w) 3681w-4b.non:lb-060897                                                                                                                 4 87

FACILITY SCALING REPORT SCALING RATIOS vs. AP600 PRESSURE (Model Pressure = 200 psia) 12 o , , , , 1.1

                                                                                                                                                                                                                                                                          ",, ^*~,,,

SP Trne I ~-

                                                                                                                                                                                                                                                               ,,,'" ,,,'                  SP Velodty 0.9                                                                                                                                                         , r'                                     -*-
                                                                                                                                                                                                                                                                                           -e-I 0.8                                                                                                           ', ar     ,

TP Velodty 0.7 0.6 y' 05 ('r = = = = = = = 0.4 ' ,,,, - 0.3 "-- o h 400 e60 abo 10'00 1200 1400 itio0 18'00 2000 AP600 PRESSURE (psia) Figure 4-7 Scaling Ratios for Steady-State Natural Circulation vs. AP600 System Pressure

                                                                               @lodel Pressure = 200 psia)

O SCALING RATIOS vs. AP600 PRESSURE (ModelPressure = 100 psia) 1A , . - ..

                                                                                                                                                                                                                                                                       ",,m,,                SPTrne 12                                                                                                                                                                           w                         ..
                                                                                                                                                                                                                                              ,,,"',,,                                         SPVelocity 1                                                                                                                                                                                             -*-

TP11me

                                                                        $0.8-                              ---

TP Velocky g , 0.6 ',/

= = = = = = =  :

0.4  ;' ~' .= , 0.2

                                                                                                                                                        '+--,'*-"-*--W=n                                                                                                                -

h 4bo 060 800 10'00 12'00 1400 1000 1800 2000 AP600 PRESSURE (psia) Figure 4-8 Scaling Ratios for Steady State Natural Circulation vs. AP600 System Pressure Sfodel Pressure = 100 psia) 368Iw-4b.non:Ib-060897 4.gg , l l

FACILITY SCALING REPORT ("'s d MODEL POWER vs. AP600 PRESSURE (Model Pressure = 375 psia) 1200 5% DemyHeat 1000- -+-- r 800-^\ \ 1-~ 3%%M 000

                                                                    -(                                         -.-

2% Decay Heat b"

• XL&% hN -

                                                                                                     ~
                                                                                                     =   .-u b 400 000 860 10'00 1200 1400 1800 1500 2000 AP900 PRESSURE (pein)

Figure 4-9 Model Power Requirements vs. AP600 System Pressure for Steady-State Natural Circulation (Model Pressure = 375 psia) MODEL POWER vs. AP600 PRESSURE (Model Pressure = 300 psia) 1200 S% DecayHest 1000 -+- 4% DecayHeat {800 ^T -m-3% Decay Heat

                                                                                                                 -e-800 2% DecayHeat 0"       ~

KM

• hh_ T 5 I

                                                                                                   =   =      n b 400 8b0 86010'001200140010001500 2000 AP900 PRESSURE (poia)

Figure 4-10 Model Power Requirements vs. AP600 System Pressure for Steady State q Natural Circulation (Model Pressure = 300 psia)

     ]

3681w-46.non lb40897 4 89

FACILITY SCALING REPORT 1

                                                                                                                                                                                                                     )

f i MODELPOWERvs. AP600 PRESSURE ! (Model Pressure = 200 psia) 1200 __ 5% DecayHeat 100G-4% DecayHeat g --m-6 800 3% DecayHeat

                                                                                                                                                                                       -e-
                                                                                        --   -                                                                                         2% DecayHeat J\h a"hk      -     -

x % -

e 2 4

h 460 e60 800 160012c0140010001800 20'00 AP000 PRESSURE W Figure 4-11 Model Power Requirements vs. AP600 System Pressure for Steady-State Natural Circulation (Model Pressure = 200 psia) O MODELPOWERvs. AP600 PRESSURE (Model Pressure = 100 psia) 1200 _ ,_ 5% Decay Heat

                                                                                                                                                                                     +

1000 4% Decay Heat g800 "T

                                                                                         ^
                                                                                                                                                                                     -m-3% DecayHeat
                                                                                                                                                                                     +

2% Decay Heat goc a -\h y .

                                                                                                                                    .                        E    =   r     _

5  ; 0 0 200 460 6do ado 1600120014'00100018002000 AP600 PRESSURE (psia) Figure 4-12 Model Power Requirements vs. AP600 System Pressure for Steady-State Natura' Luw;ation (Model Pressure = 100 psia) 3681w-4b.non:Ib-060897 4-90

FACILITY SCALING REPORT O FROUDE NUMBER VS. LIQUID LEVEL 1/4 LENGm SCALE CLOSED LOOP 1 ANNULAROISJERSED INTERMITTENT-SLUG l 0.8--- w 0.6' ' *~~~~ OAn . .e . a . 02

                                                                               $~"                                 *--
                                         $        NVFIED                                 #

M '"8' O.1 0'2 03 0'.4 6.5 0'6

                                                                                             .      0'7     03       0'.9      1 DIMENSIONLESS UQUID LEVEL (tC)
                                                      --- AP600 -*- 4* DIA. e S* DIA. -e- 6* DIA.

Figure 413 Flow Regime Transition Boundaries for AP600 and OSU Model Hot Legs O DIMENSIONLESS DIAMETER (HOTLEG)

                                                                                                                 =
                                                                          -w-,...

m l100 D* = 32 10 0 2d0 40 6d0 8d0 1h 1200 PRESSURE (psia)

                                                                      -*- AP600 + MODEL Figure 4-14 Dimensionless Diameter (D*) for the AP600 and OSU Model Hot Legs 368Iw-4cmn:1b 060897                                         49]

• FACILITY SCALLNG RETORT I

l DIMENSIONLESS DIAMETER h (COLD LEG) g1000

.=. = -;

l 100 l - r-

                                                                                   . 1 v i _ __ __ -

l l 9 l 5 10 idOO 1200 0 2do 460 edo 800 l PRESSURE (psia) l

                                                                    -*- AP600 -+- MODEL Figure 4-15 Dimensionless Diameter (D*) for the AP600 and OSU Model Cold Legs O

DIMENSIONLESS DIAMETER (PZR SURGE LINE) ! gim g - l - = g  : 100: : - f _

r. -

l -

                                                           - --                                                                                                                           o. . ,

l 10 0 2do 4d0 6do 80 idOO 1200 PRESSURE (psia) l

                                                                     + AP600 + MODEL Figure 4-16 Dimensionless Diameter (D*) for the AP600 and OSU Model Pressurizer Surge Line 3681w-4c.non:Ib-060897                                                    4-92

1 1 FACILITY SCALING REPORT O Critical Heat Flux Scaling Pressure = 0.1 MPa(14.7 psia) 1K c e [} 1.01-P

                                                                                            =       =       =

15 m ~? =

                                                        .=

h h 5 5-t

                                                                                                           ^3 g

6 0.90,;

                                                                        =     =
                                                                                     =      =       i       i- -

l 8

c 0%

2 4 8 5 1'O i2 i4 (6 i8 2'O 22 Test Facility Mass Rux. Ihn/(A^2-s)

                                             -*- heuthig m .02 -*- tudWg =.04 -#- hsWhig m .06
                                             -+- heutWhig m .08 -*- hsWhig m .1 Figure 417 Critical IIeat Flux Similarity Criteria for the AP600 Test Facility O

Critical Heat Flux Scaling l Pressure = 0.69 MPa (100 psia) l, in E h 1.01 - e-

                                   'E 1

r r 2  :  :  : 2

l m .

On s

= ,

1 2 4 8 8 (0 i2 i4 1'6 i8 50 22 Test Facility Mass Hux, Lbm/(R^2-s)

      ,                                          -+- heichig = .02 -*- hautWhig m .04 -*- hsWNg m .06
                                                 -+- hsutWhig a .06 -*- haubhlg n .1

(3 Figure 418 Critical Heat Flux S!milarity Criteria for the AP600 Test Facility

  %.)

3681w-4c.non:lb-060897 4-93

FACILITY SCALING REPORT CriticalHeat Flux Scaling Pressure = 138 MPa(200 psia) 1K

                                                                 $ 1.01 E                    = _ = .                                                        =      =    =                       =

g 3__ 2, n 5 m o.gg - 3 m 2 4 6 8 (0 i2 i4 i6 i8 2'O 22 Test Facilhy Mass Flux, Lbm/(ha2 s)

                                                                            --- taubNg = .02 -*- totbHg = .04 -e- haubNg =.06
                                                                             -+- haubHg m .08 -*- tsubHg m .1 Figure 4-19 Critical IIeat Flux Similarity Criteria for the AP600 Test Facility O

Critical Heat Flux Scaling Pressure = 2.76 MPa(400 psia) ' 1m 6 1.01 I = * * * * *

                                                              }n
                                                                              ~

1 2 0.00 s x 2 4 6 8 (0 i2 i4 (6 i8 20 22 Test Facilky Mass Flux, Lbm/(fta2.s) )

                                                                          -*- helbMig = .02 -*- hsthNg m .04 -e- heutWg = .06
                                                                          -*- hotbMig = .08 -*- hsubNg = 1 Figure 4-20 Critical Heat Flux Similarity Criteria for the AP600 Test Facility i

3681w-4c.wrdb-060897 4 94

                                                                                          -                - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _               . _ _ _ _ _ _ . _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _    __ _ ._   _]

FACILITY SCALING REIORT O 11 EATER ROD AXI AL POWER PROFILE 0.25 e 0.2- --- - - - - - - - - - - - w it 2

                                                                                                                                                                                 ~ ~ -
                              $ 0.15-      ~ ~ ~ ~ ~ - - - - - -

! B B z g o.i. m , it: 2 0 .05- - - - - i i i 0 l 0.25 0.75 1.25 1.75 225 2.75 ELEVATION (Fees) l NORMALIZED POWER PROFILE 0.6

                                              - - - - - - - - -                                                             --~~---"--"~~~~"--""-""'~~~"-""

0.5 E . _ . - - . _ - . - . - .

                                  , o4     ..

2 l o a. 2 .. --- e og- ~~.--._---------------- -. O Z g,3 ... .....-.......-.....~..~...-~~-.. - -- 0 25 1.5 2 3 0 0.5 1  ; ELEVATION (Feet)  ; Figure 4 21 Axial Linear Power Profile (Normalized) for the Model Core 368iw.4cm!b.060897 - 4 95

l FACILITY SCALING REPORT O

                                                                       ///

G G G G A.... A .. . . . . e

                                               /        o G  9    9    9    .

O G . G G e e e e e o // e e f

                                                                     .    .    .    . y G    G LEGEND GS.805 kV RODS g    g    /                     911.02 kV RODS STHERH0 COUPLE RODS Figure 4-22 Radial Power Distribution in Power Core t

I 3681w-4c.non:1b-060897 4-96 1

FACILITY SCALING REPORT 5.0 OPEN ShSTEM DEPRESSURIZATION SCALING ANALYSIS One of the significant difficulties encountered when operating a reduced-pressure facility is establishing the proper initial conditions for testing. For APEX, much of the phenomena of interest (such as long-tenn recirculation cooling), occurs at pressures well within the capability of the facility. However, to reach the onset of long-term cooling, the system first must evolve through a variety of states (mass and energy inventories); the trajectory of the states is controlled by the dominant transport processes such as break flow rate, secondary-side heat transfer, and decay heat. The initial state of the full-scale system, just prior to a transient, cannot be duplicated in a reduced-pressure system such as APEX. Therefore, a method is needed to define the initial and boundary conditions for APEX, just prior to the transient, such that the prototypic trajectory of system states can be duplicated on a scaled i basis, leading to the correctly scaled initial conditions for long-term cooling. This section addresses this need by presenting a set of similarity criteria applicable to open system depressurization transients. The specific objectives of this section are as follows: To define the initial and boundary conditions for the range of tests being considered. To define the break sizes. To provide a method of comparing full-scale pressure to model pressure by defining a reference pressure. To develop a method to achieve the properly scaled initial conditions for long-term cooling.

                   = To maintain the timing of events on a scaled basis.

To identify the transition pressure at which fluid property similitude begins. To meet these specific objectives, four methods are used. First the governing set of equations for the depressurization of a two-phase fluid system is developed. This includes the equation of two-phase elastic compliance, a two-phase mixture density equation of state, and a depressurization rate equation. Second, a top-down subsystem level analysis is performed using the integral balance equations to obtain an analytical solution for the time-dependent pressure for a break flo-w rate dominated system. The theory that forms the basis for this analysis is presented in OSU-NE-9407,W This includes the development of a critical flow relation, predictions of the time required to uncover a single break, and a comparison to experimental data. Third, a top-down system level scaling analysis is performed for a system containing multiple vent and injection paths. This includes the development of scaling criteria for break flow rate dominated and volumetric expansion dominated systems. Last, a bottom-up { analysis is performed to describe the scaling of the source tenns (such as core decay heat) and the local transport processes such as the critical flow through the breaks. Figure 5-1 presents a list of system and local processes addressed by the scaling analysis. 3681w-5.non:lb-o60897 5-1

FACILITY SCALLNG REPORT l 5.1 Description of the Depressurization Process Depressurization of the AP600 can occur as the result of a primary system break or operation of the automatic depressurization system (ADS). The ADS consists of independent tiers of valves that open in sequence to reduce primary system pressure in a controlled momer. Reducing system pressurc pennits core makeup tank (CMT), accumulator (ACC), and in-containment refueling water storage tank (IRWST) injection. The first three sages of the ADS vent steam or a two-phase fluid mixture from the top of the pressurizer directly into the IRWST through a set of spargers. He fourth stage of ADS consists c' two sets of valves located on each hot leg. These valves vent directly into the reactor cavity inside containment. For the AP600 design, breaks in the primary system result in a very rapid initial depressurization to the saturation pressure that corresponds to the secondary-side temperature. While at saturation conditions, the primary-side fluid expands rapidly as vapor is generated. The break flow is choked, and the quality at the break changes slowly. When the volumetric expansion of the two-phase fluid equals the volumetric flow out of the break, then a quasi-steady condition, with respect to break flow and system pressure, is established. His plateau pressure, depicted in Figure 5-2, serves as the reference pressure for scaling. The duration of this constant pressure period is dependent on the break size and the vapor generation rate. For practical purposes, the time at which the steam generator reaches its maximum pressure can be used as the transient start time. It will be used to define the initial conditions for the transient for purposes of comparison 'vith counterpart tests and full-scale AP600 calculations. Actuation of the ADS results in lowering the system pressum to near atmospheric conditions where fluid property similitude exists between the AP600 and the APEX, therefore pressure scaling is not required at these conditions where fluid property similitude exists. Two types of depressurization cases are examined. Fw breaks equal to or exceeding two inches in diameter in the AP600, the depressurization rate is dominated by the break volumetric flow rate. For breaks equal to or less than one-inch in diameter in the AP600, the depressurization rate is dominated by the system fluid volumetric expansion. 5.2 Governing Equations for the Two-Phase Fluid System Depressurization This section presents the governing equations for the depressurization of a two-phase fluid system. The unique features of this section lie in the development of a two-phase fluid clastic compliance equation, a two-phase fluid volumetric dilation equation (based on a volume-averaged slip ratio), and a two-phase depressurization rate equation. f l l O 368iw s-ib-oeo897 5-2 _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - - - - - - - - _ _ _ _ _ _ _ _ _ _ _ - - - . - - - l

FACILITY SCALING REPORT 5.2.1 Stress Strain Relation for a Two-Phase Fluid in Compression

As part of his studies of the mechanics of materials, Hooke first observed

"As the stretch, so the l force."W Hooke's observation that the deformation of a body, defined by the strain, is proportional to the load applied to the body, defined by the stress, became the basis for the development of a general theory of elasticity. For an isotropic body, the general theory provides the following fundamental relationship between stress and strain:W e = E" : 5 (5-1) where:

i E = strain dyadic (second order tensor) that describes the volumetric dilation of the body I 5 = stress dyadic that describes the internal stresses gc= i 1 elastic compliance tetradic, also known as Hooke's tensor

                                                                                                                                                                              )

l

    '(A)               For a homogeneous, isotropic fluid under compression at equilibrium conditions, Equatior. 5-1 reduces to the following form:

F, = E.

                                                                                                                    -                                                   (5-2)

P where the stress term is represented by the pressure (P) acting normal to any element of the surface. For a fluid, the elastic compliance term (E c) is the inverse of the bulk modulus of compression. That is:W E* = I. dp (5-3) p dP Figure 5-3 depicts a pressure vessel that has a total volume (V T ) and that is initially filled with saturated liquid. Subsequent to opening the break, the liquid flashes to create a bubbly liquid / vapor mixture with a two-phase mixture density. A vapor bubble occupies the top of the tank, while a two-

  -V(9 phase mixture leaves through the break.

I

                     ~ 36siw.5.non:ib-060897                                                                         53

FACILITY SCALING REPORT The axiom of continuity for the system is given as follows: a b + "g + "I = 1 (5-4) where: ab = given by VA'T as = given by Vg/V7 c, = given by V,/VT The volume averaged two-phase mixture density is as follows: (PT P) " ]PTPdVTp (5-5) where VTP is the volume of the two-phase fluid mixture. For uniform fluid properties, Equation 5-5 can be written as follows: UTP (5-6) (PTP)*PTP whem: ot7p=ag + a, (5-7) Furthermore, for unifonn fluid properties: (PTP)

• Pg Ug +Pt"t (5-8)

Substituting the two-phase mixture density into Equation 5-3 yields the equation of elastic compliance for the two-phase fluid mixture: l I d (5-9) E, = (pTP) dP (prp) 1 36siw.5.non iunso897 5-4

FACILITY SCALING REPORT O- Substituting Equation 5-2 into Equation 5-9 yields the stress-strain relation for the two-phase fluid mixture. E

                                                                                                                                                                                          =     1        d(pTP)

(5-10) P {pTP) dP Substituting Equation 5-6 for uniform conditions into Equation 5-10 yields the following: 1 E 1 d

                                                                                                                                                                                                       -                              U                                              (5-11)

P (pTPU TP) dP (PTP TP) 5.2.2 State Equation for a Two-Phase Fluid Undergoing Depressurization Equation 5-11 could be integrated if an expression for the volumetric dilation (E) were known as a function of pressure. The volumetric dilation is defined as the relative change in volume of the two-r' phase fluid mixture. That is: l

                                                                                                                                                                                                                                                                                                           )

AVTP E= (5-12) V TP where AVTP is the volumetric expansion or contraction of the two-phase mixture. j Postulating that the volumetric dilation is a constant which is dependent on the initial conditions of the system: E=Eo=f(P,aTPo) o (5-13) 1 Substituting Equation 5-13 into Equation 5-11 and separating variables yields the following: , 1 E d . dP (5-14) r (PTPU TP) (p7paTP)*P i 3681w-5.non:Ib 060897 5-5 Ci_____________..____________________.____________-._----- _ _ . _ _ _ - - _ _ . _ _ . - - - - _ - _ - - - - -- - - _ - - - - - _ - - - ---- _ _ . _ - - _ - - _ _ . - - .- - - - - - -

FACILITY SCALING REPORT Integrating both sides and solving for the two-phase fluid mixture density yields the following:

                                                                                                     'E, (5-15)

PTP U TP "(PTPUTP)o s os Equation 5-15 is a state equation for the two-phase fluid mixture that describes the variation of the two-phase fluid mixture density with system pressure. In OSU-NE-9407, the following relation for the volumetric dilation is obtained for the depressurization of a simple system:W l S7 2 (5-16) e = (Y-1) where: y = ratio of the specific heats of steam (i.e., y = 1.33) So= volume-averaged velocity slip ratio The volume-averaged slip ratio (S) is given by the following: O.4 So = =0.3 $ (5-17) ( Esso I Substituting Equation 5-16 into Equation 5-15 yields the final form of the two-phase mixture density state equation:

                                                             ,                       , s,y 2 U                                          (y -1)                                                                               (5-18) pTPU TP *(PTP TP)o
                                                             ,             o, O

3681w 5.nortIb-060897 5-6 i ___ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ . _ _ _ _ _

FACILITY SCALING REPORT 5.2.3 Depressurization Rate Equation a ne continuity equation for the two-phase fluid mixture is given by the following: 1 i U (pTPTP)" - I g+ Pe e c (5-19) where: s,i = the liquid mass injection flow rate into the system l Ts = vapor generation rate due to phase change  ! p, = fluid density at the break plane e, = speed of sound for the fluid at the break plane a, = break flow area Substituting Equation 5-14 into Equation 5-19 yields the following: p EopTPUTP dP rh , _ p,c,a e (5-20) P I,V T 8 VT l Dividing both sides by EnpTP UTP yields the depressurization rate equation: l dP , 1 N in _p _Pec ,a e _ (5-21) g i P dt copTPUTP VT y To evaluate Equation 5-21, relations defm' ing the coupling between the primary system pressure and

                                                 . the source and sink terms are needed. He following sections examine this coupling.

S.3 Top-Down Subsystem Level Analysis for the Depressurization of a Break Flow Rate Dominated System The analytical model presented herein is developed in detail in OSU report, OSU-NE-9407.W As indicated in that report, this model was developed as part of the ongoing effort to understand the coupling between critical flow at a pressure boundary and the time-dependent system pressure that []

 %.)

drives the flow.

                                                 . 36siw 5.non s o60897                                              5-7

FACILITY SCALING REPORT In the present discussion, an analytical model is developed to predict the pressure trends that would be observed following the puncture of a pressurized system initially at saturation conditions. The unique features of this analysis lie in the development of a two-phase critical mass flow rate relation and an expression for time-dependent pressure in a break flow rate dominated system. 5.3.1 Critical Flow Rate During System Depressurization This section develops an equation, for the system depicted in Figure 5-2, that relates system pressure to break mass flow rate during depressurization. The basis for the model is that the expansion rate of the vapor bubble located at the vessel head and upper regions of the reactor system are affected by the expansion of the two-phase fluid mixture and the rate at which the fluid mixture level recedes (which is controlled by the break flow rate). For a single vapor bubble under compression, neglecting viscous stress, the bubble growth rate can be ' described by Rayliegh's classical equation as presented in Lamb's text, Hydrodynamics.W RN+ b 4) (5-22) b b hbf "Sc (PTP) where: O Rb= bubble radius Pb = internal pressure of the bubble P., = pressure in the two-phase fluid far from the bubble The surface tension term 20/(R b<PTP>) has been neglected because the bubble radius being considered is typically quite large. The doub'e dot superscript represents the second derivative with respect to time, and the single dot superscript represents the first derivative with respect to time. Using Bemoulli's equation, the term on the right side of Equation 5-22 can be written as follows: ge (Pb - P ) = - CvMhb)2 (5-23) (PTP) where Cyg si the virtual mass coefficient expressed in a form similar to that developed by Zuber.W l O l 3681w.5.non:Ib.060897 $-8

FACILITY SCALING REPORT  ! l. That is: I b (5-24) Cyg = 2-1 *b i l l Dis equation differs from Zuber's in that the coefficient m is defined as follows:

                                                                                                                                                   \

I m= '57-6' (5-25) j s Y , For the case of y equal to 1.33, m equals 0.5 and the virtual mass coefficient is written as follows: 1 1 .5a b C (5-26) VM = 2 1 -ab The negative sign in Equation 5-23 arises due to the inotion of the bubble interface. Substituting Equations 5-23 and 5-24 into Equation 5-22 yields the following: RN+ b b (kb f " - bf

                                                                                                                                             ~

I a The volume of the bubble is the following: Vb= n Rj (5-28) l

i. De first and second derivatives with respect to time are as follows:

(- l. t 9 b= 4x Rb Nb

  .m]
                   '36siw-5. mi m 97                                           5-9

FACILITY SCALING REPORT b = 4n Rb RbNb + 2 hb Substituting Equations 5-28 through 5-30 into Equation 5-27 and rearranging yields the following: ( +1 V9=f 1 b (5-31) b b g_ 9b )2 For the break flow rate dominated case,(rh inNT) and I's are small compared to the break flow rate and the continuity Equation 5-19 reduces to the following: dVTP dPTP (5-32)

                                                          ~

PTP dt

                                                                ***             dt Recognizing that the change in the fluid density with respect to time is small compared to the other terms in Equation 5-32 allows the following simplifications:

dV7p p,c,a' (5-33) dt pTP From the time derivative of the axiom of continuity, it can be shown that: dV TP . Q b (5-34) dt Substituting Equation 5 34 into Equation 5-33 yields the following: b= (5-35) PTP 3681w $.non:lb.060897 $-10

[: i l-FACILITY SCALING REPORT j 1 where the mass flux (G,) through the break is given by the following: I G, = p,c, (5-36)

q. ,

I

                                           'Ihe second derivative is given by the following:

a, dG* V. (5-37) b - l pTP dt i l-l l where the time derivative of the density ratio is small compared to changes in the mass flow rate. Substituting Equation 5-35 through 5-37 into Equation 5-31 yields the following: ) abV1a, dG, ,_1 '(m +1)ab G,a c (5-38) pTP dt 6 {l-a), PTP , b j \ rg - i

   'Y                                      For the break flow rate dominated case, (rh                     inNT) and I'g are small compared to the break mass flow rate and the depressurization rate Equation 5-21 becomes the following:

l dP , -_ . G, a' (5-39)

                                                                                                 ..P    dt         cpTP(I'U)V o           b T Dividing Equation 5-38 by 5-39 and rearranging yields the following:

P o G,_dPdG, 6, (m +1)c (5-40) L Separating variables and integrating yields the desired expression for break mass flux as a function of system pressure:

                                                                                                                          , (m +1)c, th                                                                                               G, = G,,o

[ 6 (5-41) 36stw-5.non:Ib-060897 5-11

FAC11.ITY SCALING REPORT Substituting the definitions for Eoand m, given by Equations 5-16 and 5-25 respectively, into Equation 5-41 yields the following: e ,soy P (5-42) Ge = G e.o p _ r o s

           'Ihe simplicity of this equation is quite remarkable and can be readily tested using existing data or correlations for critical mass flux.

Figure 5-4 presents the results of several isentropic depressurizations using the homogeneous equilibrium model for critical flow at the break. Equation 5-42 is found to be in exce!!ent agreement with the results for the ideal case (HEM) with So equal to 1. That is: r sy P (5-43) Ge = G e,0 p g O j The critical mass flux ratio shown in Figure 5-4 is defined as (G,/G,,,). 5.3.2 Analytical Model for Depressurization from Saturated Conditions Having obtained the goveming equation for depressurization (Equation 5-21) and an equation that couples the break flow rate to the system pressure (Equation 5-43), an analytical expression for system pressure as a function of time can be derived for the special case when (rhi ,/VT) and I'g are small compared to the break flow rate. For this set of conditions, Equation 5-21 becomes the following: I dP , _ G,a* _ _ (5-44) P dt t opTPOC TPVT The pressure ratio can be defined as follows: P * = P/P o (5-45) 3681w-$mn:Ib 060897 5-12

FACILITY SCALING REPORT f Substituting Equations 5-41 and 5-45 into Equation 5-44 yields the following: (m +1)c, I dP + =- a,G,,,(Pj 6 (5-46) P+ dt e,pTPCI TPVT Further aore, substituting the state equation for the two-phase fluid rnixture density (Equation 5-15) yields the following: 1 dP + *- a,G,,, 1 d; (5-47) P+ ,E,pTPouTPo V, T ** (P j 6 (5-m) Equation 5-47 can be further simplified by defining a density weighted sound speed as follows: (O_/ c,, - G* (5-48) I

                                                                                                       .PTP o This two-phase, density-weighted, sound speed can be readily evaluated using existing models f6r the critical mass flux. Substituting Equation 5-48 into Equation 5-47 and rearranging yields the following:

c p5 m)-1 dP

• Csoa e (5-49)

(pg ,, dt eV TPo-I

                                            . where:
                                       ,                                                                                     +

VTPo " "WoVT (5-50) l j i o 36siw-5.non-ib.060897 5-13 p t ..

                                                                                                                                                         ]

FACILITY SCALING REPORT Defining a specific frequency, the inverse of the residence time, as follows:

                                                                     #50 **

co, . (5-51) TPo j and substituting Equation 5-51 into Equation 5-49 and separating variables yields the following: i E r l" [p gy(5 m) - IdP * = - dt (5-52) T ' Integrating both sides of Equation 5-52 yields the desired analytical solution for the pressure ratio as a function of time and initial conditions. That is: 6 p . ,) , 3 _ ' S -m (' oot **(5 *)

                                                                ,  6, Substituting the definition for m given by Equation 5-25 yields the final form of the analytical solution:

Y/E, to (5-54) P(t) = P o 1o t Y I l

                                                                                                                            \

Once again, the result is remarkably simple. I l Having developed an analytical expression for the system pressure as a function of time (Equation 5-54), an analytical expression for the critical flow at the break can now be developed. Substituting Equations 5-25 and 5-54 into Equation 5-41 yields the following-1 y -1 Ge= G e,o 1 bt (5-55) Ol 368Iw $.non:1b-060897 5-14 j 1

1 FACILITY SCALING REPORT l

                                    'Ihe enthalpy flow rate at the break is given by the following:

Y -1 4 G, h, a, = G,,oh , a, 1 bt Y (5-56) where h, is the enthalpy of the fluid at the break plane. I 5 3.3 Analytical Model for Single-Phase Vapor Depressurization

                                 . Equation 5-54 can be written in a general manner as follows:

t 2 (5 57) P(t) =o P [1 -C a>o ]C

                                 - For a pressurized system, filled with a perfect gas which is expanding isentropically and undergoing isentropic acceleration at the break, Wulff(6) shows that:

A

  'V-                                                                                                   y+1 C = -(y - 1)          2   2(y - 1)                                      (5-58) 3 2       ,y + 1, and

_,7

                                                                                               ~   -

C (5 59) 2 = (Y - 1)

                                 ' 5.3.4 Comparison to Marviken Data Figure 5-5 presents a comparison of Equation 5-57 using the appropriate coefficients for the two-phase and single-phase portions of a blowdown experiment performed at the Marviken test facility.U) A RELAPS calculation has also been performed for the same test. For this test, the fluid was initially at saturated conditions. Equation 5 57 shows excellent agreement with the data. During the two-phase portion, before the tank empties of two-phase fluid (that is, break uncovers), the coefficients given in
  \ j_                            Equation 5-54 are implemented. Subsequent to the break uncovering, the depressurization rate increases rapidly as single-phase vapor escapes from the tank. The coefficients given by 36siw.5.non:ibe60s97                                  5-15

FACILITY SCALING REPORT Equations 5-58 and 5-59 are used to predict the later portion of the test. This comparison with Marviken data provides support to the analytical approach. 5.4 Top-Down System Level Depressurization Scaling Analysis The previous section developed and evaluated (for a simple case) a depressurization rate equation (Equation 5-21) applicable to a single break. This equation will serve as the basis for the system level scaling analysis. For a primary system having multiple injection and break points, Equation 5-21 can be written as follows: I$= I (5-60) Eo pTP(ITPVT frhin - F VT - E$e] g _P dt For the purpose of system scaling, this equation can be made dimensionless by substituting the two-phase mixture state equation, dividing each term by their initial condition, and dividing both sides of the equation by the initial break or vent flow rate (Lh,,o). This yields the following: T,y,(P f*-1 dP * , g g$m* NI*g -E*e* (5-61) where the time constant for the overall system transient is given by the reciprocal of the system specific frequency. That is: 2 = 1

                                                                                                                                                                                                                                       -  PU V"                                                                                         (5-62)

U sys b e,o This time constant physically represents the time required for the two-phase fluid to leave the control volume. For the case of a single vent path, Equation 5-62 reduces to the following: t, = _ I_= " (5-63) Oo m e,o O l 3681w-5 non:1b-060897 5-16 l

                                                                                                                                                                                                                 -                              - _ - _ _ _ _ _ _ _ - _ _ _ _ _ _ - _ - _ _ _ _ _ _ _ _ - _ _ _ - _ _ _ _ _ _ - _ _ - _ _ _ _ - _ _ _ _ _ _ _ ~

FACILITY SCALLNG REPORT O Substituting Equation 5-48 yields the following: O 1 I TP to == (5-64) Oo es ,a,

                                                 ~ which is the same result given by Equation 5-51. The characteristic time ratios are given by the following:

Lh "- Um= m,t,y, = (5-65)

                                                                                                                                                       =          T Ur " Wr Tsys                                              (5-66)

C.O i where: Um = system mass flow rate ratio. For a constant injection flow rate, H, represents the total liquid mass injected into the primary system during the residence time (t,y,). For large break rates and small injection flow rates, Umwould be small. Up= vapor generation rate ratio. For a constant vapor generation rate, D represents r the total vapor mass generated in the primary system during the residence time (t,y,). The following characteristic time ratio can be obtained by performing an energy balance on the system: Mrhh)i-Uh *% Tsys = (5-67) Mrhh),,, where Uh is the enthalpy flow rate ratio. U represents h the ratio of the total enthalpy change due to

                                               - fluid injection to that lost by break flow during the residence time (t,y,). Table 5-1 summarizes the essential equations needed to scale system depressurization events. The time constant and

! [ characteristic time ratios (H groups) can be used to develop ideal scaling ratios for specific

  '(

l-368Iw.5.non:Ib-060897 5 17 E _ __ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - . _ _ _ _ _ - _ _ _ - - . - . __

FACILITY SCALING REPORT components. In general, by writing each H group in terms of a model to prototype ratio and setting the ratio equal to unity, a similarity criterion is developed. That is: Hg=1 (5-68) Evaluating the similarity criterion in terms of physical properties results in a scaling ratio that relates a geometrical or flow feature in the AP600 to the corresponding feature in APEX. The following sections develop the scaling ratios for system depressurization events. Equation 5-61 has been normalized using the break flow rate. Therefore, when the numerical value of either U mor U exceeds p one, this indicates that the process represented by that characteristic time ratio would have a greater impact on depressurization than the break flow rate. Conversely, if the numerical values of the characteristic time ratios are less than one, this indicates that the break flow rate has a greater influence on the depressurization rate than the other processes. U mwill typically be quite small for most of the depressurization transient. Furthermore, the injection flow rate is a control parameter that can be easily adjusted. Because of this, Equation 5-61 needs to be evaluated only for three sets of conditions. First, the case where break flow dominates the depressurization rate will be examined. Next, the case where dP+/dt equals zero will be examined. Last, the case where the vapor generation rate dominates the depressurization rate will be examined. 5.4.1 Scaling Depressurization Processes Dominated by Break or Vent Path Flow Rate (Breaks Greater Than or Equal to 2 Inches) For breaks greater than or equal to two-inches in diameter in the AP600, U m and Up will be less than one. Therefore, the break flow rate dominates the depressurization rate. The two-phase fluid residence time constant for the system is dictated by the break flow rate. By developing a time scaling ratio and setting that ratio to 0.5, as in previous analysec, a break flow rate scaling ratio can be 4 developed. Substituting the break flow rate scaling ratio into each of the characteristic time ratios (fI m and U p) permits one to obtain the scaling ratios pertinent to those processes. This procedure is implemented in the next sections. O 3681w s.non:lb-060897 $-18

FACILITY SCALING REPORT I Scaling the Break Flow Rate (Breaks Greater Than 2 Inches) i i The two-phase fluid residence time constants, Equation 5-62, can be written in terms of a time scaling ratio: i PTPoVTPo tsys R . (5-69) gg e.o

                                                                                                                                .R This equation can be written in terms of mass flux as follows:

xsys.R . n'o V TPo (5-70) EC p G,,,a*

                                                                                                                                  . R                                l I

I where Cois the discharge coefficient of the break or ADS valve. l p, Every process in APEX has been scaled such that: I h  ! t,y,,g = 0.5 (5-71) Furthermore, the system volumes have been scaled such that: Vg = 1/192 (5-72) Substituting Equations 5-71 and 5-72 into Equation 5-70 yields the following: EC o 0,,,a, )

                                                                                                                             , . ,                         (5-73)

PTPo .R L This requirement will be evaluated for each of the AP600 breaks (two-inches or greater) and for the

                                                ' ADS valves in the bottom-up scaling analysis.

m_ 368tw 5ma:twi60897 5-19

FACILilT SCALING REPORT Scaling the Injection Flow Rates (Um,g) 1 l Equation 5-65 can be written as a Il group ratio as follows: Gi,ain IgR . (5-74) E CD G,,oa ' . R where Gi , is the injection mass flux and ai, is the injection flow area. Because of geometric similarity, the discharge coefficients for the APEX and AP600 injection flow areas will be identical. Setting Equation 5-74 equal to one yields the following: fC pG,,,a,], = pi,a m)g MS Substituting Equation 5-73 into Equation 5-75 yields the scaling ratio for the injection flow rates: E G in a,, j s_ (5-76) PTPo .R Equation 5-76 will be coupled to the similarity criteria of Section 7.0 to obtain cross-sectional flow areas for the core makeup tank (CMT), accumulators (ACC), and in-containment refueling water storage tank (IRWST) injection lines. Scaling the System Vapor Generation Rate (Ilr,R) Equation 5-66 can be written as a scaling ratio as follows: goV T q" = (5-77) ECp G,,oa* R O 3681w 5.non:Ib-060897 5 20

FACILITY SCALING REPORT where I',, represents the vapor generation due to phase change. For a saturated system, it is expressed as follows:

                                                                                                 =    sys I'8                                       (5-78) hVg7 l

t j- where: I hg = latent heat of vaporization q,y, = net thermal power added or removed from this system. I The net thermal power added or removed (q,y,) is expressed as follows: 9sys " 9 core + 9 metal + 9SG + 9PRHR (5-79) where: 'O .q co,, = heat release rate in the core q,,g = metal stored energy release rate 9so = steam generator heat removal rate

     . qpggg = energy removal rate of the PRHR Substituting Equation 5-78 into Equation 5-77 yields the following:

SY8 11,R r = (5-80) hgEC p G,,,a, . g l Setting Il ,g p equal to one and substituting Equation 5-73 yields the following: 95YS I

                                                                                                        .                              (5-81) hgpTPo R

i L r\ ~ V Equation 5-81 is evaluated using the fluid property scaling information obtained through the bottom-up  ! scaling analysis. ' 36st w-5.non:tb.071897 5-21 l

FACILITY SCALING REPORT Scaling the System Enthalpy Flow Rate (FIh,R) Equation 5-67 is written as a scaling ratio as follows: { (rhh)I"' T[1,g = (5-82) ((ih),, . g where hi , is the enthalpy of the injected fluid which is typically subcooled. Setting FI h,R equal to one yields the scaling ratio: Mthh)in.o ,i (5-83) Md6),,, . g Fluid property similitude is maintained for every injection process, with the exception of CMT injection following prolonged CMT recirculation. CMT processes are discussed in detail in Section 6.0. By assuming fluid property si:r.ilitude for the injection processes, the following is true: (hin.o)R Substituting Equation 5-84 into Equation 5-83 yields the following expression: 8 0. R in.o)R Substituting Equation 5-76 into Equation 5-85 yields a scaling criterion for the break and vent path energy flow rate. That is: Mrhh),,, 1

                                                                                     , _ . _                                                                                                       (5-86)

PTP,o

                                                                               .R O

3681w-5.non.It>.060897 5-22

4 1 FACILITY SCALING REPORT For a single initiating break in a reduced-pressure scale test facility, the criteria given by Equations 5-76 and 5-86 cannot be satisfied simultaneously because this would require fluid property similarity at th'e break. Specifically, (h,,o)R However, as noted previously, for breaks exceeding two-inches in diameter, the volumetric flow out of the break dominates the depressurization rate. Therefore, Equation 5-76 is used to scale the break sizes for AP600 breaks greater than or equal to two-inches. His introduces some distortion in the break energy flow rate, but still preserves the depressurization rate behavior. The previous section demonstrates that when the break mass flow rate dominates, the volumetric dilation term cancels (upon integration) on both sides of Equation 5-61. Therefore, the ratio of E n need not be incorporated into the characteristic time ratios for break dominated systems. Table 5-2 summarizes the scaling ratios for break flow rate dominated system depressurization. For depressurization events in the AP600 that initiate at 400 psia or less, fluid property similitude exists between the AP600 and APEX. For these conditions, the similarity criteria for system depressurization would reduce to the set of equations presented in the lower half of Table 5-2. For the ( case of fluid property similitude, it will be shown in the bottom-up scaling analysis that the critical mass flux ratio (G,)g) is equal to one. 5.4.2 Scaling Initial Conditions (dP/dt Equal to 0) As shown in Figure 5-2, subsequent to the rapid depressurization from initially subcooled conditions, the system reaches a plateau pressure. His quasi steady pressure period exists as long as the volumetric expansion rate of the system fluid equals the volumetric flow rate out of the break. For

                  . large breaks, this period is relatively short. For practical purposes, the reference pressure is the
                 - pressure at which the initial subcooled blowdown ends. It is defined as the maximum pressure on the

secondary side of the steam generators subsequent to initiating the break. This plateau pressure has

                 , been selected.as the reference pressure (P                               o ) for scaling and is mathematically defined as follows:
                                                                                                  $dt     =0                                        (5-88) po he use of P as         o a reference pressure for scaling is consistent with the work of Moskal.(8) he reference pressure is used to scale the fluid properties and to establish system setpoints in the bottom-up scaling analysis.

368iw-5.non:Ib.060897 5 23

FACILITY SCALING REPORT 5.4.3 Scaling Depressurization Processes Dominated by System Energy / Volumetric Expansion (Breaks Less Than or Equal to 1 Inch) For breaks less than or equal to one-inch in diameter in the AP600, Up will be much greater than one prior to ADS operation. U m remains much less than Dr during this same period. Therefore, the system vapor generation rate, which dictates the fluid volumetric expansion, dominates the depressurization rate. Under these conditions, Equation 5-61 reduces to the following: d T,y,(P f , -- r (5-89) Because the vapor generation rate which drives the volumetric expansion of the system fluid dominates the depressurization behavior, it is desirable to scale the break flow using the following scaling ratio:

                                                                                       --.-      =1                                         (5-90) 0R Substituting Equation 5-80 into Equation 5-90 yields the break mass flow rate scaling ratio required to O

simulate transients dominated by the volumetric expansion of the fluid. That is: 95Y5 (5-91)

                                                                                                          -1 Ehog    ECoG,,na '      . R Equation 5-91 is used to scale AP600 breaks that are one inch in diameter or less. Table 5-3 summarizes the scaling ratios for system depressurization events dominated by fluid volumetric expansion.

5.5 Scaling Synergistic Phenomena A major contribution of a properly-scaled integral system test facility is its ability to simulate the synergistic behavior that arises from the interaction of its subsystems. This behavior is often difficult to predict and provides a good test for systems analysis computer codes. A powerful aspect of the Hierarchical, Two-Tiered Scaling (H2TS) methodology is that it provides a quantitative method of assessing the impact of subsystem processes on integral system behavior. 1 3681w-5.non. lb-060897 5-24 L___________.

l FACILITY SCALING REPORT l t []

           \s Section 5-3 presents a top-down analysis at the hierarchical level of subsystem (that is, a single break or an individual ADS valve actuation). A unique residence time constant, given by Equation 5-64, can be generated for each subsystem. Section 5-4 presents a top-down analysis at the system level. The residence time constant for the system is given by Equation 5-62 and incorporates the aggregate effect of all of the subsystems (that is, multiple breaks and ADS valve actuations). As a result, the system scaling ratios presented in Table 5-2 include the summation of multiple mass and energy transport

{ processes. If one is solely interested in overall system processes, such as the total time required to { depressurize the system or the total energy deposited during the blow-down period, maximum flexibility in experiment design can be achieved by requiring that the summaticn of the mass and energy transport processes satisfy the appropriate system scaling ratio without regard to the exact scaling of individual subsystem transport processes. This hcwever results in a loss of information at the subsystem level. The most restrictive method of satisfying the system scaling ratios, with respect to experiment design, would be to scale each subsystem transport process individually. To demonstrate this approach, let us define a system scaling ratio comprised of multiple subsystem transport processes (A ) as follows: N EA im

                                                                      .I
           ,m                                                                    = po                                     (5-95)                           l 3                                                                                  i A

i iP i where p,is a constant. This system scaling ratio can be satisfied by requiring that: l A i,g = o A.R"ho 2 (5-96) AN.R " Oo This mathematical identity can be readily proven by substitution. The system scaling ratio given (see Equation 5-95) can also be satisfied by individually scaling only those subsystem processes that would i significantly impact system behavior. Other subsystem processes could be scaled as a set or neglected if their time scales are such that they do not impact system behavior.

          ,9 f

Q Because a wide range of tests will be performed in APEX, it is difficult to envision all of the subsystem interactions that might arise. Therefore, to achieve the best scaled representation of the l 3681w 5.non:ltM6o897 5-25

FACILITY SCALING REPORT AP600 and to establish a consistent design approach, the most restrictive method of satisfying the exact scaling of individual subsystem transport processes will be applied. This, however, results in a loss of information at the subsystem level system scaling ratios, as defined by the mathematical identity stated above. This approach preserves the residence time constants for the individual subsystems and the integral system. 5.6 Bottom Up Scaling Depressurization Scaling Analysis The objective of this bottom-up analysis is to obtain the closure relations needed to evaluate the scaling ratios developed through the top-down analysis as presented in Tables 5-2 and 5-3. These closure relations are dependent on fluid properties. Therefore, the first step in the bottom-up analysis must be the development of the method to relate fluid properties in the AP600 to those at reduced pressure in APEX. Scaling relations shall also be developed for the following phenomena, transport processes, and components:

• Two-phase natural circulation during depressurization
• Critical flow through the breaks and ADS valves
• Core energy transfer
• Component stored energy transfer
• Steam generator
• Pressurizer
• PRHR heat exchanger The following sections present an analysis of each of these topics.

5.6.1 Pressure Scaling The purpose of this section is to determine the scaling relationship between depressurization transients in the AP600 (which normally operates at 2250 psia) and APEX (which has a maximum operating pressure of 400 psia). This is accomplished using a method consistent with the method developed by Kocamustafagullari and Ishii(9) and further expanded upon by Moskal.(8) However, rather than using the graphical approach used by Kocamustafagullari, an analytical solution for the pressure scaling equation is developed using the Clausius-Clapeyron equation.00) The basic scaling parameter implemented in their analysis is as follows: Ap Y= (5-97) PssPtshg 9 3681w.5.non:1b-060897 5-26

  ~                                                                                                                                                                                                                                                          _ _ _____________ _ _.________

7. I l FACILITY SCALING REPORT l 1 i i where: Q.

   /

l Ap = (pg-p) p ph = saturated liquid density pp = saturated vapor density hk = latent heat of vapor.zation Note that this scaling group appears as the fundamental fluid property scaling ratio in all of the coefficients of the cubic velocity equation presented in Equations 4-106a through 4-106c. Therefore, maintaining Y go similitude assures that the two-phase natural circulation behavior that occurs during a slowly evolving depressurization event is reasonably modeled if power is scaled accordingly. l Equation 5-97 can be expressed in terms of the specific volumes of the saturated liquid (yh) and saturated vapor (v p ) as follows: Y=h h (5-98) g () v where vg=vp - v h. Substituting the Clausius-Clapeyron equationUO) for the saturated conditions yields the following:

                                                      'I      dT'         h                                 (5-99)
                                                      ,T      dP s SAT    h g Combining with Equation 5-98 yields some additionalinsight. That is:

Y= I

                                                               ,T dP, (5-100)

SAT Requiring the Y value in the model to be proportional to that in the prototype yields the following: Y m = (Yo ), Y p (5-101) ex 36siw-5-1b-060897 5-27

FACILITY SCALING REPORT where (Yo)g si evaluated for the model and the prototype using Equation 5 98 at their respective transient reference pressures (Po). l Substituting Equation 5100 into Equation 5101 and rearranging yields the following:

                                                ,       ,,           s.  ,       ,

Tm dT P dP m (5-102) g)

• _Tp , , dTm, . s dPp ,

                                                                                    ,1 s

For saturated conditions in both the model and prototype, the first term on the left side can be approximated by the following: r , , , , , PP Tm dT P (5-103)

                                                  , p9)R           _
                                      ,Pm,                           T,p       dTm, Substituting Equation 5-103 into Equation 5102 yields the following:

PP dP m

                                                                       =1                                 (5 104)

P m, dP p Separating variables yields the following: dP m = dP p (5-105) m P Integrating both sides yields the following:

                                                   ,       ,              ,     s

• PP Ln = Ln (5-106)

                                                   , Po.m ,                 P t gp ,

3681w.5.non:1b-060897 5-28

FACILITY SCALING REPORT l O or: ,U f i ! p Po.m m pP (5-107) p s 0Pe Equation 5-107 can be used to relate the pressure in the full-scale AP600 to that in APEX test facility. For the APEX model, depressurization transients are initiated from 350 psia. Thus the reference value ((Y,)m) for the model is determined to be 0.001645. For the AP600, the reference valueo ((Y )p) is based on 1080 psia. This yields a value of 0.00061. This selection is based on the assumption that the initial depressurization from 2250 psia to 1080 psia is quite rnpid for most SBLOCA scenarios. At 1080 psia, the system is at the saturation pressure corresponding to the secondary-side fluid temperature. Thus, the system pressure may remain relatively constant for a prolonged period of time

                              . or until the ADS is activated. Substituting the reference pressures into Equation 5-107 yields the following:

Papa = 0.324 fAP600) (5 108) ,r For the AP600, Po equals 1080 psia (7.45 MPa). For the APEX, P oequals 350 psia (2.41 MPa). Figure 5-6 illustrates the pressure scaling relationship between the AP600 and APEX. Having obtained this pressure scaling criteria, it is now possible to specify the initial conditions for a I depressurization transient in the model. These conditions are listed in Table 5-4. He pressure scaling relationship also permits a comparison of fluid properties in the model to those of the AP600 during depressurization transients. Thus, fluid property dependent scaling criteria can now be evaluated for depressurization transients. His is done in the next section for natural circulation behavior. 5.6.2 Fluid Property Scaling Having developed a scaling relationship for system pressure, scaling ratios can now be obtained for important fluid properties. Figure 5 7 illustrates the fluid property ratios. Note that 'hese fluid property ratios are nearly constant with respect to pressure. That is, the fluid proper , of water exhibit self-similarity over a relatively wide range of pressures. Thus, any continus s abset of fluid properties can be scaled by a single constant to re-create a set of fluid properties appchk to a much wider range of pressures. !O 'v 368Iw.5.non.Ib.060897 5-29

FACILITY SCALING REPORT Vapor Ouality Ratio The vupor quality scaling relation was developed in Section 4.0 for two-phase natural circulation conditions. It is as follows: f i

                                                                             =

(xoc)R ,ApsR b = 0.30 (5-109) Figure 5-7 indicates that the vapor quality ratio ranges from 034 to 0.26 with an average value of 0.30. Thus, during the initial phase of the depressurization transient, the vapor quality in the model is sigr.ificantly less than in the AP600 at the core exit if the power is scaled to preserve two-phase natural circulation flow rates. The void fractions, however, will be the same in the model and the AP600.

                  - Averace System Fluid Density Ratio Equation 5-8 defines an average two-phase fluid density. This equation can be written in terms of a ratio as follows by expressing the volume fmetion in terms of vapor quality at equilibrium conditions.                                                                              '

(Pe)g (PTy)R " - '- 3p (5-110) 1+x 6 Pgs,. R Since x (Ap/pss) is much greater than one, Equation 5-110 can be written as follows:

                                                                               -      (Pe)*

(Prr)R ' ' (5-lil) AP xR - 6 b5 EsR e. i l l 36siw-5.non:ib-060897 5-30

FACILTTY SCALING REPORT Substituting Equation 5-109 into Equation 5-111 yields the following:

                                                                                                                                                             ~

(PTP)R b)R } Thus, the average two-phase fluid density ratio is modeled by the saturated liquid density ratio shown

                                      - in Figure 5-7. De liquid density ratio varies from 1.03 to 1.14.

Vapor Volumetric Fraction Ratio

                                      . De vapor volumetric fraction is related to the quality as follows:

a= xv8 ' (5-113) xv,, + (1 -x) vh lTQ where: l%./

                                                           -=            specific volume of the saturated vapor '

l. l v,, .

v6 = ' specific volume of the saturated liquid 1 i Equation 5-113 can be written as follows: a= (5-114) l Ps5  ! where: l

                                                                                                                      = xvg , + (1 -x) v 6                (5-115)
 ;-                                                                                                              Pre i
      \
                                      ' 3681w-5.non:IM)60897 -                                                             5-31
 = _ - _ _ - _ _ _ _ - _ _ _ _ _ - .                                                 -_ _ _ _ _ _ _ -_-_ _ _ _ .

FACILITY SCALING REPORT and l 1

                                                                                                    *V g5 (5-116)

Egs Writing Equation 5-114 in terms of a model to prototype ratio yields the following: X (PTP)R R NW ag= (Pgs)g Substituting Equations 5-109 and 5-112 into Equation 5-117 yields the following: f i Pts (5-118) (Gnc)g " g R O where op = p,, - p g,. Since p3 is much greater than p s, g Equation 5-118 becomes the following: (a),=1 nc WW Figure 5-7 indicates that (a nc )R actually varies from 0.96 to 0.99 over the range of pressures. Equation 5-119 indicates that the volumetric vapor fraction in the model will be the same as that in the AP600 during the initial phase of the depressurization transient when two-phase natural circulation is in progress if the power is scaled accordingly. Effect of Vapor Density and Energy Chances Durine Deoressurizatio_n The following important facts regarding the vapor density and vapor volumetric energy can be discerned from Figure 5-7. O 36siw 5.non:iu)60897 5-32

FACILITY SCALING REPORT i b' dPss ,3 (5-120) dP.R and d g (Pg, h g,) =1 (5-121) Equation 5-120 can be verified by recalling that the pressure scaling ratio is 0.33 while from Figure 5-7 the average vapor density ratio is 0.32. The ratio of the two values is close to unity. Similarly, since the vapor enthalpy ratio is close to unity, Equation 5-121 is also valid. Since the vapor volume and energy changes are scaled with pressure and since the total mass in the system is regulated by controlling break flows, the following approximation is implemented: (app)R " N' l.

                    ...That is, the average vapor volumetric fraction in the model is the same as in the AP600 during depressurization.                                                                                                                                                                  'I
                                                                                                           ~

At any'given time during the depressurization, the following equation can be applied: I aDP YT"XDP Mppg , (5-123)  ; t where: app. .= average vapor volumetric fation VT- = total system volume

                                   =      average vapor quality xDP.
                   'M7             = ~ total fluid mass inventory p,g           = density'of the saturated vapor                                                                                                                                       i i
   .[J                                                                                                                                                                                                    l

1. 36sIw.5.non: b 060897 5-33 '

FACILITY SCALING REPORT Writing Equation 5-123 as a ratio yields the following: (xDP)R R (app)R VR= (5-124) (Pss)R Since the system volume ratio is fixed at 1:192 and the total fluid mass inventory is also fixed at 1:192 by controlling break mass flow rates, Equation 5-124 can be written as follows: f  % DP (app), = (5-125) r 8s , g Substituting Equation 5-122 into 5-125 yields the following: (xDP)R * (Pgs)R (5-126) Examining Figure 5 7 reveals the following: (xDp)g = 0.33 (5-127) which is essentially the same as (x,c)g. Thus, the vapor volumetric fraction ratio and vapor quality ratio remain constant throughout the two-phase natural circulation period and depressurization period. Having examined the effect of vapor density and energy changea, it is important to mention that the contribution of these effects to the depressurization rate is small compared to the effect of the break flow rate. This is demonstrated in Sections 5.7.1 and 5.7.2. G 36Siw-5.non:Ib 060897 5-34

FACILITY SCALLNG REPORT 4 5.6.3 Transition Pressure (~ O) An important aspect of the depressurization scaling analysis is the selection of the transition pressure. This is the pressure at which pressure scaling of fluid properties ends and fluid propeny similitude begins. The following criteria have been used to select a transition pressure for the OSU/AP600 test

                                - facility:

The transition pressure should be greater than the pressure at which fourth-stage ADS operation is expected to initiate. Thus, final system depressurization can be evaluated under the same fluid propeny conditions as in the AP600 and the ADS 4 valves can be sized accordingly.

                                      . Accumulator injection should be completed.

The net integrated mass and energy into the system up to the point of the transition must be properly scaled to achieve proper initial conditions. Based on these criteria,50 psia was selected as the transition pressure. This pressure is greater than that expected for a start of ADS 4. It also represents the pressure at which the accumulators empty. Section 5.6.4 explains how the core power is scaled to satisfy the integrated energy requirement. m 5.6.4 Two-Phase Natural Circulation Scaling for Depressurization Events j During a small break loss-of-coolant accident (SBLOCA) or ADS depressurization events, the loss of reactor coolant pump forced circulation results in periods of two-phose natt:ral circulation. This natural circulation behavior continues until the steam generators become a heat source to the primary system liquid or until the primary system liquid inventory is greatly depleted. This section evaluates the similarity criteria for scaling two-phase natural circulation during depressurization transients. Natural Circulation Time and Velocity Scaling Ratios Table 4-17 presents the steady-state single-phase and two-phase natural circulation scaling ratios. The scaling criteria presented in this table may be applied to slowly evolving depressurization events provided the fluid properties are evaluated at the respective system pressures.. Using the pressure scaling relationship developed in the previous section, the single-phase and two-phase velocity and  ; time scaling ratios given by Equations 4-139 and 4-140 can be evaluated. Figure 5.8 presents the natural circulation r,caling ratios as a function of model pressure. The two-phase time and velocity l scaling ratios are independent of pressure, remaining constant at a value of 0.5. The time and velocity l scaling ratios for those regions of the primary system which remain single-phase liquid vary as a l . function of pressure and are distorted relative to their counterparts in the two-phase regions. Fortunately, the single-phase scaling ratios do not vary greatly, thus reducing the complexity of 3681w-5.non:ll>.060897 5 35

FACILITY SCALING REPORT analyzing their impact on system behavior. Furthermore, during the depressurization events, most of the primary system contains two-phase fluid, thus reducing the distortion. Natural Circulation Core Power Scaling Ratio Equation 5-97 presents the fluid property group that has been used to relate pressure in the model to pressure in the AP600. This equation can be written in terms of a fluid property ratio as follows: Yg= U B-@ Pgs Ptsh is. R This ratio scales the fluid properties in each of the coefficients (that is, Equations 4-106a through 4-106c) of the cubic velocity equation developed to predict two-phase natural circulation under steady-state conditions. Rearranging Equation 5-101 yields the following: Yg =(Yo)g (MM) O Because the transient initiation reference pressures have been previously set, the value of(Y,)g is constant. Thus: Yg =2.70 (5-130) The core power scaling ratio for natural circulation can be determined by substituting Equation 5-128 into Equation 4-138. This yields the following:

                                                         "8   f    YR                                  (5-131)

[9 core)R R O 36s t w.5.non. iso 60897 5-36

l FACILITY SCALING REPORT l l For the model, ag equals 1:48 and fg equals 1:4. Substituting these values and Equation 5-130 into Equation 5-131 yields the following: [qcore)R N' Thus, the core power ratio is constant with respect to system pressure. During a depressurization transient, the AP600 core power is generated by decay heat. Core decay heat scaling is examined in Section' 5.6.6. 5.6.5 Scaling Critical Flow Behavior l The aim of this section is to develop a reasonable method of sizing the breaks and vent paths for the APEX facility. Because the APEX is a reduced-pressure test facility, it is not possible to duplicate exactly the critical flow behavior of the AP600 for transients initiating at full AP600 pressure. Derefore, a phase level analysis to scale the local phenomena associated with critical flow at high pressure is not performed. The goals of this section are much more modest and appropriate for I integral system testing. He goals of this section are as follows: l a

                                                                                              ' To indicate that the break and path sizes should be selected such that the sequence and timing  )

of transient events are preserved on a scaled basis. ! e To indicate that the system mass inventory, subsequent to blowdown, should be preserved on a scaled basis. His mass inventory represents one of the initial conditions for long-term i l cooling. To indicate that the break and path geometries should be such that the discharge coefficient is well known. His reduces the uncertainty in modeling and permits a better assessment of the computer codes being benchmarked. l It will be shown that the theoretical development presented in Section 5.3, coupled with a two-phase critical flow model, can be used to meet these goals. De following topics are discussed in this section:

• Subcooled critical flow e Integrated mass inventory a Break and vent path geometry

                                                                                          . Two-phase critical flow a     Subcritical flow s
                                                                                . 36aiw-5.non: b-060s97                                          5-37

FACILITY SCALING REPORT Subcooled Critical Flow At steady-state operating conditions, the fluid in the AP600 is subcooled relative to the system pressure. A leak in the system at inese conditions results in subcooled critical flow at the leak location. This results in a rapid depressurization to saturated conditions. Because this flow period is short relative to the two-phase ponion of the transient, the subcooled critical flow phenomenon has not been scaled in APEX. This results in a distonion in the time it takes to reach saturated conditions I (that is, not half time scale). The mass loss during the subcooled blowdown is insignificant for larger i breaks relative to the total mass loss and therefore has no significant impact on the transient. For I smaller breaks the subcooled blowdown period can be prolonged. This distortion, however, can be easily eliminated by beginning the APEX-to-AP600 transient comparisons at the time the reference pressure is reached. Integrated Mass Inventory To assure the proper initial conditions at the onset of long-term cooling, it is necessary that the total loss out of the break during the blowdown period be properly scaled. The total mass loss during two-phase fluid blowdown is given by the following: To (5-133) hiLoss

• j CpG,a, dt l

Substituting Equation 5-55 into Equation 5-133 and moving constant values outside of the integral yields the following: t, ' y -1 MLoss = Cd G,,o a , f s 1 bt Ys dt (5-134) o Integrating Equation 5-134 yields the following: 1

                                                                 ~l Cor'e s (5-135)

MLoss * - O l

r. .e 368Iw-5.non:1b-072597 *

                                                         .                                                 Rev.I l

FACILITY SCALING REPORT l. h,O Substituting Equation 5-64 into Equation 5-135 and writing the result in terms of a scaling ratio yields

    'V                                                               the following:

L , i (MLoss)g TTP,o)R (5-136) Because the initial fluid volume will be scaled as 1:192, Equation 5-136 becomes the following: [MLoss)g

• N @-M Equation 5-137 indicates that the mass inventory subsequent to two-phase fluid blowdown would be l properly scaled, thus satisfying one of the initial conditions for simulating long-term cooling.

Break and Vent Path Geometry . The APEX is capable of simulating small breaks on the top and bottom of the hot and cold kcs and . of simulating single- and double-ended breaks on the DVI and cold leg CMT balance lines. In addition, two tiers of ADS 1-4 will bc modeled. Because the objective of the tests is to simulate integral system behavior for purposes of validating computer codes, an idealized break geometry will be implemented.' For a comprehensive examination of realistic break geometries, such as pipe cracks, the reader is referred to the work of Schrock.W I The APEX implements flow nozzles having nearly ideal discharge coefficients. That is: CD,m = 1 (5-138) I This eliminates one variable in computer code modeling and permits better comparison with other test facilities. The flow nozzles are designed to preserve the break wall thickness (f,) to diameter (D)

                                                                  - ratio.- That is:

l 1 (fw/D]R "I bM o lU L 368Iw 5.non:1b-060897 5-39

FACILITY SCALLNG REPORT Effective Flow Area Scaling Ratios for Break Flow Rate Dominated Depressurization The summation term in the system break flow scaling ratio, Equation 5-73, can be expanded as follows: (CpGo a,), + (Cp Go a)g + . = [pm], M@ where: Gi = mass flux through the individual break or vent path (Coa); = effective flow area of the break or vent path Applying the identity defined by Equations 5-95 and 5-96 yields for a single break or vent path the following: [CpGq a{,, = [pi,o]R where pi,o is the local system density at the individual break or vent path. Rearranging this equation results in an effective flow area scaling ratio for each break or vent path: l i- [Ca{,R o =1 (5-142) 96 G I 0. R This scaling ratio satisfies the system scaling requirement as specified by Equation 5-73. This equation shall be used to scale AP600 breaks equal to or greater than two inches. O 3681w-5.non:Ib-060897 5-40

FACILITY SCALING REPORT O%J Effective Flow Area Scaling Ratios for Volumetric Expansion Dominated Depressurization Equation 5-91 can be rearranged as follows': [ECpG,,oa,]g = Eh o 4. R Using the same approach as in the previous section, Equation 5-143 can be written for individual breaks or vent paths as follows: a = 'Y$ (5-144) l [Co }.R G;,ocho% R 1 This scaling ratio satisfies the system scaling requirement as specified by Equation 5-91.' This equation shall be used to scale AP600 breaks equal to or less than one inch. Critical Flow Models 1 The analysis presented in Section 5.4 reveals that the break sizes can be selected based on the initial conditions at the reference pressure. The scaling analysis is greatly simplified because the need to evaluate critical mass fluxes as a function of time-dependent pressures has been eliminated. Equations 5-142 and 5-144 can be evaluated for each break and vent path location by using a critical flow model applicable to the reference pressure conditions at that location. A wide range of critical flow models are available in the literature. The SOURCE computer code appears to be the most compn:hensive calculational tool for predicting critical flow rates for a wide range of fluid conditions, break geometries, and break orientations.(II) For ease ol application, however, the following critical I flow models have been used:

• Homogeneous equilibrium model (HEM)(30)
• Perfect gas model(4)

These critical flow models provide reasonable scaling ratios as demonstrated in Section 9.0. The HEM (Q / calculates the critical mass flux for fluids at saturated conditions for a full range of vapor qualities. Malw inon:1b-060897 5-41

FACILITY SCALING REPORT Table 5-5 presents the critical mass flux values calculated by the HEM for vapor qualities ranging j from 0 to 0.2. The perfect gas model for critical flow is given by the following the following: i

                                                                                                                       ' y +i
                                                                                                                              'H 2    Y7                                (5-145)

G,,, = yp,P,Y+1; g where: y = ratio of specific heats p,g

                                                   =                   vapor density Saturated Liquid Breaks and Vent Paths At the reference pressure, the fluid in the primary system is saturated liquid. For the AP600 reference pressure of 1080 psia, at a vapor quality of zero, the etitical mass flux is 5619 lbm/(ft2-s) using the HEM. For the APEX reference pressure of 350 psia, for saturated liquid, the critical mass flux is 2

2524 lbm/(ft -s). Thus: [Gj ,,]R " 1 l l i 3 ! The saturated liquid density at the reference pressure is 45.7 lbm/ft for the AP600 and 52.3 lbm/ft3 for APEX. Thus: l l [ pes [,g = 1.W @M Substituting Equations 5-146 and 5-147 into Equation 5-142 yields the numerical value for the effective flow area scaling ratio for saturated liquid breaks in break flow rate dominated systems. For AP600 breaks greater than or equal to two inches: [Cp a},g = M (W) l 3681w-5.non:1b-071897 5-42

FACILITY SCALING REPORT

p For saturated liquids at the reference pressures, the following property ratios are obtained:

(co) = 1.64 (5-149) I and (heg), = 1.25 (5-150) l l

                                                                   'Ihe system power is scaled as follows:

i l

                                                                                                                            *                                        ~

(9sys)R Substituting Equations 5-146, and 5-149 through 5-151 into Equation 5-144 yields the effective flow area scaling ratio fur saturated liquid breaks in volumetric expansion dominated systems. For AP600 [p breaks less than or equal to one inch:

    %)

(Coa}.R = 0.0113 (5-152) l Saturated Vapor Breaks and Vent Paths a Under large break loss-of-coolant accident (LBLOCA) conditions, high point vents (such as the ' l~ ADS 1-3 vent valves) may vent steam instead of a two-phase mixture. Under these conditions, the j'- perfect gas critical flow model would be more appropriate. For steam, having a specific heat ratio (pl.3), Equation 5-145 yields for the AP600 a value to 2,333 lbm/(ft2-s) and yields for the APEX a ' 2 value to 738 lbm/(ft -s). Thus: l I~' [Gio ]g = W ' @M n Y) i 3681w-5.nortib-060897 5-43 l

FACILITY SCALING REPORT 3 For the APEX, For the AP600, the saturated vapor density at the reference pressure is 2.445 lbm/ft 3 the saturated vapor density at the reference pressure is 0.7544 lbm/ft Thus: [pgo]R I Substituting Equations 5-153 and 5-154 into Equation 5-142 yields the numerical value for the effective flow area scaling ratio for saturated vapor breaks: [C9a{.R Fluid Property Similitude For transients in the AP600 that initiate below 400 psia, the APEX facility fluid properties would be identical to those in the AP600. The ratios of G,, hig , Eo, and ph would be identical. Thus, G g would be unity and Equations 5-142 and 5144 would reduce to the following: O [Coal,g = HO This equation is directly applicable to the ADS 4 vent valves, which will generally open below the transition pressure of 50 psia. 'llu NJ property similitude will exist. Break and Vent Path Dimensions For purposes of computer code benchmarking, the discharge coefficient ratio (C o )g will be set to unity. This permits the evaluation of the flow area scaling ratios presented in the previous sections. Table 5-6 presents the break and vent path sizing calculations for APEX. These calculations implement the fact that the diameter ratio is related to the area ratio as follows: DR = af NR O 368Iw-5.non:Ib-o72597 3 44 Rev.I

l FACILITY SCALLNG REPORT T The break and vent path sizes are specified for the local fluid conditions at the reference pressure. l Table 5-6 presents the scaling ratios for the diameter, flow area, and wall thickness to diameter ratio for breaks ranging from 1 inch to 4 inches and for the ADS 1-4 throat areas. 5.6.6 Core Decay Power Scaling Analysis no purpur of this analysis is to obtain time-dependent core power profiles for APEX that best , simulate the AP600 depressurization behavior while establishing prototypical initial conditions for long-term cooling. The following specific requirements need to be. atisfied: l The total core energy (integrated power) deposited into the system must be properly scaled to i assure prototypical initial conditions for long-term cooling. l

                                                        . When fluid property similitude exists, the core power must be scaled according to the ideal power scaling ratio,1:96. Fluid property similitude will exist subsequent to reaching the transition pressure of 50 psia.

Core Decay Power Profiles

  . /--
  ~'

, Figure 5-9 presents the time dependent decay power for the AP600. It.was developed using the American National Standard for Decay Heat Power in Light Water Reactors, 1 L (ANSI /ANS-5.1-1979).02) .j l \ l Figure 5-10 presents the ideal power for APEX assuming fluid property similitude from the onset of

                                               . the transient.. That is:
                                                                                                                                                           )
                                                                                                  ' (qco,'bdeal,R
                                              ' Ihe actual APEX core power is also shown on Figure 5-10. Approximately one minute after transient initiation, the APEX core power crosses the ideal power curve. APEX power is held constant to l             <

140 seconds to obtain the correct integrated energy. After 140 seconds the power is permitted to drop I'

                                               ' to the ideal power curve for the remainder of the transient. : Note that the time scale for the APEX       ,

decay power curve is half that of the AP600 to be consistent with the other transport processes in the facility. l Because ADS actuation occurs early in most transients of interest (that is less than 600 seconds for j a 2-inch break), adjusting the decay power to achieve pressure-scaled two-phase loop natural i circulation is not necessary since this portion of the transient is very short. 'Ihe APEX power curve 3681w-5.non lb-060897 5-45

l l FACILITY SCALING REPORT satisfies the requirement that subsequent to reaching the transition pressure to fluid property similitude, 50 psia, the power must be scaled according to the ideal power scaling ratio given by Equation 5-158. Integrated Power Profiles Establishing initial conditions for ADS 4 operation and long-term cooling requires that the total energy input into the system, to the point of reaching the transition pressure, be properly scaled. The integrated core power ratio is written as follows: dt m (Eco

                                                                                  ,,), = ['" qcore.m                          (5-159) fo 9 core.p dt p i

where: tm = time it takes in the model to reach the transition pressure tp = time it takes in the AP600 to reach the transition pressure Since the ideal core power scaling ratio is 1:96 and the ideal time scale is 1:2, the integrated core power ratio should be the following: l l (Ecor,),= M N@ l Figure 511 presents the integrated core power for the profile given by Figure 5-10. Note that the integrated power for the model is relatively close to that of the ideal integrated power throughout the transient. After reaching approximately 200 seconds, the APEX and ideal curves overlap. 5.6.7 Scaling Component Stored Energy During the transient depressurization of the AP600, a significant amount of component stored energy l is released to the fluid. The high temperature metal components that comprise the reactor vessel and l its internals are the primary sources of this heat addition. The purpose of this section is to determine l the scaling relationship between stored energy in the AP600 and the OSU model. l

                                                                                                                                      \

3681w.5.non.It4)60897 5-46

i FACILITY SCALING REPORT The average energy release rate (dE,/dt) from the system metal components during a transient is given j as follows:

                                                                                                                                                                                                                        )

dE, , { ' dT, '(5-161) dt la dt i i where: { l M, .= mass of a metal component j C,, - = specific heat of the metal { T, = is the temperature of.the component i Equation 5-161 assumes that the specific heat is constant with respect to temperature. Equation 5-?61 can be written in terms of an average stored energy release rate by assuming that all of I

                                            ; the reactcr vessel metal structures are at the same temperature. Making this assumption and writing Equation 5-161 in finite difference form yields the following:

i q t

 \ _,)

dE, ATs (5-162) dt At * i

                                                                                                          . Avg l

i where: AT, = change in metal temperature from the beginning of the transient to the end of the transient At = duration of the transient An average stored energy release rate ratio is written as follows: dE s 1 (5-163) l - dt R. Avg 5 8 8 R l l( ) 3681w-5.non: b-071897 5-47 t

FACILITY SCALING REPORT Over the course of a full depressurization transient (depressurization to IRWST injection) in the ( AP600, it is expected that the metal structure temperature would change from an initial value of 600 degrees Fahrenheit (315.6 degrees Celsius) to the saturation temperature at atmospheric conditions, 212 degrees Fahrenheit (100 degrees Celsius). For the scale model, the initial metal structure temperature would be 430 degrees Fahrenheit (221.1 degrees Celsius). Subsequent to a full depressuri-zation, the metal temperature would be 212 degrees Fahrenheit (100 degrees Celsius). Thus, the f a full depressurization transient is given by the following: (AT,)g or 1 (AT,)g = 0.5619 (5-164) The time scaling ratio, Tg, equals 0.5. Thus, Equation 5-163 becomes: dE s (5-165)

                                                                       = 1.124 [I(M,Cy ,)],

R. Avg Table 5-7 presents a list of reactor vessel metal components for the AP600 and the OSU model. The specific heat of each component, the mass of the components, and the stored energy released during full depressurization have been included. Because of the reduction in scale size, it is typically the case that the amount of stored energy in the model relative to the reactor vessel fluid volume is larger in the model than in the plant. This creates some distortions in the depressurization and flow characteristics in the model vessel relative to the AP600. To reduce the potential problems with stored - a energy release, significant efforts have been made to reduce the amount of metal mass in the model. 1 In particular, the vessel wall thickness in APEX has been reduced to a minimum (while still meeting American Society of Mechanical Engineers (ASME) pressure requirements), a low mass, high strength, ceramic has been used to simulate the stainless steel reflector in the AP600 and a Teflon coating to reduce heat transfer from the APEX lower plate flange has been used. Table 5-8 presents the downcomer stored energy results, which will be implemented in Section 7.8 to determine the downcomer gap dimensions. (This is an iterative seledion process). The downcomer mass is based on the core barrel and pressure vessel steel from the centerline of the direct vessel injection (DVI) line to the top of the lower support plate and includes one-half the mass of the reflector. O 368Iw-5.non:Ib.07:597 5-48 Rev.1

FACILITY SCALING REPORT 1 l !O Table 5-9 presents the heat capacity ratios; the average stored energy release rate ratio; and the  ! average volumetric energy release rate ratio, which is defined as follows: l l l

                                                                                          = 215.8[I(M,Cy ,)],                         NO I

When fluid property similitude exists, the following criteria should be met: 5

                                                                                       = 0.0104 (1:96)                                (5-167) dt.R and
                                                                                                                                                ]
                                                                                  .I    dE'     =20                                   (5-168)   ;
   %)                                                                                    dl . R The special case of downcomer depressurization is examined in Section 7.0.

5.6.8 Scaling the Net System Power The net system power given by Equation 5-79 can be written as a ratio as follows: [9sys)R "(9 core +%etal +9so +9PRHR)R Equation 5-94 presents the following similarity criterion for the case of fluid property similitude: [q,y,( = 0.0104 (1:96) (5-170) (~') v I 368iw.5.non:1b-060897 5-49

l FACILITY SCALING REPORT l Substituting Equation 5-169 into Equation 5-170 yields the following: [9 eon + 9 metal + 9so + 9PRHR)R Using the identity specified by Equations 5-95 and 5-96, the criterion of Equation 5-171 can be satisfied in a more restrictive manner by requiring that: (qcom), = am @@ (q,g), = 0.m @@ (qso)g

• EN @M (qpagg), = Rm @@

It is noted that three of the heat late terms are controlled by test procedures: qcoy, the heat addition rate of the core heater rods; qso, the heat removal rate of the steam generators; and qPRHR, the heat removal rate of the PRHR. Thus, the scaling criteria presented in Equations 5-172,5-174, and 5-175 can be satisfied using careful operating and measuren.ent (system energy balance) procedures. The steam generator and the passive residual heat removal (PRHR) scaling analyses are presented in Sections 5.6.9 and 5.6.11. The scaling criterion presented in Equation 5-173 depends on insulation thickness, local convective heat transfer coefficients, and insulation thermal conductivity. Variable control over heat loss could be achieved by applying strip heaters to component surfaces, if necessary. The criteria presented in Equations 5-172 through 5-175 are based on average system behavior.

 'Iherefore, the criteria can be reasonably satisfied. It is not likely that the local stored energy release rates and local component heat loss rates can be simulated exactly during transient depressurizations.

Areas requiring more detailed analysis, such as the CMTs, have been identified and addressed elsewhere in the report. For purposes of benchmarking the computer codes, satisfying the system scaling criteria given by Equations 5-172 through 5-175 yields adequate results. 36siw-5.non; b-060897 5-50

FACILITY SCALING REPORT 5.6.9 Steam Generator Scaling Analysis The AP600 will use two Delta-75 U-tube steam generators, each capable of removing 970 MWth from the primary fluid during full power operations. Each steam generator produces approximately 4.2x106 lbm/hr of steam. The OSU APEX incorporates two U-tube steam generators modeled according to the Delta-75 design. Because of the experiment design, the following functional requirements must be met: Each of the OSU AP600 steam generators must be capable of removing 350 kw of thermal energy from the primary fluid while single-phase natural circulation conditions exist on the primary side. This represents a combined energy removal rate equivalent to 3.5 percent of scaled decay power. To properly model the initial conditions of SBLOCA scenarios, the OSU AP600 steam generators must be capable of removing the core energy deposited in the primary fluid while operating with secondary-side pressures that are very close to primary-side pressures. This requires that the tube surface area be sufficiently large to permit heat transfer with small temperature differences between the primary and secondary sides. In addition to the functional requirements listed above, the following scaling requirements must also be met: The tube side of the steam generators must satisfy the length and flow area scaling requirements of Section 4.0. The steam generator dryout time (tm) must satisfy the following scale requirement: (tpo)R ~ Steam Generator Dimensions Applying the length and flow area scaling ratios of Section 4.0 to the AP600 steam generators yields the model steam generator dimensions presented in Table 5-10. By using the same tube diameters and tube pitch as in the AP600 steam generators the number of tubes can be determined from the following relation: aR

• ftube)R ( tube)R 3681w-5.non:ibaos97 5-51

FACILITY SCALLNG REPORT Since a n quals e 1:48 and (a be)R tu equals one, 1 (Ntube)g = 0.0208 (1:48) (5-178) Furthermore, the tube surface area ratio is found to be: A s.R = (NDf) tube,R Substituting (Drube)R equals one, (ftube)R equals 0.25 and (Neube)R equals 0.0208 yields the following: A,,g = 0.005208 (1:192) (5 180) Using these scaling ratios to establish the tube number and dimensions satisfies the functional requirements set for the model steam generators. Steam Generator Dryout During the course of a SBLOCA, the steam generators eventually are isolated and transitioned from being a primary-side heat sink to a primary-side heat source. A simple energy balance can be used to estimate the initial secondary-side liquid mass (Mpw) required to satisfy the steam generator dryout time scaling criterion (Too)g equal to 0.5. For a constant pressure boil-off process: Mpwhg = [*D qcomdt (5-181) O 3681w.5.norciba60897 5-52

FACILITY SCALING REPORT l ( - Rearranging and writing in terms of a model to prototype scaling ratio yields the following: I f 9 core dt (Mpw)g = W 82) j Ig}R ID qco,,dt 0 . AP600 Because the core power ratio is scaled as 1:96 and the time scaling ratio is 0.5, the initial liquid mass { in the steam generator is scaled as:

                                                                                                                                                                                                )

I 1 hiFW)R 192 (his)R i Table 5-10 presents the values of initial steam generator liquid mass for the AP600 and the model. Equation 5-183 can also be written in terms of initial liquid level is the steam generator by 4 I recognizing that: hi FW

                                                                                                                                )R = aRhFW)R                                           (5-184) l Substituting Equation 5-184 into 5-183 and rearranging yields the following:                                  l l

1 R 192aR(his )R Setting ag equal to 1:48 yields the following: I hFW)R 4 (5-186) 1 3681w 5.non:Ib-060897 5-53

l FACILITY SCALING REPORT For the case of an AP600 secondary-side pressure of 1080 psia and a model secondary-side pressure of l 350 psia, (Lpw)g = 0.2053. l 5.6.10 Pressurizer Scaling Analysis The AP600 pressurizer consists of a 1300 3ft (36.8 m3 ) cylindrical tank, which contains 1.3 MW of immersion heaters to generate a steam bubble for primary-side pressure control. The steam and liquid in the pressurizer tank are maintained at saturation conditions. 'Ihe pressurizer includes a nozzle capable of spraying liquid from the cold leg onto the steam bubble. System pressure is reduced as a result of the steam volume reduction caused by direct contact condensation on the spray. The pressurizer sprays will not function subsequent to reactor coolant pump trip and coast down. The following functional requirements must be met for APEX:

• The pressurizer must be capable of providing pressure control for the OSU APEX model.

                          . The pressurizer should be capable of providing system de-gassing.

In addition to the functional requirements listed above, the following scaling requirements must also be met:

                          . The length and flow area scaling criteria of Section 4.0.

• Average stored energy release rates in accordance with Equation 5-173.

Pressurizer Dimensions The length and flow area scaling requirements of Section 4.0 have been used to scale the pressurizer. As shown in Section 7.0 for other primary system tanks, applying the criteria of Section 4.0 will properly scale tank filling and draining rates, liquid and steam volumes, immersion heater power, and component lengths. Table 5-11 presents the pressurizer dimensions. Pressurizer Stored Energy Release Rate The pressurizer average stored energy release rate has been determined using the method presented in Section 5.6.7. 5.6.11 PRHR Scaling Analysis The AP600 design implements two,100 percent capacity, passive residual heat removal heat (PRHR) exchangers each capable of removing 2 percent of the core power using natural circulation. The C-type tube heat exchangers are located in the IRWST and can operate at full-system pressures. i 3681w.5.non:Ib.060897 5-54

FACILITY SCALING REPORT 1 The OSU AP600 includes a single C-type heat exchanger scaled using the length and flow area scaling . criteria developed in Section 4.0. Table 512 presents the scaling ratios and dimensions for the PRHR heat exchanger As can be observed from this table, all tube lengths, the total tube-side volume, and the total cross-sectional flow area have been ideally scaled. l l PRHR Heat Transfer Surface Area l l ) Because the OSU facility operates at reduced pressure and reduced height, the heat exchanger transfer l area has been selected such that 403 kw (scaled equivalent of 2 percent decay power) can be removed. I by natural circuh. tion. 5.7 Evaluation of Depressurization Specific Frequencies, Characteristic Time Ratios, and Scaling Distortions l To assess the relative importance of the various depressurization processes occurring in the system, the specific frequencies and characteristic time ratios presented in Tables 51 and 5 3 must be evaluated for both the APEX model and the full-scale AP600. A comparison of these values reveals which of I the processes are dominant and to what extent the proposed APEX break and vent path geometry and operating conditions introduce scaling distortions 7 5.7.1 Evaluation of Break Flow Rate Dominated Depressurization Processes i i All cases of ADS operation result in the opening of a valve having a throat diameter of at least

          .l. 2 inches in diameter. Therefore, based on the previous analysis of depressurization behavior, scaling i of the ADS flow area should be dictated by the same rationale used to scale the system breaks having I diameters larger than 2 inches.
              ~ Two depressurization cases are examined in this section: a two-inch cold leg break and a double-ended DVI line break. The two-inch break represents the smallest break flow dominated
               ~depressurization case, and the double-ended DVI line break represents the largest break to be studied in the facility. For breaks smaller than two-inches in diameter in the AP600, fl-i becomes increasingly greater than one, which indicates that the depressurization is no longer break flow rate dominated.

Therefore, all cases of ADS operation are break flow rate dominated. Tables 5-13 and 5-14 present the numerical values for the system fluid residence time, the specific process frequencies, the characteristic time ratios, and the distortion factors for each of these cases. An examination of Tables 5-13 and 5-14 reveals that the vapor generation rate ratio is small compared

              - to one, and thus, both cases are break' flow rate dominated. Because all of the characteristic time

( ! ratios are less than one, an examination of their distortion factors is not needed. O y As anticipated, the fluid residence time for the DEDVI break case is significantly shorter than that of the two-inch break case. For both cases, the APEX residence times are one-half that of the AP600 as required. 36siw.5.non:ib-072597 5-55 Rev.I w_______-___________-__________________-___--_ __ ___ _ _ _ _ - _ - _ _ _ .

FACILITY SCALING REPORT 5.7.2 E+aluation of Volumetric Expansion Dominated Depressurization Process A one-inch cold leg break case shall be examined in this section. Table 5-15 presents the numerical values for the system fluid residence time, the specific process frequencies, the characteristic time ratios, and the distortion factors. For one-inch diameter breaks or smaller, the volumetric expansion due to vapor generation dominates the depressurization rate. This is reflected in the value of U r, given in Table 5-15, which exceeds one for the AP600 and APEX. Table 515 indicates that, in an effort to preserve the desired power ratio of 1:96, the residence time constant in APEX is close to that of the actual AP600. Thus, the one-half time scaling ratio is used. Using a larger power ratio (1:41) would permit the use of a larger break pipe, thus preserving the one. half time scale. However, for this case, CMT draining will be governed by the system energy. That is, interruption of CMT recirculation flow and the subsequent initiation of CMT draining will depend on the rate at which steam accumulates in the cold leg /CMT balance lines. Draining the CMT will eventually actuate ADS, which places the system in a break flow dominated mode. By preserving H g, the system volumetric expansion will be properly scaled and the rate at which steam accumulates in the CMT balance lines should also be reasonably modeled. Thus, the timing for I the onset of CMT draining and ADS actuation should be preserved. See Section 6.1 through 6.3 for a 1 detailed discussion. 5.8 Conclusions This section has implemented the H2TS methodology to provide the structure for the depressurization scaling analysis. The following results have been obtained:

• A relationship between fluid pressure, properties, and depressurization rate in APEX and the AP600
• APEX initial conditions, system setpoints, break sizes and ADS vent sizes
• APEX core power l
• APEX metal stored energy
• APEX steam generator dimensions
• APEX pressurizer dimensions i l
• APEX PRHR dimensions O

3681w-5.non:lt@2597 5-56 Rev.1

FACILITY SCALING REPORT [b' ) It is shown that depressurization behavior can be divided into two main categories: break flow dominated behavior or fluid volumetric expansion dominated behavior. By implementing the depressurization rate equation, break sizes for APEX corresponding to AP600 scenarios, can be selected. An evaluation of the characteristic time ratios reveals that the break and ADS vent sizes selected will adequately model AP600 behavior. In conclusion, the scaling analysis presented in this section indicates that the depressurization behavior obtained in APEX will be representative to the AP600 and of sufficient range and quality to validate computer codes. 5.9 References

1. Reyes, J.N., A Theory of Elastic Decompression of Two-Phase Fluid Afixtures, Oregon State University, OSU-NE-9407, December 1,1994.

2. Higdon, A., et al., Afechanics of Afaterials, John Wiley and Sons, Inc., New York,1976.

3. Nadeau, G., Introduction to Elasticity, Holt, Rinehart and Winston, Inc., New York,1965.

4. Shapiro, A.H., The Dynamics and Thermodynamics of Compressible Fluid Flow, Vol.1, The

 ,3
  ,                       Ronald Press Company, New York,1953.

L )

5. Lamb, H., Hydrodynamics, Dover Publications, New York,1945.

6. Zuber, N., "On the Dispersed Two-Phase Flow in Laminar Flow Regime," Chemical Engineering Science. Vol. 19, 1964.

7. Marviken Project, "The Marviken Full-Scale Critical Flow Tests, Vol.18: Results from Test 10, Electric Power Research Institute," Palo Alto, California, May 1982.

8. Moskal, Thomas E., Examination of Scaling Criteria for Nuclear Reactor Thermd-Hydraulic Test Facilities, Ph.D. Dissertation, Carnegie-Mellon University, Pittsburgh, PA, March 1987, 1 9. Kocamustafaogullari, G., and M. Ishii, Pressure and Fluid to Fluid Scaling Lawsfor Two-Phase Loop Flow, NUREG/CR-4585 (1986).

10. Lahey, R.T., and F.J. Moody, The Thermal Hydraulics of a Boiling Water Nuclear Reactor, 2nd Edition, American Nuclear Society, LaGrange Park, Illinois,1993.

I1. Schrock, V.E., " Critical Flashing Flow in Pipes and Cracks," Dynamics of Two-Phase Flows, Begell House, CRC Press, Tokyo,1992.

                  ') 12. Decay Heat Power in Light Water Reactors, American National Standard, ANSI /ANS-5.1-1979, V                       American Nuclear Society, La Grange Park, IL,1979.

368i 5 non:1b-060897 5-57

FACILITY SCALING REPORT TABLE 51

                                ' < STEM LEVEL SCALING ANALYSIS:

SYSTEM TIME CONSTANT AND CIIARACTERISTIC TIME RATIOS Depressurization Rate Fquation: l T,y,(P f -1 dP +, gg ,+ _Qp+-Es g * (5-61) Two-Phase Fluid Residence Time (Break Uncovering Time): PTPoYTP x'= (5-62) I [di,o e Characteristic Time Ratios: S "Wm Tsys " g FVso T System Vapor Generation g, ,'N' Erig Rate Ratio (5-66) E(sh)in o System Energy Flow S "%Tsys " Rate Ratio (5-67)

                                      )

Property Ratio: 2 S7o Volumetric Dilation (5 16) (Y-1) 1 0 3681w 5.non ItM60897 5-56

FACILITY SCALING REPORT TABLE 5-2 SCALING RATIOS FOR SYSTEM DEPRESSURIZATION EVENTS DOMINATED BY BREAK OR VENT PATH FLOW RATE  ! 1 Residence Time Scaling Ratio: t,y,,a = 0.5 (5-71) I Break Flow Scaling Ratio: E G,,o D a, _1 (5-73) PTPo R Injection Flow Scaling Ratio: EGinain = t (5-76) PTPo .R Net Thermal Power Scaling Ratio: 95Y5 I

                                                                                                                       -                                                                 (5-81) hgpTPo     g __ 96                                                                                         l A

V Fluid Property Similitude Break Flow Scaling Ratio: [ECoa,]R " Injection Flow Scaling Ratio:

                                                                                                                                                                                         "U Foie=ie)R -d Net 'Ihermal Power Scaling Ratio:

f9sys R 3 (v I 3681w-5.non:ll> 060897 5-59 { L _ _ _ _ _ . - _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ - _ _ _ _

FACILITY SCALING REPORT TABLE 5 3 SCALING RATIOS FOR SYS'IEM DEPRESSURIZATION EVENTS DOMINATED BY VOLUMETRIC EXPANSION Residence Time Ratio: t sys,R = 0.5 (5-71) Break Flow Scaling Ratio: 95Y5 =1 (5-91) Eh o ls Co C1,,o a* . R f i l 9 l l l O 3681w-5.non:Ib-060897 5-60

FACILITY SCALING REPORT TABLE 5-4 IDEAL INITIAL CONDITIONS FOR A DEPRESSURIZATION TRANSIENT AP600 OSU APEX Reactor Coolant System Pressurizer Pressure: psia (MPa) 2250 (15.5) 385 (2.66) Hot Leg Temperature: "F (K) 602 (590) 420 (489) Reference Pressure: psia (MPa) 1080 (7.25) 350 (2.41) Secondary Side Steam Generator Relief Setpoint: psia (MPa) 1080 (7.46) 3'O (2.48) Steam Generator Shell Pressure: psia (MPa) 918 (6.3) 300 (2.07)  : Steam Generator Steam Temperature: 'F (K) 534 (552) 417 (487) U Safety I Accumulator Charging Pressure: psia (MPa) 715 (4.9) 231.7 (1.56) i 4 3681w-5.non:Ib-060897 5-61 e

FACILITY SCALING REPORT TABLE 5 5 HOMOGENEOUS EQUILIBRIUM MODEL (IIEM) CRITICAL MASS FLUX 0 s X p, s 0.2 Vapor Ouality 0.0 0.1 0.2 Pressure (nsia) Mass Flux (Ibm /ft2-s) 1080 5619 4405 3824 1000 5360 4157 3600 950 5155 3979 3430 900 4950 3800 3260 850 4745 3621 3090 800 4540 3443 2920 750 4335 3248 2750 700 4130 3054 2580 650 3925 2859 2410 600 3720 2665 2240 550 3487 2469 2065 500 3253 2273 1890 450 3020 2077 1715 400 2786 1881 1540 350 2524 1661 1360 300 2262 1440 1180 250 2010 1213 1004 200 1758 986 828 150 1437 755 642 100 1116 523 456 0 3681w-5.non:lb-060897 5 62

FACILITY SCALING REPORT

  -('                                                                                                                                                                           ,

TABLE 5-6 I BREAK AND VENT PATH l FLOW DIAMETERS AP600 REFERENCE PRESSURE: 1980 psia APEX REFERENCE PRESSURE: 350 psia

1. HOT LEG BREAKS

l. Scaling Ratio AP600 APEX Wall Thickness (in.) 0.160 3.25 0.521  !

l l Break Diameter (in.) 0.1063 1 0.1063  ! ! Break Area (in.2) 0.0113 0.785 0.0089 (IJD) Thickness / Diameter 1.51 3.25 4.90 l \ Break Diameter (in.) 0.160 2 0.321 l Break Area (in.2) 0.0257 3,14 0.081 l- (IJD) Thickness / Diameter 1.00 1.63 1.63 Break Diameter (in.) 0.160 4 0.642

Break Area (in.2) 0.0257 12.57 0.323 (IJD) Thickness / Diameter 1.00 0.813 0.813

, f~4 LQ 2. COLD LEG BREAKS Scaling Ratio AP600 APEX Wall Thickness (in.) 0.160 2.56 0.411 Break Diameter (in.) 0.1063' 1 0.1063 Break Area (in.2) 0.0113 0.785 0.0089 (IJD) Thickness / Diameter 1.51 2.56 3.87 Break Diameter (in.) 0.160 2 0.321 Break Area (in.2) 0.0257 3.141 0.081 (IJD)'Ihickness/ Diameter 1.00 1.28 1.28 Break Diameter (in.) 0.160 4 0.642 Break Area (in.2) 0.0257 12.57 0.323 (IJD) Thickness / Diameter 1.00 0.64 0.64 L Note: Numerical values may not reflect as-built conditions. Refer to APEX Facility Description Report. O 36si 5.non:ib-060897 5-63

FACILITY SCALING REPORT w TABLE 5-6 (Continued) l BREAK AND VENT PATH FLOW DIAMETERS

3. ADS THROAT AREA (LARGE BREAK-SINGLE PHASE VAPOR)

Scaling Ratio AP600 APEX Single Line Throat Diameter (in.) 0.104 2.42 0.253 Broat Area (in.2) 0.0109 4.6 0.050 Two Lines Combined Equivalent Throat Diameter (in.) 0.104 3.42 0.357 Equivalent Throat Area (in.2) 0.0109 9.2 0.100 2nd and 3rd Stage ADS Single Line Throat Diameter (in.) 0.104405 5.17 0.540 Throat Area (in.2) 0.0109 21 0.228908 Two Lines Combined Equivalent Throat Diameter (in.) 0.104405 7.31 0.763 Equivalent Throat Area (in.2) 0.0109 42 0.458

4. ADS 4TH-STAGE (FLUID PROPERTY SIMILARITY)

Scaling Ratio AP600 APEX Single Hot Leg - 100% flow (No failures) Flow Nozzle Diameter (in.) 0.1021 9.834 1.00 Flow Nozzle Area (in.2) 0.01042 76 0.792 Single Hot Leg - 50% flow (Single Failure) Flow Nozzle Diameter (in.) 0.1021 6.6 0.674 Flow Nozzle Area (in.2) 0.01042 38 0.396 Note: Numerical values may not reflect as-built conditions. Refer to APEX Facility Description Report. 1 O l l 368Iw.5.non:1b-060897 5-64

l FACILITY SCALING REPORT TABLE 5-6 (Continued) BREAK AND VENT PATH FLOW DIAMETERS l

                                                            ~5.      COLD LEG /CMT BALANCE LINE BREAK                                                                                                                          I Scaling Ratio                                             AP600           APEX Wall Thickness (in.) -                                                                                              0.160          0.906           0.146 2" Cold Leg Balance Line Break Break Diameter (in.)                                                                                                0.160              2           0.321 Break Area (in.2)                                                                                           0.0257                  3.14           0.081 (IJD) Thickness / Diameter                                                                                                  1.00   0.453           0.453 Double-Ended Cold Leg /CMT Balance Line Break Break Diameter (in.)                                                                                                                6.81             1.12 Break Area (in.2)                                                                                                                  36.42           0.985 Tube Length / Diameter:

Model/AP600 1.52

6. DVI LINE BREAKS Scaling Ratio AP600 APEX O' Wall Thickness (in.) 0.160 0.906 0.145 2" DVI Line Break Break Diameter (in.) 0.160 2 0.321 Break (in.2) . 0.0257 3.14 0.08 (L/D) Thickness / Diameter 1.00 0.453 0.453 Double-Ended DVI Line Break Break Diameter (in.) 6.81- 1.12 Break Area (in.2) 36.42 0.985 Tube Length / Diameter:

Model/AP600 1.52 Note: Numerical values may not reflect as-built conditions. Refer to APEX Facility Description Report. O 36stw.5 non:lw>os97 5-65

FACILITY SCALING REPORT TABLE 5-6 (Continued) BREAK AND VENT PATH FLOW DIAMETERS

7. ADS THROAT AREA (SMALL BREAK-TWO PHASE FLUID)

Scaling Ratio AP600 APEX 1st-Stage ADS Single Line Throat Diameter (in.) 0.160 2.42 0.388 Throat Area (in.2) 0.0257 4.6 0.118 Two Lines Combined Equivalent Throat Diameter (in.) 0.160 3.42 0.549 Equivalent Throat Area (in.2) 0.0257 9.2 0.2367 2nd- and 3rd-Stace ADS Single Line Throat Diameter (in.) 0.160 5.1 0.829 Throat area (in.2) 0.0257 21 0.540 Two Lines Combined Equivalent Throat Diameter (in.) 0.160 7.31 1.17 Equivalent Throat Area (in.2) 0.0257 42 1.08 Note: Numerical values may not reflect as-built conditions. Refer to APEX Facility Description Report. O1 l 3681w-5.non:lb-060897 5-66 _ _ __ . _ _ _J

FACILITY SCALING REPORT I Od TABLE 5-7 STORED ENERGY OF REACTOR VESSEL STRUCTURAL COMPONENTS AP600 OSU Model Reactor Pressure Vessel Material: CS SS Mass: lbm (kg) 5.52x105 (2.5x10 ) 5 4.2877x103 (1.945x10 ) 3 Specific Heat: Btu /lbm *F (kJ/kg *C) 0.1 (0.0862) 0.12 (0.103) 4 4 l Heat Capacity: BtufF (kJ/*C) 5.52x10 (2.16x10 ) 514.5 (200.3) l

• Energy Release: Btu 7 7 5 5 (kJ) 2.142x10 (2.26x10 ) 1.12x10 (1.18x10 )

Reactor Vessel Head Material: CS SS 5 4 Mass: lbm (kg) 1.34x10 (6.08x10 ) 1.014x103 (459.9) Specific Heat: Brullbm *F (kJ/kg *C) 0.1 (0.0862) 0.12 (0.103) 4 3 Heat Capacity: Bru/*F (kJFC) 1.34x10 (5.24x10 ) 121.68 (47.37)

• Energy Release: Btu 6 6 4 4 (kJ) 5.2x10 (5.49x10 ) 2.653x10 (2.798x10 )

m Lower Intemals l Material (1): SS SS 4 l Mass: lbm (kg) 1.33x105 (6.03x10 ) 229.1 (103.9) Specific Heat: Btu /lbm *F (kJ/kg *C) 0.12 (0.103) 0.12 (0.103) i 4 3 Heat Capacity: Btu /*F (kJ/*C) 1.596x10 (6.213x10 ) 27.49 (10.70) 6 6 3 3

• Energy Release: Btu (kJ) 6.192x10 (6.532x10 ) 5.99x10 (6.32x10 )

Material (2): MgO Mass: lbm (kg) 13.94 (6.32) Specific Heat: Bru/lbm- F (kJ/kg *C) 1.180 (1.017) r Heat Capacity: Btu /*F . (kJ/ C) 16.45 (6.427) 3 3

• Energy Release: Btu (kJ) 3.58x10 (3.78x10 ) '

Note: Numerical values may not reflect as-built conditions. Refer to APEX Facility Description Report. 1 L) 368Iw-5.non:Ib-071897 5-67

l FACILITY SCALING REPORT TABLE 5 7 (Continued) , STORED ENERGY OF REACTOR VESSEL STRUCTURAL COMPONENTS AP600 OSU Model Reflector (Includine Lower Core Barrel) Material (1): SS SS 5 4 Mass: lbm (kg) 1.42x10 (6.44x10 ) 370.57 (168.06) Specific Heat: Btu /lbm *F (kJ/kg *C) 0.12 (0.103) 0.12 (0.103) 4 3 Heat Capacity: Bru/*F (kJ/*C) 1.704x10 (6.64x10 ) 44.47 (17.31) 6 6 3 4

• Energy Release: Btu (kJ) 6.612x10 (6.975x10 ) 9.69x10 (1.02x10 )

Material (2): PSM Mass: lbm (kg) 383.13 (173.76) Specific Heat: Btu /lbm *F (kJ/kg *C) 0.2 (0.172) Heat Capacity: Btu / F (kJ/ C) 76.63 (29.9) 4 4

• Energy Release: Btu (kJ) 1.67x10 (1.76x10 )

Uoner Intemals Material: SS SS 4 4 Mass: lbm (kg) 8.7x10 (3.946x10 ) 947.44 (429.7) Specific Heat: Bru/l:~ T (kJ/kc *C) 0.12 (0.103) 0.12 (0.103) 4 3 Heat Capacity: Bru/*F (kJ/ C) 1.044x10 (4.M5x10 ) 113.69 (44.3) 6 6 4 4

• Energy Release: Btu (kJ) 4.05x10 (4.27x10 ) 2.48x10 (2.61x10 )

1.048x106 5 3 3 Total Mass: lbm (kg) (4.755x10 ) 7.246x10 (3.286x10 ) Total Energy Release: Btu (kJ) 4.347x107 (4.587x10 ) 7 1.993x10 5 (2.102x10 ) 5

    \ssumes a full depressurization to IRWST injection with uniform metal temperatures.

Note: Numerical values may not reflect as-built conditions. Refer to APEX Facility Description l Report. CS - Carbon Steel SS -- Stainless Steel l PSM -- High Strength Ceramic O 368iw.5mn:ib-071897 5-68 1

FACILITY SCALING REPORT D TABLE 5-8 l DOWNCOMER STORED ENERGY l l AP600 OSU Model Reactor Pressure Vessel Material (1): SS SS 5 5 Mass: lbm (kg) 4.069x10 (1.845x10 ) 2.783x103 3 (1.262x10 ) Specific Heat: Btu /lbm *F (kJ/kg *C) 0.12 (.103) 0.12 (.103) Heat Capacity: Btu /*F (kJ/*C) 4.883x104 4 (1.90x10 ) 334.0 (130.0) 7 7 4 4

• Energy Release: Btu (kJ) 1.895x10 (2.0x10 ) 7.28x10 (7.68x10 )

Material (2): PSM ~ Mass: lbm (kg) 191.6 (86.88) Specific Heat: Btu /lbm *F (kJ/kg *C) 0.2 (0.172) i Heat Capacity: Bru/*F (kJ/*C) 38.33 (14.94) {

• Energy Release: Btu (kJ) 8.356x103 3 (8.809x10 )

i Total Mass: lbm (kg) 4.069x105 (1.845x10 ) 5 2.975x103 3 (1.349x10 ) Total Energy 9 Release: Btu (kJ) 1.895x10 9 (2.0x10 ) 8.ll6x10 4 4 (8.56x10 )

• Assumes a full depressurization transient with uniform metal temperatures.

Note: Numerical values may not reflect as-built conditions. Refer to APEX Facility Description Report. SS -- Stainless Steel PSM - High Strength Ceramic i l 1 36siw.5.non:ib-060897 5-69

FACILITY SCALLNG REPORT IIEAT CAPACITY RATIOS AND TOTAL ENERGY RELEASE RATIOS TABLE S 9 Ol' FOR REACTOR VESSEL STRUCTURAL COMPONENTS E #s (MCy ,)g dt s 1V dt g R. Avg R. Avg l l Reactor Pressure Vessel: 1:107.3 1:95.6 2.01 Reactor Vessel Head: 1:110.1 1:98.0 1.96 Lower Internals: 1:363.2 1:323.5 0.59 Reflector: 1:140.7 1:125.3 1.53 Upper Internals: 1: 91.8 1:81.7 2.35

                               *Downcomer                                                                1:133.1                                                                                                                           1:116.7                         1.65 Reactor Vessel Totals:                                                    1:122.5                                                                                                                          1:108.9                          1.76
                             *Used in Section 7.8.2.

O O 368Iw-51 ion Ib-060897 5-70

FACILITY SCALLNG REPORT

  'O
  -U                                                                             TABLE 510                                                     i l

STEAM GENERATOR SCALING RATIOS AND DIMENSIONS 1 Ideal OSU Model l Scaling i Parameter Ratio - AP600 Ideal Actual (TBM) l Total Tube - 3 1:192 895.8 ft 4.666 ft 3 4.626 ft 3 I Volume 3 3 25.3 m 0.132 m 0.131 m3 Number of Tubes . 1:48 6,307 132 133 l! Tube Inside 1:1 0.607 in 0.607 in 0.607 in Diameter ' l.542 cm 1.542 cm 1.542 cm Tube Outside 1:1 0.687 in 0.687 in 0.687 in Diameter 1.745 cm 1.745 cm 1.745 cm Average Tube 1:4 848.13 in 212.0 in 207.68 in

. Length 2154.3 cm 538.56 cm 527.50 cm Cross-sectional 1

48 12.67 ft2 0.264 ft2 0.267 ft2 Flow Area (Tube 1.18 m 2 0.025 m 2 0.025 m2 l

l size) 4 I Total Inside' 1:192 7.083x10 ft 2 368.9 ft 2 365.78 ft2 I Surface Area 6581.0 m2 34.3 m2 33.98 m2  ; Steam Shroud 1:6.93

                                                                                   "'* I"                  '
                                                                                                             *      I"              **    I" Inside Diameter                                                                em               ,

em em Tube Pitch 1:1 0.98 in 0.98 in 0.92 in (Triangular) 2.49 cm 2.49 cm 2.34 . cm Tube Pitch /O.D. 1:1 1,43 1.43 1.40 Ratio Heat Removal 1:96 2.31x108 B;u/hr 2.39x106 Bru/hr 2.39x106 Btu /hr Rate (~3.5% 67.66 MW 700 kw 700 kw Decay) Steam Generator *1:233.8 1.0745x105 lbm 459.58 lbm 459.58 lbm i Initial Liquid 4.873x104 kg 208.43 kg 208.43 kg Mass ?

• Based on AP600 secondary side pressure of 1050 psia and model secondary side pressure of 400 psia.

Note: Numerical values may not reflect as-built conditions. Refer to APEX Facility Description Report. b G l l l 3681w 5.non:lb-060897 5-71

FACILITY SCALING REPORT TABLE 5-11 PRESSURIZER SCALING RATIOS AND DIMENSIONS Ideal OSU Model Scaling Parameter Ratio AP600 Ideal Actual (TBM) Pressurizer 1:192 1300.0 ft3 6.77 ft 3 6.82 ft 3 Volume 36.8 m 3 0.19 m 3 0.19 m3

                                                                                        'ac                                  'ac in                           in                    a.c       in Shell Inside                  1:6.928 em                           em                              em Diameter
                                                                                        'ac               in                 'a c in                           'a.c      in
                 *Overall Intemal             1:4 cm                          cm                                 em Length 4                              4 Immer:: ion Heater            1:96       4.437x106 Btu /hr                                        4.622x10         Btu /hr       4.622x10        Bru/hr Power                                    1.30                                       MW         13.54               kW           13.54             kW Wall Thickness                1:3.8
                                                                              '8c                       in                 %c in                           'a c in cm                           em                                 em 5

Mass of 1:107.9 1.197x10 lbm 1109.4 lbm 432.5 lbm 4 Pressurizer 5.427x10 kg 503.1 kg 196.2 kg Cylindrical Shell 6 4 4

                ** Stored Energy              1:96       5.57x10                                    Btu           5.80x10          Btu           1.13x10          Btu 4

Release 5.87x10 6 kJ 6.12x104 kJ 1.19x10 kJ Note: Numerical values may not reflect as-built conditions. Refer to APEX Facility Description Report.

        *OSU model uses elliptical end caps instead of hemispherical heads.
        ** Assumes full depressurization transient with uniform metal temperatures.

t l l O i l 368Iw-5.non:Ib-060897 5-72

FACILITY SCALING REPORT , (^\  ; () TABLE 512 l PASSIVE RESIDUAL HEAT REMOVAL HEAT EXCHANGER SCALING RATIOS AND DIMENSIONS (SINGLE HEAT EXCHANGER) Ideal OSU Model Scaling Parameter Ratio AP600 Ideal Actual (TBM) Total Tube Volume 1:192 Number of Tubes -- Tube Inside --- Diameter _ _ Tube Outside --- 0.75 in -- 0.375 in Diameter 1.91 cm 0.953 cm Average Tube 1:4 Length Average Vertical 1:4 Length

         ]                                   Average Horizontal Length 1:4 Cross-sectional         1:48 Flow Area Total Inside Surface Area Total Heat Load         1:96              1.319x108 ~ Btu /hr     1.374x106     Btu /hr    1.374x106 Btu /hr (2% Decay Heat)                          38.66        MW         0.403         MW          0.403                                 MW Maximum Flow            1:96            117.0       lbm/s        1.22           lbm/s      1.22                                  lbm/s Rate at 2% Decay                         53.2       kg/s         0.55           kg/s       0.55                                  kg/s Note:      Numerical values may not reflect as-built conditions. Refer to APEX Facility Description Report.

i O V L l l 368Iw-5.non:1b-072197 5-73 w __ . _ _ - _ - _ _ _ _ _ _ - -

FACILITY SCALING REPORT TABLE 513 EVALUATION OF TWO PHASE FLUID DEPRESSURIZATION RESIDENCE TIMES, CHARACTERISTIC TIME RATIOS AND SPECIFIC FREQUENCIES FOR A TWO INCH COLD LEG BREAK (1) APEX AP600 System Reference Pressure: 350 psia 1080 psia System Fluid Residence Time:(2) t,y,,, 2154 s 4308 s System Process Specific Frequencies: com 4.83x104 s-1 2.77x10-6s 'I co r 1.07x10 4 s'l 1.53x104 s'l g 2.24x10'7 s'l 1.90x10'7 s'l Characteristic Time Ratios: H ml) 0.010 0.012 Up 0.231 0.660 IQ3) 4.84x10 4 8.18x104 Distortion Factors (DF): All Characteristic Time Ratios < l.0 (I) Assumes bottom of cold leg break without ADS. (2%ased on volume of fluid above break elevation. (3%ased on CMT flow equal to break flow. O l l l ( l l O 368Iw.5 non:Ib-060897 5 74

I l l FACILITY SCALING REPORT l l 1

                                                                                                                                                                  \

l (p\) TABLE 5-14 l EVALUATION OF TWO PHASE FLUID DEPRESSURIZATION RESIDENCE TIMES, l CIIARACTERISTIC TIME RATIOS AND SPECIFIC FREQUENCIES FOR A DOUBLE ENDED DVI BREAK APEX AP600 l System Reference Pressure: 350 psia 1080 psia System Fluid Residence Time:(I) Tsys.o 92.9 s 185.8 s System Process Specific Freauencie_g: (om 9.93x10 4 s-3 5.68x104 s-1 1.07x104 s'l 4 ~l (or 1.53x10 s og 4.61x10-5s 'l 3.90x10 5 s -1 Characteristic Time Ratios: TIm( ) 0.0922 0.1055 TI r 0.0966 0.0285 FI() h 0.0043 0.0072 i Distortion Factors (DFh All Characteristic Time Ratios < 1.0 (1} Based on volume of fluid above break elevation. (3) Based on maximum accumulator flow.

                                                                                                            .o
 ,Ph 3681w-5.non:lb060897                               5-75

FACILITY SCALING REPORT TABLE 515 EVALUATION OF TWO PIIASE FLUID DEPRESSURIZATION RESIDENCE TBIES, CIIARACTERISTIC TIME RATIOS AND SPECIFIC FREQUENCIES FOR A ONE INCII COLD LEG BREAK APEX AP600 System Reference Pressure: 350 psia 1080 psia System Fluid Residence Time:(3) T,y,,, 20238.1 s 17232.9 s Volumetric Dilation: co 8.79 5.36 System Process Specific Frequencies: com 5.15x10~7 s'1 2.95x10'7 s'l coy- 2.14x104 s'l 1.53x104 s'1

                                                                                                                                                                                                          -1                       -1 g                                                                                                         4.78x10'8 s             2.02x10-8  3 Characteristic Time Ratios:

Um 0.0104 0.0051 Ur* 4.34 2.64 Uh 9.67x10 4 3.48x10 4 Upc o 0.493 0.493 Distortion Factors (DFh System Vapor Generation Rate -0% Ratio (Ugeo)

• Dominant process.

(l@ reserves power scaling ratio of 1:96. O 368iw.5.non:1b-060897 3 76

FACILTTY SCALING REPORT O SYSTEM DEPRESSURIZATION PHENOMENA l 1 1

                                                                                                                                                   ~

Top-Dows / System Botton UpIProcess Scaling Scallag

                                                                                                                                                                      . Pleid Property Scallag
                                                                                                                                                                      .            e Natural

• System Depresserization Rate , ,

                                                                                                                                                                      . Critical Plow
               .                                   Two Phase Natstal                                                                                                  . Core Decay Power Circulation                                                                                                        . Componest Stored Baergy
                                                                                                                                                                      . Sisaa Generator Scallag
               .                                    Not System Power                                                                                                  . Presseriter Scallag
                                                                                                                                                                      . PRHR Scallag O                                                                                                   "                                     "

System Depressurisation 11Groeps and Siellarity Criteria i i 1r Evaluate Scallag Distortless I !- o Pressere Setpoista, Core Power. and , ADS / Break Sise Specifications O Figure 5-1 Scaling Analysis Flow Diagram for System Depressurization

368iw.5.non:Ibe60897 5-77

FACILITY SCALING REPORT O i 3 . e>. w N *U,

                                       *i.

ea O ............................................................ ... Z a o W e O m M

                                           "(                                                                                                                            )

4 O kO - O A ................... g ........................................... m Gl

                                           *ll, E         8.ao i

2 jl5 80 8) Z v einssey Figure 5 2 Typical Depressurization Curve for Small Break Loss of Coolant Accidents O 3681w 5.non:Ib-060897 5-78

l FACILITY SCALING REPORT i o I i 2 [ Vapor Bubbled m i Vb ' Pb i

                                                                                                                                   ,1 O                                   O O                                                             0 0                            0        O O

Two-Phase Om ) o Mixture o' E' i O O O O $ 0 0 o o TP [ o O 'O O O O O O O a O O O . i O O j 1r 1r [u p, c, a,(Break Flow Rate) i 4 f I Figure 5 3 Control Volume Representation of the Primary System J

                                                                                     -                                              l 3681w 5.non:Ib-060897                                                5-79

m FACILITY SCALLNG REPORT CRITICAL MASS FLUX RATIO (lsentropic Expansion) 1

                                                                                                                                                           - ~ ~ -                -----
                                                                                                                     -----          ---~~-

0.8" -~~~~~~~

                                                                                                      ~ ~ - - - - -                             ,

g 0.6" 3 - a m in 0.4"- - - - - - - - - - - - - - - - - - a QQ. s .- . C . 0.6 0.8 1 0 02 0.4 PRESSURE RATIO

                                                                                                                                                                                                             .i Figure 5 4 Comparison of Critical Mass Flux Ratios (G/G,) versus Pressure Ratio (P/P) as                                                                                                                ,

Predicted by the HEM (Boxes) for Isentropic Expansion and by Equation 5.43 i (Solid Line) 36aiw.inomib.osos97 5-80

l f l FACILITY SCALING REPORT l i MARVIKEN TEST #10: Large Break LOCA (Saturated initial Conditions) 6 5'L-*- m ---

                                                                                                        . 4. y           ....     . . ...... -. _ .
                                                                                                                                  - ~ ~ - - - - - -

3"- ~ ~ ~ ~ - - - - ~ ~ - ~ ~ ~ - - - - - - - - - ~ ~ ~ ~ ~ 2- - - - - - - - - - - - - - l l j. 0 10 $0 50 100 0 1'O 2'O $0 40 50 50 TIME (Seconds) l I l l J Figure 5 5 Comparison of Equation (5.57) (solid line) to Marviken Data (boxes) and RELAP5 Calculation (crosses) 368iw.5mn:1b.060897 5-81 L______---_______________ _ _ _ _ _ _ _ _ _ _ _.

FACILITY SCALING REPORT l l 1 l O sww AP600 AND APEX PRESSURE RELATIONSHIP 1200

                                                                                                                                                                                                                                         ~ ~ ~ ~ ~ " * " " "

1000-~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ " * ~ ~ " " " " * " * " " " * * ~ ~

                                                                                                                                                                                                                                    . . . .          ~ . . . .

g . . _ - . W g .. . .... . ... W W Q. y ...............-.~..~..~-~~.~...-~~~~~~~~~~"."~.

                                                                                                                                                                                                                                 ~ . . ~ . - ~ ~ ~ ~ -

g . . . . . . . . . .. .- . - . ~ .. ~ ..... ~ . .- G O 50 100 150 200 2'50 2' E @ APEX PRESSURE (psia) , Figure 5 6 Pressure Scaling Relationship Between the AP600 and the APEX Model O 3681w.5.non: b-060897 5 82

FACILITY SCALING REPORT l l l I l FLUID PROPERTY MTIOS ,. 12 - -

                                                                                                                            -a i                                                                      _
                                                                                             =   =    -      -       -

t M_- = - - l j. .

                                                                               ..m..
                                                                                       ..m.

o

e. =. = ._ x .....::

l . , 1 1 0.8-- ~g_ ~ . ~ ~ ~ _ . ~ ~ ~ _ ~ ~ ~ _ . ~ ~ ~ _ ~ . ~ ~ ~ _ . ~ ~ _~ _ ~ ~_ ~ ~ _ . ~ ~ ~ ~ ~ ~ . - - .. W 0.6- ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ - - ~ ~ ~ ~ ~ ~ - ~ ~ ~ ~ ~ ~ ~ ~ ~ - ~ ~ - ~ ~ ~ ~ ~ 0 ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ - - - ~ ~ - ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ - l (v) 9 4-~2 2 E E E  :  ; = = =  :: , D - - -

                                                                                                      -      =       =       ::    j d 02- - - - - ~ ~ ~ ~ ~ ~ - - - ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ - ~ ~ - ~ ~ ~ ~ -                                       ;

1 0 0 5'O id0 150 2d0 250 30 350 MODELPRESSURE(psia) L 1

                   -*- Vapor Quality Ratio -+- Void Fraction Ratio -*- Vap. Density Ratio
                   -e- Liq. Density Ratio -M- Vap. Enthalpy Ratio -*- Liq. Enthalpy Ratio t

1 ll O v Figure 5-7 Fluid Property Scaling Ratios as a Function of APEX Model System Pressure 3681w.5.non:Ib 060897 5-83

FACILITY SCALING REPORT I O i SCAUNG RATIOS vs.MODEL PRESSl'RE (Depressurization Transients) 1

                                                                                                                                                                  +

SP Phase Trne

                                                                                                      ^
                                                                                                           - - ;--                         *-- l-~~~

• 0.8--- :: -

                               '---^^                                                                  ~

SP Phase Velocity

                                                                                                                                                                 -e-TP PhaseTkne 0.6 - ~ ~ - ~ ~ ~ - - - - - ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ - ~ ~ ~ ~ ~
                                                                                                                                                                 -M-
                         ===                                   =                                      =       =                            =    =        ::      TP Phase Velocity      g 0.4 - ~ ~ ~ ~ ~ ~ ~ ~ ~ - - - - ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ - -
                         * * *                                                                        =       =                            =    =        ::

0.2 - - ~ ~ - - - - - - - - ~ ~ - - - - - - - ~ ~ ~ - - - - - - - 0 . . .- . . . 0 50 100 150 200 250 300 350 MODELPRESSURE(psia) Figure 5 8 Natural Circulation Scaling Ratios as a Function of APEX Model Pressure for G Depressurization Transients l 36alw 5.nonso60897 5-84

FACILITY SCALING REPORT O AP600 DECAY POWE9 14C 12 yg t- -- --- 114 100 ~g E O II~- N 76llt

                                                                                  ~,

1* - N . l g . . . . .g . . . . .g . u{g i a i "{g i i i i 'fd meesecmos) O Figure 5 9 AP600 Time Dependent Decay Power Based on 1979 ANSI Standard 3681w 5.non:Ib-060897 $.85

FACILITY SCALING REPORT O l APEXCORE DECAY POWER , (1979 ANS Decay Power) 1400 1327 1siCG 1138 m. { iuw--

                           \                                                           m lu

u. .== =_ G

l. '

                                                                                       -    g>

siDG 190

g. . . . . . . . . . . . . . . . . . . . . . . . . . . . .

mecr<noe

                              -*- DEALPOWER-*- APEXPOWER O

Figure 5-7" APEX Decay Power Profile 3681w-5.non:Ib-060897 5-86

FACILITY SCALING REPORT O l l

                                                                                                                                                                     )

i APEX CORE INTEGRATED POWER (1979 ANS Decay Power) 180 --171 180- --  ! 152 9 140 - - - 1m B. 120 - - 114 as O $@E1100- ~ ac

                                                                                                          /._
                                                                                                                                                      ~

37 hE

• 20

[ m 19 - l o' so 160 1iio 260 250 S - TIME (SECONDS)

                                                                                                     -*- 1979 ANS -e- APEX                                           '

I I O Figure 511 Integrated APEX Core Power 36stw-5.non:ib-072197 5-87

FACILITY SCALING REPORT 6.0 CORE MAKEUP TANK SCALING ANALYSIS This section presents scaling analyses for all of the core makeup tank (CMT) recirculation and draining processes. The AP600 design utilizes passive methods to perform core and containment cooling functions for a postulated loss of primary system coolant. The CMT shown in Figure 6-1 is an important feature of the AP600 passive safety system. Each tank stores cold borated water at - reactor coolant system (RCS) pressure that can be gravity-injected into the RCS to provide reactivity control and core cooling. The CMTs provide the same function as the high-pressure safety injection system in existing pressurized . water reactors (PWRs), except that current plants require the availability

                  'of AC power to perfctm their safety function, whereas the CMTs perform this function using only

gravity-driven flows.

The CMTs are connected to the RCS, as shown in Figure 6-1, by isolation valves en the cold leg balance line and the CMT discharge line. The cold leg balance line and CMT discharge valves open on an S- (safety) signal and remain open. During normal operation, the CMTs are completely filled. I

                   .In addition to adding coolant and boron to the reactor systems, the CMTs have an additional safety .

function. As the tanks continue to drain, they indicate that an unrecoverable loss-ef-coolant accident i

               !- (LOCA) has occurred. ~ When approximately 33 percent of the tank liquid has drained, the CMT level p'            sensing device activates the first stage of the automatic depressurization system (ADS) and the plant begins'a controlled blowdown through the ADS valves into the in-containment refueling water storage tank (IRWST). The ADS 2 and ADS 3 valves open in sequence based on a timer activated by the opening of the first stage of ADS valves (ADS 1). If the CMTs continue to drain, the ADS 4 valves, which are located on the hot legs, open when their volume reaches 20 percent, providit.g a large vent path directly to containment to depressurize the RCS. As the RCS depressurizes, the CMTs continue to add coolant to the RCS to maintain core cooling during the depressurization.

3 Each AP600 CMT, shown in Figure 6-2, is a 2,000 ft tank fabricated of hemispherical heads and a cylindrical shell. The hemispheried heads are carbon steel, lined with 3/8-inch stainless steel, for a total wall thickness of 4.6 inches. The cylindrical portion of the tank is also stainless steel clad, having a total wall thickness of 7.78 inches. The top of the tank is located 28 feet above the RCS

                  . cold legs. This elevation difference provides the gravity head to drive the flow into the reactor vessel downcomer. There is a cold leg balance line attached to the top of the CMT and a drain line at the l

bottom. The drain line is connected through an isolation valve and two check valves to the direct  ! Lvessel injection line. The cold leg balance line is an 8-inch Schedule 160 pipe with an inside diameter

                  ' of 682 inches. There is an isolation valve near the top of this balance line. The AP600 piping schematic for the CMT and its balance lines is shown in Figure 6-3.

p ' There are two modes of operation for the CMTs: recirculation and draining. During the initial phase

                  / of a small break LOCA (SBLOCA), a steam line break, or a steam generator tube rupture event, the RCS inventory is still at or near the steady-state value. When an S-signal occurs (typically because of f _ () -

low press'urizer pressure), the reactor coolant pumps trip and the CMT isolat;on valves open. The 36siw.6.non:1b.072597 6.}- Rev.I f

       = _ _ -         . _ - . - . _ -

FACILITY SCALING REPORT elevation head in the pressurizer is sufficient to force fluid into the top of the cold leg balance line, f i condensing any steam there such that a complete flow path is formed between the reactor cold leg, the , cold leg balance line, the CMT, and the direct vessel injection line. With the liquid circuit formed, f I the:e will be a gravity-driven flow from the CMT to the reactor vessel and a return flow from the cold leg to the top of the CMT. This is the recirculation phase of operation for the CMTs. The colder, denser CMT water drives flow into the reactor vessel because of the density difference (Ap/p) between the CMT water and the cold leg balance line (approximately 20 percent with Tcold at 550 degrees Fahrenheit, and CMT water at 120 degrees Fahrenheit). Figure 6-4, from the AP600 Standard Safety Analysis Report (SSAR) calculations,W shows the calculated CMT draining flow during the recirculation period for a 2-inch diameter cold leg break. This flow continues and steadily decreases as the colder CM1. vater is replaced by hotter water from the balance line, thereby decreasing the gravity draining head. As the break continues to drain the RCS, the cold leg balance line begins to void (as seen in Figure 6-5), the recirculation flow path is broken, and the CMT drains as the water volume is aplaced by steam from the cold leg. This begins the draining mode of the CMT. The CMT injection flow rate is larger in this mode because of the greater density difference between the colder CMT water and the steam or two-phase mixture in the balance line. A hot liquid layer forms at the top of the CMT; its thickness depends on how long the CMT is in the recirculation mode. This reduces the steam condensation when the CMT transitions into the draining mode because the hot liquid layer can flash as the RCS depressurizes, causing mixing and reducing the effects of condensation in the CMT. If the break is larger, the recirculation period is reduced because the cold leg balance line voids sooner, breaking natural circulation. In these cases, the hot liquid layer in the CMT is thinner or may not exist. For large break LOCAs (LBLOCAs) there is no recirculation period since the entire RCS voids quickly. The accumulators begin injecting very early in a LBLOCA transient and have sufficient gas pressure such that they close the check valves in the CMT discharge line until the accumulators have finished injecting. The CMTs begin injection after the accumulators are empty. By this time, the reactor vessel is refilled, the core peak cladding temperature (PCT) has been reached, and most, if not all, of the core is quenched. 6.1 CMT Phenomena This section provides a description of two important CMT phenomena: CMT recirculation and CMT draining. 6.1.1 CMT Recirculation Thermal-Hydraulic Phenomena The CMT recirculation is driven by the density difference between the cold CMT water and the hotter cold leg balance line water. The result of the recirculation is to provide colder, denser water to the reactor vessel from the CMT, which is replaced by hotter, less dense cold leg water from the balance line. There is a net mass transfer from the CMT to the RCS due to the density difference, as well as a 3681w-6.non;1ts060897 6-2

1 FACILITY SCALING REPORT  ! i net energy transfer from the RCS back to the CMT. The rates of energy and mass transfer will { depend on the buoyant differences and the hydraulic resistances in the flow path. The recirculating flow continuously diminishes with time as the CMT heats up and the resulting buoyant head decreases. As the cold leg piping and cold leg balance line void, the buoyant head increases, which increases the discharge flow as the tank begins to drain. Both single-phase and two-phase recirculation occur for small break LOCAs, while single-phase and perhaps two-phase recirculation may occur for steam line breaks and steam generator tube ruptures. There is CMT wall heat transfer occurs during recirculation where the hot fluid from the cold leg balance line transfers heat to the initially cold CMT walls. This heat transfer pracess serves to cool the recirculating fluid. Circulation patterns form within the hot liquid layer. The cooler water at the walls flows downward and mixes within the layer. As the flow recirculates, the CMT walls heat up,

         ~

and the potential for large condensation to occur is reduced at a later time when the CMT drains.

    - Also,if the CMT walls become fully heated when the RCS depressurizes, there can be reverse heat transfer from the CMT walls back to the fluid in the tank. Because the walls are thick, the amount of heat transfer to the wal'a is scenario-dependent. This was observed in the ROSA-V experiment by 1

Yonomoto et al. (1993).(2) Fluid mixing can also occur at the top of the CMT during the recirculation phase of the transient. 'Ihe f- hot liquid from the cold leg PBL is injected into the CMT through the nozzle and mixes with the k initially colder water in the CMT. The analysis of the ROSA-V experiment by Yonomoto et al. (1993)(3) indicates that the mixing behavior is one-dimensional in nature and that conduction effects between fluid layers are negligible, while the convection effects appear to dominate. As recirculation continues, the PBLs begin to void and a two-phase mixture is swept into the CMTs. Depending on how long recirculation has occurred, the CMT walls and the mixed liquid layer could be hot or cold. If the liquid layer and the CMT walls are hot, dere is limited condensation and the vapor collects at the top of the CMT. This breaks the natural circulation flow path, permitting the draining to begin. If the walls and liquid are cold, condensation occurs that can draw more two-phase mixture up the balance line to the CMT. For these conditions, water recirculation can be re-established and an intermittent recirculation can occur. 6.1.2 CMT Draining Phenomena CMT draining is the flow of liquid from the bottom of the CMT with a decrease in the water volume in the CMT. As the CMT begins to drain, it can uncover CMT metal surfaces, which may be cold or hot depending upon the length of the recirculation period. Depending upon the wall temperature, various amounts of steam are condensed. The thick CMT walls soon become conduction limited such that the condensation rate is determined by the CMT wall temperature distribution, surface area, l l q thickness, and degree of preheating due to the recirculation. ) b , 3681w-6.non:Ib-060897 6-3 I L

l FACILITY SCALING REPORT f Condensation can also occur on the CMT water interface as the tank drains. The amount of surface condensation is a function of the water temperature, the velocity of the steam as it impacts the surface, and the rate of condensate that flows from the walls to the liquid interface. The thickness of the hot liquid layer also influences the interfacial condensation since it tends to insulate the colder CMT water from the steam flow. The location of the liquid level, relative to the CMT inlet nozzle, also influences the amount of condensation since the steam jet from the nozzle may not be able to expand sufficiently and penetrates into the liquid layer, increasing the amount of mixing and generating surface waves. If the liquid is sub-cooled, the surface waves increase the interfacial heat transfer and condensation, which draw more steam flow into the CMT. As the CMTs continue to drain, all of the above effects, which enhance interfacial heat transfer, should diminish and a thicker, wann liquid layer should form both from the direct interfacial heat transfer and from the tank wall condensate. This layer of hot condensate can later flash during RCS depressurization, when the later stages of the ADS valves open and the CMT depressurizes. The flashing further enhances the mixing in this hot layer and can pressurize the top cf the CMT such that the delivery is increased. Once additional cold CMT wall area is exposed, the additional steam generated from the flashing condenses and is recycled back to the liquid layer. As the CMT drains there will be some liquid-to-liquid mixing; however, this effect is believed to be small since the system is in a thermally stable operating mode with the hot layer on the top. The axial temperature gradients in the ROSA-V test, given by Yonomoto et al. (1993) and Yonomoto and i Kukita (1993), indicate a relatively sharp interface between the heated layer and the cold CMT liquid.(2m As the CMT drains, there is also heat transfer from the hotter liquid layer to the CMT walls by convection. The amount of heat transfer that occurs depends upon the thickness of the hotter layer, its temperature, the draining rate, and the CMT wall initial temperature. If the convection from the hot liquid layer heats the CMT walls, the condensation heat transfer from the steam is reduced once the walls uncover. 6.2 Scaling Analysis for CMT Recirculation In the recirculation mode of operation, there is volume replacement of the cold borated CMT water with hot water that flows up the cold leg balance line to the top of the CMT. He CMT remains full during this time period. This process could continue until the CMT is fully heated and the natural circulation driving head for flow decreases to zero. l De ability of the OSU CMT to simulate the recirculation performance of the plant CMT is examined. Both top-down and bottom-up scaling analyses are perfonned to assess the capability of the test CMT to provih th key thermal-hydraulic data for the recirculation period. This section develops the governing 1 pations and examines the scalmg differences between the AP600 plant CMT and the OSU CMT to ensure that the key thermal-hydraulic phenomena presented in the PPIRT are captured in the CMT test. l ! 3681w.6.non:lb-060897 6-4 I

                                                                                    -          -           _ _ _      _  _ ______m

FACILITY SCALING REPORT I 6.2.1 CMT Recirculation Behavior - Top-Down Analysis j Natural circulation behavior has already been examined in Section 4.1 for loop recirculation in an .; integral system. The general governing equations derived for the top-down scaling of single-phase fluid recirculation are applicable to fluid recirculation in the CMT loop.

                                                                 ' Top-down analyses were performed at both the constituent level for the core and at the system level                             I for the loop. At the constituent level, the generalized system of equations that describe the natural circulation behavior through the core or any individual part of the recirculation loop is given by Equations 6-1 through 6-4 as follows:
                                                                   =    Mass (Pt V)"A      t     [ Pet)                                      @I)

• Momentum i , ,

                                                                                                                                                                '      4                   (6-2)

(Pf u,V,)=A [p ug,]$r8PtA D a , "'" g

                                                                   . Energy (p Cy ,(r)V,)=A g              [pfp,TS,]+H,A,(r,-T,)                     Boundary               N
                                                                  . Solid Energy Equation p,Cy ,V,           (T,) = H, A,(T,-T,)              Boundary 95 O
                                                            , . 3681w4.non:lb-060897 -                                                   6-5'

FACILITY SCALING REPORT The nomenclature for these equations is provided in Section 4.1.1. These equations can be normalized using their initial and boundary conditions to obtain the following set of equations:

• Mass

                                                                                                      *V ,1=A

• p,*Q,+ (6-5)

Tg.o Pa 1

• Momentum t,,, p, u, V,* I = A p,*u,*Q,* +11gi
• p,*AT (6-6)

                                                                                                                          <          s.

u it

                                                                                                       -I7p Pe t. .Q .

t

                                                                                                                                 +K.

2 , ,dh ,

• Energy 1

l Pe*C *, T,* V,* = A p,*Cp*,T,*Q,* t.o (6-7)

                                                                                                              +qtH,*(T, -T,)* g

• Solid Energy Fauation Ti.o T, = c,,, H,*(r, - T,)]3,,ng,ry + Q,* H l

1 where the characteristic time ratios are given by the following:

• Richardson Number hTo8dIo$

fliu = (6-9) 368Iw-6.mn:IW897 66

FACILITY SCALING REPORT

• Friction Number flF= +K (6-10) dh

                                                                        ,o e    Stanton Number "50 ^ 5 (6-11) 4 = pg Cpg Q,o

• Heat Source Number 9o5 (6-12) g, = p,Cy ,ugATo V, O

The liquid residence time is dcfined as follows: V t,,, = (6-13) t,0 The transport time constant ratio is defined as follows: g , Tt .o , H,oA , V ,,o Ts ,o Ps vsCYko s t As demonstrated in Table 4-24, the dominant phenomena are described by the Richardson number and the Friction number. This suggests that to preserve the recirculation flow behavior of the AP600 plant CMT in the OSU CMT, the following criteria should be satisfied: bi3" O w ..

                                                                  =1                              (6-15) 368Iw-6.non:Ib-060897                                     6-7

FACILITY SCALING REPORT and

                                                         %,m
                                                                 =1                                           (6-16) 4,p This is supponed by performing an order of magnitude analysis on the momentum balance equation for CMT loop recirculation.

Since the liquid draining rates are so small (typically 50 lbm/s or 0.8 ft.3/s) the liquid residence time given in Equation 6-13 becomes very large (2500 seconds) such that the time-dependent terms in the momentum equation can be neglected. The change in the momentum flux term, the first term on the right-hand side of Equation 6-6, is also small since the change in the water density is given by the ratio of the density of hot water to the initial density, The maximum change is 26 percent effect if hot water at the saturation temperature at 1100 psia and cold water at 120 degrees Fahrenheit are considered. Therefore, the momentum balance becomes a balance between the buoyant forces and the frictional

       ' forces:
                                                                  ..    .'    r         s.

P' ' (6-17) 4 p

• p,*AT * =Q s

                                                                    ~

s f h+K Equation 6-17, as defined, represents the momentum balance equation for a specific component within the CMT recirculation loop By integrating over the entire CMT recirculation loop, Equation 6-17 becomes the following:

s. r 32

                                                                   ' '                                       (6-18) k,CMT h PfAT T      * = I},CMT                              +K 6

2 , i. dh ai,

                                                                                           ,i s l

l l l j 3681w-6.non:Ib-060897 6-8 l

FACILITY SCALING REPORT 6 4 y j where: 0To g AT,L th,CMT

                                                                                       %),CMT*                                                                (0'I9}

2 fo l l l and l

                                                                                                                      ,2
                                                                                                          +K                                                  (6-20) 4.CMT =           d i =1
                                                                                                    'h       '8, '

a;' CMT,o where Lih,CMT is the thermal driving length for the CMT recirculation loop. In Equation 6-20, the mass balance equation has been simplified as follows:

   \                                                                                           pg uag o =pgua  n;                                            (6-21)

I where: 1 I ao= reference flow area for the CMT recirculation loop ai = flow area of the i* component within the CMT recirculation loop 6.2.2 CMT Recirculation Behavior- Bottom Up Scaling Analysis The bottom-up scaling analysis calculates the recirculation performance of both the AP600 CMT and the APEX CMT to verify that similarity of performance exists between the two CMTs. The assumptions used in this analysis are the following:

1. Single-phase recirculation.

l 2. One-dimensional flow. 3.' Linear interpolation of the density gradient in the CMT.

4. Essentially zero velocity in the reactor vessel downcomer annulus.

5. Loss coefficients are independent of Reynolds number.

L O- j 6. Quasi-steady state conditions. l l. 3681w-6.non:Ib 060897 - 6-9

FACILITY SCALING REPORT The genermed mechanical energy balance equation is given for the generalized system equation: 2 d+U p b 2gc

                                                                  . 3g           u 2gc 2
                                                                                         + .8._

ge L+W =0 5 (&22)

                                                                     ,d h       ,

where W, is the rate of shaft work and is set to zero. Since the starting point and end point of the calculations are the same, then:

                                                                              $ =0                                              (6-23)

P Expanding Equation 6-22 around the piping network using the component density gives the following: 2 2 2 ' 2 P2ugy, - p3uct fLf' u CL fLr ' u ggt

                                                        +                                +    K T +K                 pg Pt 2gc       2ge      ,

h,CLa

                                                                    +h d ,cL        Ec      s cm a+ d h,BAL         Sc (6-24) 2
                                         +

ECMTgx +KCKv +KDVI, + P2

                                                                                                +

PIL- P2L=0 h ,pyj 2Sc Sc 8e where: p3 = hot RCS and balance line flow p2 = cold CMT fluid density The form loss coefficients and velocities are defined in the nomenclature. By applying the Boussinesq approximation, all of the fluid densities in the CMT loop are assumed equal- except those which comprise the buoyancy term. Thus, the mass conservation equation becomes as follows: l UCLaCL = uBALaBAL = UDVIaDVI " Of ) O 3681w-6.non:IM)60897 6 10

FACILITY SCALING REPORT A Q Transposing the gravity head term to the left side of the equation and using Equation 6-25, the system

         . equation (Equation 6-24) becomes the following:

2 g ' K + -1 3 _ge- L(P2 ~ PI) = 2gc ct N dh,CL a 2

                                                                      ,s                                                                        CL f                         %

fL t P'

                                                    +

K7+K 2 (6-26) CMTu +h T ,BAL a BAL , l-t fL t P2

• I CMTgx + KCKY +KDV1 3

                                                                                   +7h ,Dy1                                     a 2

DVI i As the CMT recirculates, the top of the CMT fills with the less dense, hot fluid from the cold leg at [d~~~% density pg, Recirculation ends when all of the denser liquid (p2) has been replaced with the less dense pg fluid. To represent the decrease in the driving head, the buoyant term becomes the )

followmg

I I

1 i 8 {(L-L 3)p2 +L3 pi -pi L (6-27) bC which simplifies to the following:

p-L ) (p2-PI), (6-28) where

Li = height of less dense liquid pii n the CMT L = overall height r% 1  : V His expression has the correct limits since at the beginning of recirculation, L3 =0 and the full driving head is available. While L =L, 3 there is no driving head and recirculation ends. 3681w-6.nortIb-071897 6-11

i FACILITY SCALING REPORT j j The value of L ican be calculated as follows: l l s L, .j rh(t)dt P1ac Mr (6-29) o where: I acur = a function of L for i the CMT heads th(t) = mass flow into the CMT Equation 6-29 must be solved in an iterative fashion for the CMT head volume but can be solved directly for the cylindrical portion. The hydraulic resistances for APEX were measured in the pre-operational tests and were found to be j very close to standard textbook values. The current design values for the AP600 CMT balance line and direct vessel injection (DVI) line resistances were used, as well as an estimate for the reverse flow cold leg nonle loss and the DVI nonle expansion. Equations 6-26,6-28, and 6-29 were programmed for the AP600 and the OSU CMTs. Figure 6-6 shows the mass flow in the cold leg balance line as a function of time for the AP600 CMT balance line circuit. Figure 6-7 shows the development of the hotter, less dense fluid level in the AP600 CMT as it circulates. Figure 6-8 shows the mass flow in cold leg balance line as a function of time for the OSU CMT test circuit. Figure 6-9 shows the development of the hotter liquid layer in the OSU CMT test vessel. The analyses for the mass flows and the hot layer in the CMT show the same trends for both the AP600 CMT and the OSU test CMT. To compare the two calculations, the calculated values for the OSU CMT mass flow were multiplied by the OSU to AP600 mass flow scale factor of 96, and the

 ~ OSU time os multiplied by a factor of 2 to account for the time scaling of the OSU test. The resulting time scale and mass flow scale shifted OSU results were then divided by the calculated AP600 plant results.

Figure 6-10 shows the ratio of the rescaled OSU CMT mass flows to the calculated AP600 plant mass flows. As the figure indicates, the ratio is approximately 0.8 and eventually increases to unity. This indicates that the OSU CMT recirculation behavior, resistances, and gravity driving heads produce approximately the same recirculation flow behavior as the AP600 plant. A similar comparison has been made on the development of the hot liquid layer in the CMT for both calculations. The calculated OSU hot liquid layer values were multiplied by the height scale factor of 4, and the time scale was shifted by a factor of 2 to relate the calculation to the AP600 plant. Figure 6-11 shows the 368Iw-6.non:1b-060897 6-12

FACILITY SCALING REPORT ratio of the OSU CMT hot liquid layer to the AP600 calculated hot liquid layer as a function of time. The agreement between the two calculations is not as good with a ratio of approximately 0.65 to 0.7. i This indicates that the hot liquid layer develops at'approximately about the same rate in the test as in the plant. His comparison confirms that the scaling approach used for the OSU facility will yield the same CMT recirculation behavior as calculated for the AP600. j 6.3 Scaling Analysis for CMT Draining l

                                                       ' CMT draining occurs when the natural circulation flow path, established during a recirculation period, is broken and steam flows from the cold leg balance line to the CMT. The steam condenses on the cold CMT walls and water surface, drawing more steam into the CMT. As condensation proceeds, the
                                                       . wall and water surfaces are heated and the tank drains. Eventually, the draining process approaches a volume replacement process and the tank drain rate is based on the elevation head within the tank.

The phenomena of interest during the CMT draining process are as follows: 1

• CMT wall and interfacial condensation
• Transient conduction in the CMT walls Figure 6-12 is a flow diagram for the CMT draining scaling analysis. ' First, a top-down scaling _

y analysis is performed. The objective of the top-down scaling analysis is to obtain the transport time scaling constants and the characteristic time ratios that govern the CMT draining process. Following the top-down scaling analysis, a bottom-up scaling analysis is performed to scale specific processes including condensation heat transfer and conduction heat transfer. ' The parameters that have been scaled to address the phenomena of interest to CMT draining are as follows:

• Metal mass l

                                                              . Internal volume e' Length e   Diameter
                                                              . Heat transfer time constant
                                                              . Inlet nozzle diameter
                                                              =   Fluid initial conditions e   Rehtive elevation 6.3.1 Top-Down Subsystem Level Scaling Analysis for CMT Draining O

The coupled behavior of the CMTs makes scalin'g of CMT draining a challenging process, particularly for a test facility that will operate at pressures significantly below 2,000 psia. His section develops the similarity criteria needed to scale the various processes occurring in the CMT during draining. 3681w4non:lt460897 6-13

I FACILITY SCALING REPORT I Two CMT draining cases are examined in this section. The first case is CMT draining following a prolonged period of balance line recirculation. As the system depressurizes, the CMT fluid becomes two-phase. Because the CMT walls and liquid interface are already preheated, steam condensation is negligible. A two-phase depressurization rate equation, as developed in Section 5.0, will be implemented. The second case is CMT draining following a short period of recirculation. As the CMT drains, steam entering through the balance line condenses on the CMT walls because the walls are cold. A single-phase vapor depressurization rate equation will be implemented for this analysis. CMT Depressurization Rate Equation (Pre Heated Walls) Following a prolonged period of CMT recirculation, saturated liquid occupies the CMT. As the system depressurizes, the CMT fluid expands as it becomes two-phase. The continuity equation for a two-phase fluid mixture in the CMT is given by the following: l AL _ b CMT

                                                               ,                                       (6-30) t p.I'P(*TP)"y CMT              CMT O

where s tag is the stea.n mass flow rate entering the CMT through the cold leg pressure balance line. The flow leaving the CMT is given by ACMT, which is the CMT draining mass flow rate. For the case of pre-heated CMT walls, the steam mass flow rate (sa gt) is negligible because condensation is not taking place. The CMT draining rate is governed by the CMT fluid elevation head. Substituting Equation 5-14 into Equation 6-30 and neglecting s34t yields the following: l 1 th CMT p, (6 31) P_dtdP , _c pTPCI 8 V CMT o TP This equation can be made non-dimensional by dividing each term by its initial condition. Substituting Equation 5-15 and rearranging yields the following: dP* T CMT (P *)e -1 =- U'CMT f, I*g+$*MT. (6-32) O' 368Iw-6.non:Ib-060897 6-14

FACILITY SCALING REPORT l where the CMT fluid residence time constant is given by the following: CMT l tcMT = (6-33) kMT o .r and the CMT vapor generation rate ratio is given by: FV TY,cMT* 8 cMT (6-34)

                                                                                                                                            *CMT g o                                               j For the case where the CMT walls are superheated relative to the saturated liquid, the vapor generation                                                ,

rate is given by the following: - 1 pg, HwtAwnf,-TSAT) (6-35)

      -;                                                                                                                                hg,V CMT.         ..

where: i

                                                                                             =             heat transfer coefficient for heat transfer from the superheated wall to the Hwt saturated liquid (that is, vaporization heat transfer coefficient)

A,yt = total CMT wall surface area exposed to saturated liquid (T, - TSAT) = wall superheat Substituting Equation 6-35 into Equation 6-34 yields the followly HwtAwtf,-TSAT) (6-36)

                                                                                                                     . k CMT ,

rh cMT h g y) ( The CMT vapor generation rate characteristic time ratio is further evaluated in the bottom-up scaling

                                         . analysis.
                                          . 3681w-6.non:1b-071897                                                                       6-15

FACILITY SCALING REPORT CMT Depressurization Rate Equation (Cold Wall) l' For CMT draining which occurs following a short recirculation period, the CMT inside surfaces would still be at ambient temperature. Steam would be drawn through the balance line at a rate given by thBAL, which is equal to the steam condensation rate on the cold surfaces given by F, VcMT-

 'Iherefore, the continuity equation for the steam at the top of the CMT can be written as follows:

d (p gV g) = rhggt - F,VCMT 4 (6-37) The continuity equation for the liquid inside the CMT can be written as follows: d (6-38) Pt 3 (V,) = F,VCMT - rhey7 where the liquid density remains constant throughout the draining process. The total CMT volume is given by the following: VCMT = V s +V t (6-39) The time derivative of which is the following: d V, dV 8 _ (6-40) dt dt Substituting Equation 6-40 into Equation 6-38 yields the following: dV T,

• I,VcMT - thcMT ( I) 1 0

3681w 6.non:ib-oeos97 6-16

        ' . F FACILITY SCALLNG REPORT Expanding Equation 6-37 and rearranging yields the following:

dV 8 V

                                                                 =                           8. dp8 (6-42) dt                             pg    dt
              ' Substituting Equation 6-42 into Equation 6-41 yields the following:

Vgp, d p "IV f cMT 4 cMT (6-43) p, dt i I For isentropic expansion of the vapor during depressurization, the vapor density can be expressed in terms of pressure as follows: Pg

• Pgoc (P +MT}t4 (6-44) 3 l 1

1'

              - where PC *MT is the pressure in the CMT divided by the initial CMT pressure (PcMT,o) .The derivative of Equation 6-44 with respect to the dimensionless pressure is the following:

l d g 1-1 p'c dPCMT

                                                                .h(i                                Y                                                            (6-45) i Substituting Equations 6-44 and 6-45 into Equation 6-43 yields the fol!owing:

c gfa ,pycMT-dkMT (6-46) yPcMT I i r U ,. 36stw4.non:Ib 060897. 6-17

l l FACILITY SCALING REPORT j

                     'I7nis equation can be made non-dimensional by dividing each term by its initial condition. Further, dividing by ScMT,o yields the final form of the dimensionless CMT pressure rate equation:

T.CMT g V, dPC *MT , kond I+-m.fCMT

                                                                                                                                          +                     p y                dt p C+MT where the residence time constant (T   s .cMT) is the following:

8 (648) i,cMT* g m. C M T,o By allowing V togo be the volume of liquid that must be drained before depressurization begins (that is, ADS 1 actuates), then t g.cMT Physically represents the time to ADS I actuation subsequent to the end of CMT recirculation. The net condensation rate ratio is given by the following: IVt CMT (6-49) O k ond , b CMT o To evaluate 0,, it is necessary to perform an energy balance on the steam cor.*. acting the cold CMT surfaces. Figure 6-13 depicts an idealized view of the condensation processes within the CMT. It is assumed that steam enters the CMT at saturated conditions. The steam contacts the cold CMT walls where it condenses to form a liquid film. The liquid film flows down the wall and eventually mixes with the cold liquid. The steam also contacts the cool liquid surface. Depending on the momentum l flux of the steam jet entering the top of the CMT, the liquid interface may be quiescent or wavy to the point of sloshing. The interface type impacts the reent of condensation on the liquid surface. Steam l condensation and condensate film draining quicidy causes a thermal layer to form at the steam-liquid interface. Because of this, steam condensation at the liquid interface can be neglected. Writing an energy balance for the steam in terms of the wall film condensation mechanism yields the following: 5 (6-50) r,VcMT " h tg O 36stw.6.non:ltm897 6.] g

7 FACILITY SCALLNG REPORT where it is assumed that the steam gives up its latent heat (h g ). In this equation, q,, is the energy transferred to the condensate film. The. amount of energy transferred to the liquid film is given by the following: q,3 =Htp Awsfd - T,) (6-51) where: l

        . Hop          .=          heat transfer coefficient for condensation of steam on vertical surfaces. Obtained in the bottom-up scaling analysis
                             =     total CMT wall surface area exposed to the steam Aws
        - (Td - Tws) =             temperature difference across the condensate film as depicted in Figure 6-13 l'         It'has been assumed that conduction through the thermal liquid layer is insignificant relative to film condensation on the CMT walls.

Substituting Equations 6-50 and 6-51 into Equation 6-49 yields the following: Htp A ,(I'd -Tw) q , (6-52) hgrhcMT- .o Because the CMT pressure rate equation has been normalized using the CMT draining rate (thcMT).

        ' Equation 6-47 suggests that the CMT pressure behavior is dominated by the CMT draining rate (that is, the elevation head of the liquid in the CMT) when:

kond<1 (6-53) Conversely, the CMT pressure behavior is dominated by the condensation processes within the CMT when: I Rond>1 (6-54) y>s -6.nonitb4)60897 6-19

FACILITY SCALING REPORT This is particularly important to understanding CMT refill at low system pressure. Equation 6-52 is examined further in the bottom-up scaling analysis. CMT Structural Energy Equation Because the CMT walls are thick structures that are :nitially at room temperature, a more detailed solid structure energy equation, which assumes a cylindrical one-dimensional heat transfer in the cylindrical portion of the CMT, is written as follows: BT, k, a BT s (6-55) P:Cy , = r Br , with the following inner boundary condition: k,

                                                                                                                                           =HLF fd -T,)                                                 NO Br , g' 9

and with the following outer boundary condition:

                                                                                                                                                   =0                                                   (6-57) k:    Br   ,,g, This is based on the assumption that the heat transfer from the outer surface of the tank is sufficiently small such that the tank is adiabatic. For the heat conduction in the head of the CMT, a spherical coordinate system may be used as follows:

BT, ks 3 BT s 2 (6-58) p,Cvs 8t , "7E ,r The boundary conditions are stated above. O 368Iw-6.non;1M)60897 6 20

FACILITY SCALING REPORT i Equations 6-55 through 6-58 can also be expressed in non-diinensional form by making the following y definitions: l l I T, l- T,. = (6-59) . l hd -T,)o I hd ~Iw ) = fd -T,) (6-60) Fa -T.).  ;

• r  !

r = (6-61) hoS)i l Substituting Equations 6-59 through 6-61 into Equations 6-55 through 6-58 and rearranging yields the g following: L

• CMT Cylinder BT,* 1 3 BT,* (6-62)

Ts .cMT 8t " ~; &+ r {

• CMT Spherical Head I

l- , , BT,* 3 BT,* (6-63) I Ts cMT BT 1 g & + (ri I l 3681w-6.non:lb-060897 6-21

FACILITY SCALING REPORT

                        . Inner Boundary Condition BT*^                                                   (6-64)

• Si (Td-T,)*

Br . 7,g, a Outer Boundary Condition DT,' (

                                                                                  =0 r=RO where:

p,Cvs (Ro-R;)2 (6-66) Ts ,CMT , ks and N g, . LF hoS)i (6-67) k, Equation 6-66 is the conduction heat transport time constant for the solid which can be expressed as the Fourier number. Equation 6-67 is the Biot number. It can be written in terms of a heat I conduction specific frequency (cos .cMT) as follows:

                                                                  %i " S ,cMT s      s.cMT T                                    (6-68) l O

3681w 6.non:lt>M0897 6-22 L_____-__________________-____-_-_____-____________.---__-______

i FACILITY SCALING REPORT

  . ,m
    !                                            where:

(6-69) cos ,CMT = Ps C VS(Ro -R ) A characteristic time ratio for heat conduction in terms of the fluid residenceg time constant (T .C can be written as follows: bC

• U ,CMT s gI CMT )

That is: g= LFPfY g g escVS(Ro -R i) thCMT,o l l D \ ,) i Tables 6-1 and 6-2 summarize the non-dimensional equations, time constants, and characteristic time j j ratios for the CMT.  ! 6.3.2 Bottorn up Scaling Analysis for CMT Draining Processes l l . To evaluate the characteristic time ratios developed through the CMT top-down scaling analysis and

                                                                                                                                                                                )

summarized in Tables 6-1 and 6-2, constitutive equations are needed for the following: L- l Vaporization heat transfer coefficient (Hwt.) j Condensation heat transfer coefficient (Htp) l l v10 3681w4.non:Ib-060897 6-23

FACILITY SCALING REPORT Convective Heat Transfer on Superheated CMT Walls I-Subsequent to a long period of CMT recirculation, the CMT walls reach the saturation temperature at the recirculation pressure. When the system depressurizes, the saturated liquid in the CMT flash'es and the CMT serves as a heat source. For these conditions, Rohsenow's correlation for nucleate boiling is implemented.C3) That is: r *3 P"hig Ja 4.1 (6-72) H = Pr Afw-TsAT), k where C,7 equals 0.013 for stainless steel in contact with water, and where: g, C (6-73) g(PtsPss). The Jakob number is defined as follows:

                                                                                                                                                                 , -TSAT)

Ja = PI (6-74) h is The Prandtl number is defined as follows: Pr, = PtPh (g,73) k,3 Substituting Equation 6-72 into Equation 6-36 yields the final form of the CMT vapor generation rate characteristic time ratio: r '3 4.1 Pts Awt Ja Prf (6-76) k,CMT* A --C,g, thCMT,o 368iw4.norcit> o60897 6-24

FACILITY SCALING REPORT Condensation on Cold CMT Walls Subsequent to short periods of CMT recirculation, the CMT walls will still be cold during the draining process. It is assumed that film condensation exists on the CMT walls. The condensation heat transfer coefficient is a function of the film Reynolds number, which is defined as follows: 4 *fm Refm = (6-77) Efm where:

• fm = film flow rate per unit width fm = film properties The wall condensation model used in the CMT scaling analysis is the Nusselt (1916)M) laminar film condensation coefficient for Ref s 2000; given as (see Kreith,1962):(5)

O pg , hiskfm 3

                                                                                                                                                 /

(6 78) Rtp = 0.94 Htm Ls fd -T,), where L, is the length exposed to the steam, and where: Tw+T SAT Td" (6-79) 2 such that: T d -T, = hSAT-I) w ( } where T, is the surface temperature, which is calculated from the one-dimensional CMT wall O conduction solution. For turbulent film condensation, the modified Colbum (1993) correlation is used, as recommended by Kreith (1962).(5) 3681w-6 non:Ib-060897 6 25

FACILITY SCALING REPORT The correlation is given as follows: u3 I" * (6-81) Hg = 0.056 Prf} 1

                                                       , pfm ,             p This is a local correlation since *fm is the local film flow. However, in the CMT heat transfer analysis, an average film flow is calculated separately for the dome region and for the combined dome and cylindrical wall region. Therefore, the *fm in Equation 6-81 represents the average film flow, such that the heat transfer coefficient is also taken as an average, over these regions.

6.3.3 Evaluation of CMT Dimensions ) This section implements the results of the top-down and bottom-up scaling analyses to determine the physical dimensions of the CMTs. CMT Diameter By fixing the time and length scaling ratios, it is possible to obtain a scaling ratio for the cross-sectional area of the cylindrical portion of the CMT. Equation 6-33 can be written as follows: PTPaCMTbCMT (6-82) tCMT,R . b CMT

                                                                                   .R Substituting Equation 7-3 into Equation 6-82 for a single CMT injection line yields the following:

TCMT,R = 96 (aCMTbCMT)R (' } Because all of the time ratios have been set to 0.5 and di of the length scale ratios have been set to 0.25, Equation 6-83 can be solved to obtain a scaling ratio for the CMT cross-sectional area: aCMT R " O 368 t *4.non:lt> 060897 6-26

FACILITY SCALING REPORT (3 Equation 6-84 can be written in terms m CMT diameter as follows: I a,c DCMT,R * } The inside diameter of the full scale CMT is 381 cm (150 in). Therefore, the ideal model CMT diameter would be 54.8 cm (21.6 in). CMT Volume The required CMT valume scaling ratio is 1:192. The volume of the full scale CMT is 56.7 m3 3 (2000 ft ). Therefore, the ideal volume of the model CMT would be 0.295 m3 (10.4 ft 3). l CMT Length The required CMT length scaling ratio is 1:4. The full-scale CMT has an internal length of 624.8 cm

       ~ (246 in). Therefore, the ideal internal length of the model CMT is 156.2 cm (61.5 in.).                                                                                                I
    ,s

( ( CMT Metal Mass If one recognizes that the principal objective of scaling CMT condensation behavior is assuring that CMT liquid inventory and draining rates are reasonably modeled by the test facility so that initial conditions for long-term cooling can be properly established, then the following condition must be satisfied: wcond)m , I (6-86) 192 [Mwcond), where M wcondi s th: total mass of condensate produced by steam condensation on the CMT internal metal surfaces during the period in which CMT wall heat-up occurs. The total CMT wall condensate is defined by the following: t l tr M dt (6-87) wcond "f wcond dl f *

  'w
       .368Iw 6.non:1b-060897                                                   6-27

FACILITY SCALING RF.' ORT where tf is the time required for the CMT wall to reach T347 Performing a global energy balance on the CMT yields the following: M,CvshSAT-I w)0

                                                              *WCODd (6-88) g is Substituting Equation 6-88 into 6-86 and solving for the metal mass ratio yields the following:
                                                                               %c h'8

[M,)"= (6-89) hSAT-Tw} . R where CMT structural material similarity has been assumed. The CMT metal mass ratio is 1:127.5. This result, in conjunction with Equation 6-71, is used to determine the CMT ./all thicknesses. Summary of Model CMT Dimensions Table 6-3 summarizes the dimensions of the model CMT based on the scaling analysis. The following sections examine scaling distortions which arise as a result of implementing the selected CMT design. 614 Evaluation of CMT Draining Processes and Scaling Distortions To assess the relative importance of the CMT draining processes, the characteristic time ratios presented in Tables 6-1 and 6-2 must be evaluated for both the APEX model and the full-scale AP600. A comparison of these values reveals which of the processes are cominant and to what extent the proposed APEX CMT geometry and operating conditions introduce scaling distonions. Evaluation of Convective Heat Transfer on Superheated CMT Walls Table 6-4 presents the numerical values for the CMT fluid residence time, the specific frequency, the characteristic time ratio, and the distortion factor for the CMT draining processes that occur when the walls are superheated relative to the two-phase fluid mixture in the CMT. This occurs following a prolonged period of CMT loop recirculation. For purposes of calculation, it is assumed that ADS actuation quickly depressurizes the AP600 and APEX from their respective reference pressures (that is, 1080 psia and 350 psia) to atmospheric pressure. This would be the maximum vapor generation case. The CMTs are assumed to be filled to 60 percent of their total volume which would approximate the case shonly after ADS 1 operation. I l 3681w-6.non:Ib-o72297 6-28  : I \ l l t.-.._.._...

l FACILITY SCALLNG REPORT

                 )   For these operating conditions and CMT geometry, the residence time of the fluid in the APEX CMT l                      is approximately 325 seconds as compared to 650 seconds for the AP600. Thus, the one-half time scale factor is preserved.

I Table 64 indicates that for these conditions, the vapor generation rate characteristic time ratio (D r.cMT) is significantly laig $.n one, thus, indicating that the vaporization process is important. l The distortion factor for U r,cMT si found to be 17.4 percent. This factor could be reduced to 2 percent,if the APEX reference pressurc were at 400 psia. Table 6-4 indicates that the APEX CMTs adequately simulate the draining processes subsequent to [ prolonged CMT loop recirculation. l Evaluation of Steam Condensation on Cold CMT Walls ) Table 6-5 presents the numerical values for the CMT transport constants, specific frequencies, characteristic time ratios, and distortion factors for the CMT draining processes that occur when the l walls are cold relative to the steam entering the CMTs through their balance lines. This occurs when the CMT loop recirculation period is short. For purposes cf calculation, it is assumed that the AP600 and APEX are at their respective reference pressures. This calculation examines the effect of condensation during the initial draining of the CNT to the l sensing level for ADS I actuation. That is, the fluid residence time represents the time to ADS 1 l' actuation subsequent to the start of draining. The CMTs are assumed to be 6tled to 60 percent of their total volume. For these operating conditions and CMT geometry, the time to ADS 1 actuation is approximsely 104 seconds as compared to 208 seconds for the AP600. Thus, the one-half time scale factor is preserved in APEX. Table 6-5 indicates that the characteristic time ratio for condensation is preserved in the APEX CMTs as indicated by its distortion factor, which is less than 3 percent. This indicates that the condensation process is reasonably modeled in the APEX CMT. Table 6-5 presents the characteristic time ratios for heat conduction (that is, Biot number) in the CMT head and cylindrical shell based on the solid transport time. The numerical values are relatively large, j which indicates that a lumped parameter conduction model (H ai ess l than 0.1) would not be applicable. In order to assess the importance of the conduction process on CMT draining, the characteristic time ratio for conduction needs to be evaluated on the fluid residence time scale as given by Equation 6-71 for DHC. The numerical values for HHC show good agreement for the APEX and AP600 CMT heads as indicated by its distortion factor of 6.3 percent. However, the numerical values of UHC for the (q/ cylindrical portion of the CMT indicates that conduction in the cylinder is not simulated as well in the APEX. 3681w-6.non:Ib.060897 6-29 l e

l FACILITY SCALING REPORT l l 1 Table 6-5 indicates that condensation and conduction in the CMT head region are reasonably simulated. It also indicates that additional analysis is warranted for the cylindrical portion of the CMTs. This analysis is presented in the next section. Calculation of CMT Wall Heatup Due to Steam Condensation The conduction equations coupled with their respective initial and boundary conditions as specified in Equations 6-55 through 6-58 have been solved numerically for the AP600 and APEX CMTs. Figures 6-14 through 6-23 present the intemal wall surface temperatures as a function of time for different fluid volumes and pressures in the AP600 and APEX CMTs. These figures show the same trends for both the APEX and AP600: a rapid initial dse in wall surface temperature approaching the

<aturation temperature. Comparing the AP600 CMT wall heat-up rate with the APEX CMT wall heat-up rate at the scaled pressure indicates that the initial temperature rise in the wall surface temperature does not necessarily occur on the half-time scale basis. (Note that an APEX reference pressure of 385 psia was used for calculational purposes.) However, the long-term approach to the saturation temperature is correct because the metal mass as a function of CMT length is properly scaled.

Therefore, Equation 6-89 is satisfied for each section of the CMT as shown in Table 6-3. 6.4 Conclusions This section has implemented the Hierarchical, Two-Tiered Scaling (H2TS) methodology to obtain the CMT scaling criteria for CMT loop recirculatica and draining. The important CMT processes have been identified and scaling distortions evaluated. For CMT loop recirculation, the dominant transport processes are charactenzed by the Richardson number and the Friction number. An assessment of the scaling distortions indicates that the dominant processes for CMT loop recirculation can be adequately simulated in the APEX test facility. For CMT draining, two cases have been examined: CMT draining in the presence of hot CMT walls and CMT draining la the presence of cold CMT walls. For draining while the CMT walls are hot, the CMT wall vapor generation number (DrcMT) represents the dominant process. An assessment of the scaling distortion, presented in Table 6-4, indicates that this process is reasonably simulated in the APEX test facility. For CMT draining while the CMT walls are cold, the condensation number (Ucond) and the head and cylinder conduction numbers (UHC. Head and UHC cyl) represent the dominant processes. An assessment of the scaling distortions indicates that the condensation and conduction processes in the head will be reasonably simulated in the APEX test facility. The cylinder conduction number (UHC.cyl) is not preserved in the APEX CMTs. Additional analysis indicates that the long-term behavior is reasonably siinulated in APEX because the CMT mass is properly scaled. In conclusion, the scaling analysis presented in this section indicates that CMT loop recirculation and draining data of sufficient quality to validate computer codes can be obtained with the APEX test facility. 3681w-6.non:lb-060897 6-30

FACILITY SCALING REPORT l-6.5 References 1

1. AP600 Standard Safety Analysis Report (SSAR).

l 2. Yonomoto, T., Kukita, Y., and Anoda, Y, " Passive Injection Experiment at the ROSA-V Large Scale Test Facility," 1993 National Heat Transfer Conference, Atlanta, GA, August 8-11, 1993. l_ 3. Yonomoto, T., and Kukita, Y., "RELAPS Analysis of a Gravity-Driven Injection Experiment at l ROSA-V Large Scale Test Facility," ASME Winter Annual Meeting, paper 93-WA/HT-74, New Orleans, LA, Nov. 28 - Dec. 3,1993.

4. Nusselt, W., " Die Oberfluchenkcondensatiin des Wasserdampfes," Z. Ver. Deutsch, Ing. Vole 80, l pg 541,569,1916.

5. Kreith, F., Principles of Heat Transfer, International Textbook Company, Scranton, PA.

l l l l 1

  '3
  /

V l l-i

  .A 1

f 36stw-6.non:tbes97 6-31 Q-_____-_______-_____-____.-__..--_ __ .- -

l FACILITY SCALING RE*'3RT TABLE 6-1 TOP DOWN SUBSYSTEat LEVEL SCALING ANALYSIS DI51ENSIONLESS EQUATIONS FOR l TITE CMT DRAINING PROCESSES (PRE IIEATED WALLS) Depressurization Rate Equation:

                                                                                                            -1 dP *                          +                           +

TCMT (P po ,_ (6-32) Fluid Residence Time: PTPU TPVcMT (6-33) 7CMT ,

                                                                                                                         $ CMT          ,o Characterisuc Time Ratio:

HwtAwn(T,-TSAT) (6 36) b CMT Ilfg ,o O 3681w-6.non:iw)60897 6-32

FACILITY SCALING REPORT o l TABLE 6 2 l . TOP DOWN SUBSYSTEM LEVEL SCALING ANALYSIS DIMENSIONLESS EQUATIONS FOR CMT DRAINING PROCESSES (COLD WALLS) Depressurization Rate Equation: T.CMT g V*g dP'MT C .+ g) odI+-S i y dt t CMT i p C+MT Structural Energy Equation: CMT Cylinder i BT,' i a . BT,* (6-62) i Ts,CMT * - T l 6 g . CMT Scherical Head f  %

BT,* 1 3

• BT,* (6-63)

Ts,CMT BT (f )

        ,                                                                                           (r *)2 Or * ,          Or

• Inner Boundary Condition l

BT'* (6-64)

                                                                                              .      " bi fd -T,)*

I r4g Outer Boundary Condition BT'* (6-65)

                                                                                                              =0 r40 Transport Time Constants:

E 8 t.CMT " (6-48) g E.' CMT,o l Cvs (R,-R;)2

                                                                                         ,s.CMT , Ps                                              ( 66) k*

f. -

7 88 3681w-6.non:1b-072297 6-33

1 FACILITY SCALING REPORT l 1 l TABLE 6-2 (Continued) TOP DOWN SUBSYSTEM LEVEL SCALING ANALYSIS D151ENSIONLESS l EQUATIONS FOR CMT DRAINING PROCESSES (COLD WALLS) Fluid Characteristic Time Ratios: Hy t A,,(Td -T,)

                                                                             ,                                                                                                (6-52) hgthcMT       ,o SIC "                                                                                                                            (   I)

PsCvs(R,-R i) rincMT o Solid Characteristic Time Ratio: g; = 15 ' ' i) (667) O O 368Iw.6.non:1b-072297 6 34

FACILITY SCALING REPOS T I \ V TABLE 6-3 MODEL CMT SCALING RATIOS AND DIMENSIONS g OSU APEX Parameter Ratio AP600 Ideal Actual Head Inside Radius 1:6.928 75.0 in 10.825 in 10.50 in 190.5 cm 27.50 cm 26.67 cm Cylinder inside Radius 1:6.928 75.0 in 10.825 in 10.782 in 190.5 cm 27.50 cm 27.39 cm Internal Length 1:4 246.0 in - 61.5 in 57.2in 624.8 cm 156.2 cm 145.29 cm Internal Volume 1:192 2004.40 ft3 10.44 ft 3 10.44 ft 3 I 56.75 m3 0.295 m 3 0.295 ft3 l Metal Volume 1:127.5 420.55 ft 3 3.30 ft 3 3.20 ft 3 3 16.65 ft 0.131 ft 3 0.091 m 3 Head Wall Thickness -- 4.6 in --- 1.5in

        .p                                                                                                           11.68 cm                          3.81 cm t               \
         'V                                                 Cylinder Wall Thiciaess                    --

8.0 in --- 1.218 in  ! 20.32 cm 3.094 cm I Length to Diameter Ratio (L/D) 1:1 1.0 1.0 1.73 l Numerical values may not reflect as-built conditions. Refer to the APEX Facility Description Report. i I i

         -f%

3681w4non:Ib-072297 6-35

FACILITY SCALING REPORT TABLE 6-4 EVALUATION OF Ch1T DRAINING FOLLOWING PROLONGED CSIT LOOP CIRCULATION (IIOT CMT WALLS). RESIDENCE TIMES, CIIARACTERISTIC TIME RATIOS, AND SPECIFIC FREQUENCIES APEX AP600 CMT Pressure 14.7 psia 14.7 psis CMT Fluid Residence Time: tcMT 325.6 s 651.1 s CMT Process Specific Freauency: Gr.cMT 0.208 s'l 0.126 s*l Characteristic Time Ratio: fir,cMT ti7.8 82.1 Distortion Factor (DF): Wall Vapor Generation Number (FIr.cMT) 17 4 % O1 O 368Iw4.non:t b-072297 6-36

FACILITY SCALLNG REPORT O (] TABLE 6-5 EVALUATION OF CMT DRAINING WITH COLD WALLS. RESIDENCE TIMES, I CHARACTERISTIC TIME RATIOS AND SPECIFIC FREQUENCIES l APEX AP600 CMT Pressure: 350. psia 1080 psia CMT Transport Process Times: ADS 1 Actuation Time: T.cMT g 104.2 s 208.2 s l CMT Head Conduction Time: l Ts. Head 359.8 s 3383.3 s CMT Cylinder Conduction Time: T3.cyl 237.2 s 10233.2 s l CMT Process Snecific Frequencies: mcond 0.0589 s~l 0.0303 s'3 m,,ges 0.467 s'l 0.249 s'3 m,,cyi 0.575 s'I 0.143 s*3 Characteristic Time Ratios: Ucond 6.14 6.32 Usi,geg 167.9 843.6 Hai,eyi 136.3 1467.2 1 U HC. Head 48.6 51.9 D HC.cyl 59.9 29.9  ; I Distortion Factors: i (Ucond) 2.8% (D Hc. Head) 6.3% l (UHC.cyl) -100.6 % i l I I I l rn D l l

                          %8Iw.6.non:Ib-072297                                                                 6 37

l. FACILITY SCALING REPORT I O

             .          .9  _9 '"
          'YO Y         ,9,_ ** y, 9_ *' M.

M

                                                                  -e                        conc m e amn                      TANK (I or 2) pun  _t_

eness a mt TW>< M

                %)

51. ~  %) ACCW.

Ior2) es 57/> 4; kl '"

                                         &L A

M

                                                                            ~~

I g ~i ' l' i) (A T ._ " b a a l I r .*

                          %.                                    7,'";7         I   ._v;A coac atAcron wsm O

Figure 6-1 AP600 Passive Safety System Design 36sIw-6.nen: u)72297 6-38

( 1 FACILITY SCALING REPORT f l O l l l c' to N a

                                                                                                                                                   /     ----

7.78 in. %  %

                                                                                                           .d                     ~

d l a o W N $

                                                                                                                                #          150.0 in.                       :

u V 4.6 in. N r ,  % AP600 CMT hemispherical heads (top and bottom)

                                                                                                                                             /////////

O . Figure 6 2 AP600 Core Makeup Tank 368:=4mn:Ib-072297 6 39

FACILITY SCALING REPORT i O O n

                                                                                                                                                                 &                                    CORE MAKEUP r,

TAMK (10F 2) a'a LEG () - PBL pgog IRWST k ] ACCUMULATOR (1 OF 2) HOT COLD 4 LEG , ,- LEG DVI (1 OF 2) a cone M E'I O l Figurc 6-3 AP600 CMT Piping Layout 3681w-6.non:Ib 072297 6 40 t L

FACILrn' SCALING REPORT O l l l l l

             .-        Recirculation !

4-a- -+,I

• w l g-k .

v l l

              -                                                                                                                                                Y.

i' l i l g l le .. , _ wue wun--- I Figure 6 4 AP600 SSAR Calculation of CMT Draining Flow for 2. Inch Cold Leg Break 3681w 6.non:Ib-072297 g,4 g

FACILITY SCALING REPORT O APG00 2 INCH CL TRANSIENT 9 - 59 VFMFN 65(1)TP1 1.

                                                        .9-L .8--

I

                                                   *    .7 ~

u. S '8-li!

                                                       .s--

h, s

                                                       .4-d h.5-9
                                                   ?

a. .2--

J

                                                       .1-1500        2000       2500       5000
                                                         *0        500       19'08 TIME (SEC)

O Figure 6-5 Cold Leg Balance Line Void Fraction for 2-Inch Cold Leg Break

                                           ' M81w-6.nonsom97                             6-42

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                                                                                                            ' 0 F                                                                                                       0 2

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         =                                                                                                         )

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• e* >2g >[ 3o4 e F oo5 rA =$Ra CE

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      =        C P         t n                                                                                                     0
       ,   . a                                                                                            '

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            =                                                                                1 d

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                                                                         ;i l'     Ijl!l              ilIi ! l   1 m>Or       5>e2C             3Wd 4l 0

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                             -      -     -      -   -        -       -         -               5 2

r e i L s p m 0 r 0 0 . a t 0 5 W 2 3 T M

       =      C P       U S
        ,     O F

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       =                                                                                          )

c t e o s h ( T e F m 0 0 1 . t 0 0 i 0 T O

       =                                                                                        1 d

l o c T U S O 0

            .                                                                                e  0 r                                                                                         5 o

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0 0

                             -         -         -        -           0 3

_ u l - F s 0 i s 0 a I 5 2 M _ 0 0 6 _ P A _ g n / 0 i . l e U f 0 0 a S 2 c O S _ n o ) i _ t a l 0 c e (s - ui 0 c f 5 r 1 e i c m i e _O R T l a r u 0 t a I 0 a 0 _ N 1 0 _ I 0 5 0 1 8 6 4 2 a 0 0 0 0 o se i O

                   ,*~ os -@ *g FF5 2 f % omC o>1 moE w. so4 5i& 8=

f$ .L<

l l j

                                                                  ,>O"dmOe2o:t8a     =n 0

0 0 O 3 s s e n k 0 i c ' 0 i 5 h 2 T r e y a L g n 0 0 0 0 i e ' l 6 0 a P 2 c A S / n U o S

                                                                                  )

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                                          >$ 8       :::% $ E r d 4 H:r b :s g 57Mk{b"                                 Ts*                                                _

FACILITY SCALING REPORT

 .I CMT CONDENS ATION AND DRAINING PHENOMENA Top Down/ Subsystem                              Bottom-Uphrocess Scaling                                         Scaling

• CMT Draining Rates -

CMT Condensation

                                                                                . CMT Depressurization

• CMT Heat Conduction Rate Chapter 5 CMT Balance Line CMT Condensation and and injection Line Depressurization

                                                                                                --.       Draining II Groups and      -

Scaling Scaling Analysis Similarity Criteria IP Evaluate Scaling Distortions l Design Specifications , ( Figure 612 Scaling Analysis Flow Diagram for CMT Condensation and Draining Processes 3681w-6.non:lt472297 6-49

l FACILITY SCALING REPORT l O eb4 8\ ANNNNuxxw J. N }5 b L S' 1

                                               /

a m \-

                                                   <-6 m
                        +>

d V i C O O O 5 E l . xwwwwwwwwwwwwwwx T T' , f I II I I t t y

                       -    c

• I t i 1 er l ->l"l 0 ',i, p E

I E

                                                                                             *E I

h *8 41l$ t t E E I I SS $ f  ; II II 'I t I w w w w w w w w w w w w s:.,,w w Figure 6-13 Idealized Model for CMT Cold Wall Condensation Behavior e 368Iw4non:1b472297 6-50

FACILITY SCALING REPORT O ANeo CMT WaB &at Up ante (prissure tese psia) 600 - N 500- _ q_ -I I-- 2 I -- % I E N S 4001 g 300 - 9 200 - 100 X 0 500 1000 1500 2000 2500 3000 3500 lime (s)

                                     %75           %60 -*- %50 + %40                                                                                                   -X- %25 Figure 6-14 AP600 CMT Wall Heat Up Rate for Different Fluid Volumes at 1080 psia p

AN00 CMT Wa5 Heat Up ante (pressured 80 pein) 600-

                                                        %_ -;                                            X                           X                       %    %    %    R-R-x 400-l 300-200 -

100X l 0 500 1000 1500 2000 2500 3000 3500 l Time (s)

                                     %75            %60 -*- %50 + %40                                                                                                    -X- %25
 '()
  /"%

Figure 6-15 AP600 CMT Wall Heat Up Rate for Different Fluid Volumes at 800 psia 368Iw-6.non:1b 072297 g,$ }

l FACILITY SCALING REPORT AN00 CMT Wd Heat Up Rete (pnseure=400 puis) 500-

                                                                                                                                                                                          ^
                                                                                                                                                                                                ;__     y. T    -R--R--R-R-R-R
                                                                                                                                                                            ^ - ^

400 - I 300 - E H 200 100 X 0 500 1000 1500 2000 2500 3000 3500 T1ame(s)

%75 %60 -X- %50 + %40 -x- %25 Figure 616 AP600 CMT Wall Heat Up Rate for Different Fluid Voluines at 400 psia O

AN00 CMT WaII Hast Up Rate (pnssun=200 pois) 400 2 -: 2 -R X R

                                                                                                                                                                                            ^
                                                                                                                                                                                                 -_g_..  -7_   7-.

350 - * ^~ g300; 5 C 250-8. E H 200 150 - 100 X 0 500 1000 1500 2000 2500 3000 3500 Thue (s)

%75 %60 -*- %50 ---O- %40 -X- %25 Figure 6-17 AP600 CMT Wall Heat Up Rate for Different Fluid Volumes at 200 psia 3681w4non:Ib-M2297 6 52

FACILITY SCALING REPORT I l l I AP600 CMT Wd Hast Up Rate 4, . . paes) 300 -

                                                                                                     %         %        %        %       R            X    R q_        q.        %-       %

250- ' b < 200 - I 150 - 100 X 0 500 1000 1500 2000 2500 3000 3500 1 11 mas 4)

%75 %60 -*- %50 + %40 -x- %25 i

i l Figure 6-18 AP600 CMT Wall Heat Up Rate for Different Fluid Volumes at 50 psia l bv Medal CMT Wd Heat Up Rate OPressm=385 pois) 450-

                                                                                                       ^          ,__           _ P =%-                  %    %

400 - n=y-N ^~ 350< g 300 - l l

1 250 h 200 - )

i 150 -  ! 100 X i 0 50 100 150 200 250 300 350 j line 4)

%75 - %60 -*- %50 3-- %40 -X- %25 Figure 6-19 APEX CMT Wall Heat Up Rate for Different Fluid Volumes at 385 psia ,

l 3681w-6.non:lb472297 6 53

FACILITY SCALING REPORT l l Model Orr Wall Heat Up Rate (primuren285 paia) O' l 450-

                                                                                                                                                                          ,oo .                                                                                                            + x-x-x-x-x-x-x                                                                                                             )

350 - E 300 250- , 200-150 100 X 0 50 100 150 200 250 300 350 400 450 500 Timme4)

%75 %60 -*- %50 --0-- %40' -X- %25 Figure 6-20 APEX CMT Wall Heat Up Rate for Different Fluid Volumes at 285 psia O

Nedet Off was Heat Up Rate (pressan=1424 pele) 400 - , 350 - - _ g g-2-R-x-x-x-x-X g 300 - 250-200 - 150 - 100 X 0 50 100 150 200 250 300 350 400 450 500 Time 4)

                           /

% 75 %60 -*- %50 --0-- %40 -X- %25 Figure 6 21 APEX CMT Wall Heat Up Rate for Different Fluid Volumes at 142.6 psia 3681w-6.norrit>072297 6-54  :

FACILITY SCALING REPORT Ch '

   \

V} Mead cur was that up anse (pressar.-71.s pesa) e 350 - l soo- _ _ n

                                                                                      - -x       x    x-x-x-x-x-x G-250-l 200                                                                                                                                              l l                                                                                                                                                                                 l 150 100                                                                                                                                              l o           100        200          300            400        soo               soo                                      700         l time (s) l
                                                    %75               %60        I        % 50       0      %40     -X- %25 l

l Figure 6 22 APEX CMT Wall Heat Up Rate for Different Fluid Volumes at 71.3 psia ( ,v l' Medal O4T wad Heat Up Rate (pressmev=17.825 pela) 240-t l 210 - _ h2 X X-X X X X X X l 200 - G i 180 - . I 160-l 140 - L 120 - 1YU X 0 203 400 600 600 1000 1200 Det 4)

% 75 %80 -*- %50 + %40 -X-- %25 b\

Figure 6-23 APEX CMT Wall Heat Up Rate for Different Fluid Volumes at 17.8 psia 36sIw-6mibm2297 6-55

FACILITY SCALING REPORT i

                . 7.0 VENTING, DRAINING, AND INJECTION SCALING ANALYSIS This section presents scaling analyses for all of the important venting, draining, and injection processes expected to occur during the course of a loss-of-coolant accident (LOCA). A scaling analysis of l                     venting processes governs the sizing of the balance and automatic depressurization ' system (ADS) l                     lines.' A scaling analysis of draining processes govems the sizing of die accumulator (ACC) tanks, the

in-containment refueling water storage tank (IRWST), the upper support plate, and the downcomer bypass holes. A scaling analysis of injection processes govems the sizing of the accumulator,IRWST, l and direct vessel injection (DVI) lines. A scali.ig analysis of injection processes within the l downcomer governs the sizing of the DVI nozzle and the downcomer gap. Thus, the following l scaling analyses are presented h this section

i. I.

                         *-    Accumulator tank scaling

• IRWST scaling

                         *-    Injection line scaling

• Balance and vent line scalmg

                                                                                                                                                   ]

l L

• Upper support plate scaling

                         *   ' Downcomer scaling Lp                  Section 7.1 presents the depressurization scalinq requirements of Section 5.0 that are applicable.to V~                 venting, draining, and injection processes. The remaining sections rely on the criteria presented in                          ;

Section 7.1. - Section 7.2 presents a general scaling analysis for tank draining processes. It supports the accumulator scaling analysis presented in Section 7.3 and the IRWS f scaling analysis presented in Section 7.4.

                                                                                                                                                   )

The safety injection line analysis (given in Section 7.5) consists of a bottom-up scaling analysis, which

                 'is tied to the top-down scaling analysis of safety tank draining. Section 7.6 presents a bottom-up scaling analysis of the balance and vent line mass flow rate which is tied to the top-down depressurization scaling analysis of Section 5.0.

Section 7.7 presents the bottom-up scalinF analysis for upper core support plate draining. Section 7.8

                . presents a scaling analysis for the reactor vessel downcomer processes.

The parameters that inve been scaled to address the phenomena of interest of the venting, draining,

                 . and injection processes are as follows:

• Accumulator intemal volume

                         * - Accumulator liquid and N2 volume e    'IRWST liquid level 368Iw.7.non:Ib40897                                                 7.]

FACILITY SCALING REPORT

• IRWST liquid volume
• CMT injection line diameters
• CMT injection line venical lengths
• CMT injection line resistance
• ACC injection line diameters
• ACC injection line venical lengtns
• ACC injection line pressure drop

             =   M'.WST injection line diameters

• IRWST injection line vertical lengths
• IRWST injection line resistance
• DVI line diameter

           ~*    DVI line length

• DVI line resistance
• Cold leg-CMT balance line diameters
• Cold leg-CMT balance line length a Cold leg-CMT balance line resistance
• ADS 1-3 vent line diameter
• ADS 1-3 vent line length
• ADS 1-3 vent line resistance
• ADS 1-3 vent line sparger
• ADS 4 vent line diameter
• ADS 4 vent line length
• ADS 4 vent line resistance
• Upper core support plate geometry
• Downcomer length
• Downcomer gap width
• Downcomer stored energy release rate to fluid volume
• DVI diffuser geometry 7.1 Depressurization Scaling Requirements for Venting, Draining, and Injection Processes l

This section pre ents the depressurization scaling requirements that are applicable to senting, draining, and injection processes. 'Ihese scaling requirements were developed in Section 5.0 of this report. l l

                                                                                                                     )

l l \

• l l 3681w.7.non:1b-060897 7-2 I  !

l

FACILITY SCALING REPORT I l ' {f% 'Ihe first requirement specified by Equation 5-82 is that the total mass flo.v rate injected into the primary system be scaled as follows: { l Th = R (71) I i As shown in Figure 3-2, the AP600 utilizes three passive safety injection systems:

• CMT injection system

                                                    =   Accumulator injection system
                                                  -*    IRWE T injectica systeri.

In addition, the non-safety injection systems can be used:

• NRHR system a CVCS charging system

  ,                                           Equation 7-1 can be expanded as follows:

MT* bACC+ b!RWST

• bNRHR+ bCVCS)R This scaling requirement can be satisfied by imposing the identity given by Equations 5-94 and 5-95 as follows:

                                                                                                                                      .        b0)R                           (73)

CMT)R 96

                                                                                                                                                 'O R                         (7-4)

ACC)R 96 O 36si. 7.non:1b-060897 73 e =- - ____-________ _ _

FACILITY SCALING REPORT (7-5) O IRWST)R96~ [PTP,o)R [D- )R (7-6) F=xRnR]R 96 [PTP,o)R (7-7) [Em.cycs)g " g Fluid property similitude is maintained for all of the injection processes- with the exception of CMT injection following prolonged CMT recirculation. CMT processes are discussed in detail in Section 6.0. By assuming fluid property' similitude for the injection process, the following would be true: (h,n,o)R O Therefore the injection energy flow rate will be properly scaled if Equations 7-3 thmv.3.i: 7-7 are satisfied. Substituting Equation 7-8 into Equation 5-89 yields the scaling requirement for the vent line and balance line energy flow rates. That is: E (rhh),,o , = pm,]R where the subscript e refers to conditions at the exit of the vent line or break. Substituting Equation 7-1 into Equation 7-9 yields the following: R (7-10)

                                                                                                        'E (rhh),*0                   =
                                                                                                                                . R        96 O

l 3681w-7.non:ltm897 7-4 _ = _ _ .

FACILITY SCALING REPORT (' ty\ imposing the identity given by Equations 5-94 and 5-95 yields the scaling requirements for specific components: E (sh) ads,1-3 g" bD [PTPo) E (rit) ads.4, g

• 96 l

l E (sh)34t, y = M 1 j., ' Equation 5-69 requires that all vent line and balance line mass flow rates be scaled by the following: l (/ l [PTP o) (EsOQAL,R 96

                                                                         ' Applying the identity given by Equations 5-94 and 5-95 to the ADS lines and the balance lines yields                                                                                                                             l the following:
                                                                                                                                                      * (PTP,o)R                                                                                                                                   (7 15)

BADS,1-3)R 96 , [PTP,o)R (7-16) SADS,4)R 96 6 3681w-7.non:Ib4297 75 o:_ _ :._ a_ _ _ _ __.____._u_ .____._u_________.______ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . . _ . . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

FACILITY SCALING REPORT

                                                                    ,  [PTPo)R                                                         (7-17)

O BAL)R 96 h 'Aese scaling requirements are implemented throughout the remainder of Section 7.0. 7.2 General Scaling Analysis for Tank Draining Processes This section develops a set of balance equations and R groups which are applicable to draining processes in general. Figure 7-1 defines the control volume boundaries for a typical safety tank. Table 7-1 presents the control volume balance equations for the tank draining processes. Equations 7-18 and 7-19 are the fluid conservation equations for the tank. The following nomenclature is used in these equations: p, = density of the ligtdd inside the control volume } V, = volume of liquid inside the control volume A[pQ] = net rate of fluid mass entering or leaving the control volume qt = net heat rate into the control volume In Equation 7-19 (found in Table 7-1), e, consists of internal energy, kinetic energy, and potential O energy as fcilows: 2 e,=ei,+ + gz (7-20) The following term is the net rate of liquid energy entering or leaving the control volume. A pQ e,+ P" _ (7-21) Ps. This term includes the work done by pressure on liquid flowing across the control surface boundaries. Equations 7-18 and 7-19 describe + average liquid ma~ and energy transfer for the control volume. The following assumptions will be applied to the governing equations:

1. The liquid inside the control volume is incompressible.

2. There is no work done by the liquid.

3681w.7.rmn:ltW4897 7-6

FACHJTY SCALING REPORT 1 O ~ Q '

3. : Changes in kinetic anci potential energy internal to the liquid inside the control volume are small compared to its internal energy.

The liquid volume inside a tank having a constant cross-sectional area (arggg) over the entire height of liquid (L,(t)) is given by the following: V, = STANKb t(t) (7-22)- l Substituting Equation 7-22 into Equation 7-18 yields the following: Pt4ANK (b t) = A[pQ) (7-23) 'i Recognizing that the mass flow rate s equals pQ, Equation 7-23 becomes the following: 91 DANK (bt ) " b -Eb2 1 (7-24).

v  ;

                                              ' Applying Equation 7-20 and the assumptions stated above to the liquid energy equation
                                            . (Equation 7-19) yields:

Paa7Agg Cygf,)L,) = E(sh) -I(thh)2 9 (7-25) These equations are summarized in Table 7-2 and can be expressed in terms of dimensionless quantities using the initial and boundary conditions presented in Table 7-3. Substituting the dimensionless quantities into Equations 7-24 and 7-25, dividing the liquid mass conservation equation by (IA 2)o, and dividing the liquid energy conservation' equation by (T4ri>' ',), yields the non-dimensionalized balance equations listed in Table 7-4.

j. 7.3 ? Accumulator Scaling Analysis lq This section presents the scaling analysis for the accumulator tanks. The AP600 accumulators (ACC)

Q ~ consist of two,2,000-cubic foot tanks filled with borated water and pressurized with nitrogen to

                                               ' 700 psig. For depressurization events which result in primary system pressures below 700 psig, the l

[ 36si.-72on::tM)60897 ' 7-7'

l FACILITY SCALING REPORT I l check valves which isolate the contents of the accumulators from the primary system open due to differential pressure. Thit . mits the accumulators to inject coolant through the DVI lines into the reactor vessel downcomer. Accumulator injection systems are also used in standard PWRs and are designed to provide a large volume of coolant at intermediate pressures. l In general, the phenomena of interest during accumulator injection are as follows:

              . Accumulator injection rate

• Charging gas injection Figure 7-2 is a flow diagram for the accumulator scaling analysis. First, a top-down scaling analysis is performed to obtain the trransport time scaling constant and the characteristic time ratios that govem the ACC injection process. Following the top-down scaling analysis, a bottom-up scaling analysis is performed to scale accumulator nitrogen injection.

The parameters that have been scaled to address these phenomena are as follows:

• Accumulator charging pressure
• Internal liquid volume
• Internal gas volume a Fluid initial conditions
• Gas discharge mass 7.3.1 Top-Down Scal *ng Analysis for Accumulator Injection The general non-dimensionalized balance equations for safety injection tanks presented in Table 7-4 can be simplified to reflect ACC conditions. That is:

No heat loss; q, = 0 No mass flow into the ACC; rh, = 0

• A single injection nozzle connection to the DVI line a Reversible-adiabatic expansion Thus, the non-dimensionalized balance equations for ACC injection are given in Table 7-5.

O 4 I l 3681w-7.nortItMt;0897 7-8 j j L-.--_____________--____---._.-________ . ___ __ J

i FACILIT) SCALING REPORT (G

 ,4
    !   Accumulator injection occurs during system depressurization. The importance of this process to                                                                                                           j
      ' system depressurization can be determined by comparing the accumulator fluid residence time constant to the system depressuritation time constant. That is:                                                                                                                                                   ;

T ACC , Pt.ACC V,ACCE$e f (7-33) Tsys VTPPTPbACC ,o l

      -If this ratio yields numerical values on the order of unity, then the accumulator injection process can be considered to be strongly coupled to the depressurization process.

The time constant and ratios presented in Table 7-5 are evaluated follow" the bottom-up scaling analysis. 7.3.2 Bottom-up Scaling Analysis for Accumulator Gas Carryover AP600 Accumu!ators O V The AP600 accumulators are initially charged to 714.7 psia (4.9 MPa) with 300 ft3 (8.5 m3 ) of nitrogen. For a reversible, adiabatic expansion .of a perfect gas, the following equation is applicable:

                                                                                                 ,7 P final                      Vinitial
                                                                               ,                                                                                                                  (7-38)

P initial, ,Vgo,i , where: P = gas pressure V . = volume occupied by the gas y = ratio of specific heats (C/C y). For nitrogen, y = 1,4. initial = initial conditions final = final conditions The pressure at which the AP600 accumulator empties of liquid can be calculated using Equation 7-38 by setting Vnng equal to 2000 ft3 (56.6 m3). The resulting pressure is 50.2 psia (346 kPa). I l i L.) 36stwn.nonab 060897 7-9 t - - -- _ _ _ _ _ _ _ _ _ - - --- - . _ _ _ _ _ _ _ _ _ . - _ _ _ _ _ _ _ - - - - - - - - - - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

FACILITY SCALING REPORT Similarly, the volume of nitrogen injected (V inj ) into the AP600 primary syste.a can be calculated by setting the fmal pressure to 14.7 psia (101.3 kPa) in the following equation: r ' lly l Piniti al I _y (7-39) V.inj. = V.iniuta

                                                ..l                                       ACC pfins.1, The volume of nitrogen injected from each accumulator is 2,810.5 ft3 (79.6 m3) at expanded conditions. Through conservation of mass considerations, the mass of nitrogen injected is given by the following:

V iI M inj

• y final Mg,3 (7-40)

The mass of nitrogen ejected is 566.6 lbm (257.5 kg) per accumulator. In addition to the nitrogen mass injected into the primary through gas expansion, a significant quantity of nitrogen will come out of solution as the accumulator liquid depressurizes. The amount of nitrogen contained in the accumulator liquid is determined using Henry's Law. That is: PN2 = Hf (7-41) where: PN2 = Partial pressure of the nitrogen cover gas H = Henry's constant, which has been tabulated for various gases X = ratio of moles of nitrogen to moles of solution Assuming an accumulator liquid temperature of 120 degrees Fahrenheit (48.9 degrees Celsius) for the AP600, the partial pressure of the nitrogen cover gas is 712.39 psia (4.91 MPa), Henry's constant is 5 found to be 1.985 x 10 atm / (moles N2/ moles solution). Substituting these values into Equation 7-41 in the appropriate units yields a value of X equal to 2.444 x 104. Knowing the initial number of moles of water in the accumulator and the value of X yields the moles of N2 in solution. 368Iw.7.non:Ib-060897 7-10

FACILITY SCALING REPORT Assuming all of the N 2comes out of solution, each accumulator releases 40.0 lbm (18.18 kg) of N 2 into the primary. APEX Model Accumulators Equations 7-38 through 7-41 can also be applied to the OSU model to determine 33 P ,i, Vig and the volume of N2in solution.- The initial charging pressure of the model accumulators is set by Section 5.0 depressurization requirements. That is:

                                                                                                              , (Pinitial)AP600                                  (7-42)

(Pnitial)model 3.086 Thus,(Pjoiti,3)model equals 231.7 psia (1.60 MPa). Similarly, the. final pressure is 16.27 psia (112.2 kPa). This is required to assure proper timing of events and appropriately scaled injections flow rates. This result is achieved by satisfying the volume scaling criteria for both the liquid and nitrogen. Thus, O [V ini ,;,i]AP600 (7-43) finitialbodel 192 i l and [ VACC}AP600 (7-44) ACChodel 192 i This yields an initial nitrogen volume in the model accumulator of 1.56 ft3 (0.044 m3) and an 3 3 accumulator tank volume of 10.42 ft (0.295 m ). Using the nitrogen and accumulator volumes specified by Equations 7-43 and 7-44 and performing the same analysis as for the full-scale accumulator.

            ,6                                        Note that to properly scale the amount of nitrogen introduced into the primary system, an additional

() 2.91 lbm (1.32 kg) of nitrogen per accumulator must be introduced into the primary subsequent to 36siw.7.non:ttuan 7.I1

                                                           ~

L ___ _ _ __----__ ____--------_--_ _____ _ - - - - - - _ - _ _ - - - -

FACILITY SCALING REPORT emptying the accumulator of liquid. This can be done by attaching a small nitrogen charging flask to each of the a:cumulator injection lines. Another approach to scaling the accumulator injection process is to select an initial nitrogen volume in the model such that the accumulators are empty of liquid at approximately 50 psia. This distorts the injection flow rates slightly but satisfies the transition pressure acquirement of Section 5.0. Table 7-7 presents the model accumulator conditions for complete accumulator liquid discharge at 50 psia. Because the ADS reduces system pressure quite rapidly, the time elapsed between 50 psia and 15 psia is small compared to the rest of the transient. Therefore, the values in Table 7-6 will match the injection flow rates while resulting in a minimum amount of distortion when ADS is operational. 733 Evaluation of Accumulator Injection Process and Scaling Distortions The purpose of this section is to evaluate the accumulator fluid residence time constants and ratios presented in Table 7-5 and to examine process distortions. Table 7-8 presents the numerical values for these parameters, as defined in Table 7 5, for the AP600 and APEX. As demonstrated by Table 7-8, the accumulator fluid residence time in APEX is scaled by one-half as required. The accumulator liquid specific heat ratio is preserved in APEX because the ratio of the liquid internal energy to its enthalpy is essentially unity, indicating that the accumulator pressure has little impact on the fluid properties. l Figure 7-3 is a flow diagram for the IRWST scaling analysis. First, a top-down scaling analysis is performed to obtain the characteristic time ratios that govem IRWST draining and heat up. Next, a bottom-up scaling analysis is performed to scale containment back pressure and condensate return rate. The primary distortions for the APEX accumulator injection process, using the scaling approach presented in Table 7-6, arise in two creas:

• The total mass of nitrogen gas discharged into the primary system subsequt + , . ..iptying the accumulator liquid is low by 2.91 lbm (132 kg). This can be corrected using 4 supplemental nitrogen flask.
• The pressure at which the APEX accumulators empty of liquid is approximately 30 psia less than that of the AP600. Because the ADS rapidly depressurizes the primary sy. item, the time for the pressure to decrease from 50 psia to 20 psia is very short. Therefore, emptying the accumulator at the lower pressure has a negligible impact on the overall transient when the ADS is operational.

The primary distortion introduced by using the scaling approc.ch presented in Table 7-7 is that the injection flow rates are larger than the required scaled value. 3681w.7.nort ib-060897 7-12

FACILn Y SCALLNG REPORT It is recommended that the scaling approach presented in Table 7-6 be implemented. If necessary, the mass of nitrogen discharg A into the primary can be varied parametrically to study its effect on the transient. As indicated by Table 7-8, the accumulator injection process will be adequately simulated in APEX. 7.4 In Containment Refueling Water Storage Tank Scaling Analysis This section presents the scaling of the IRWST. The AP600 IRWST injection system consists of a 530,000 galica water container, which is elevated above the reactor core. Subsequent to primary sys, tem depresturization below 28.9 psia, due to a LOCA or an ADS blowdown, the liquid level in the IRWST provides sufficient gravity head to begin draining its contents into the reactor vessel through the DVI lines. In general, the phenomena of interest during the IRWST draining process are as follows:

• IRWST draining rate
• IRWST liquid heat up rate Containment condensate return rate
• Containment backpressure Figure 7-3 is a flow diagram for the IRWST scaling analysis. First, a top-down scaling analysis is performed to obtain the characteristic time natios that govern IRWST draining and heat up. Next, a bottom-up scaling analysis is performed to scale containment back pressure and condensate return rate.

The parameters that have been scaled to address these phenomena are as follows:

          *  . Intemal liquid volume

• Liquid initial conditions
• Liquid height
• Liquid surface area a Heat sources
• Relative elevations Provisions have also been made to parametrically vary containment back pressure and containment condensate retum rates.

O 368Iw-7.non:1b-ocos97 7 13

FACILITY SCALING REPORT 7.4.1 Top Down Scaling Analysis for In Containment Refueling Water Storage Tank Injection The general non-dimensionalized balance equations for safety injection tanks, presented in Table 7-4, can be simplified to reflect IRWST conditions. That is:

• Energy input is due to PRHR, ADS 1-3, and containment condensate.
• Mass input is due to ADS 1-3 and containment condensate.
• Heat loss and tank structure energy transport is unimportant.

Thus, the non-dimensionalized balance equations for IRWST injection are given in Table 7-9. The IRWST draining rate time constant and H groups are evaluated in Section 7.4.4. 7.4.2 Bottom up Scaling Analysis for In Containment Refueling Water Storage Tank Injection This section examines the scaling of the containment feedback processes (containment condensate return rate and containment backpressure) and the IRWST liquid heat up rate. Contalament Feedback Processes To evaluate the H groups presented in Table 7-9, models to predict the containment pressurization rate and the containment condensate retum rate are needed. These models can be in the form of three-dimensional code calculations specific to AP600 containment pressurization and cooling. Evaluation of these models are beyond the scope of the present scaling analysis. However, to providc maximum flexibility in testing, provisions have been made to pressurize the IRWST and sump tanks up to 65 psia (0.448 MPa). Thus, full containment pressure can be simi' lated. In addition, die APEX model is designed so that a range of containment condensate return rates can be tested during IRWST operation. For purposes of design, the APEX test facility is capable of providing approximately 6.8 lbnVs (3.1 kgh) of heated liquid to the IRWST. This scaled amount would correspond to approximately 650.0 lbm/s (295.5 kg/s) of containment condensate retum in the full scale plant. This value exceeds maximum estimates of steam release rate from the AP600 during a double-ended DVI line break. (Based on NOTRUMP calculations.)(l) O 368Iw 7.nortIb-060897 7-]4

FACILITY SCALING REPORT 9 In summary, the APEX test facility model is designed so that parametric testing of the effects of containment backpressure and condensate retum rates can be performed. The facility is designed to simulate maximum containment pressures and a wide range of condensate retum rates as follows: 14.7 psia s Pconi s 65 psia (7-55) and 0 $ the, s 7 lbm/s (7-56) IRWST Liquid lleat Up Rate Subsequent to actuation of the ADS, there are two energy sources for the IRWST liquid: l

                                               =

ADS 1-3 vent flow through the spargers

• PRHR system heat transfer A

i i (,/ . Both the PRHR heat exchanger and the ADS 1-3 spargers preferentially heat the liquid in the upper l portion of the IRWST. Therefore, the temperature of the liquid in the bottom of the IRWST remains at its initial value. Proper modeling of the IRWST requires that the PRHR heat transfer rate and ADS 1-3 energy flow l rate be scaled. PRHR heat transfer has been discussed in Section 4.0, and ADS 1-3 venting is  ! discussed in Section 7.6.1. 7.4.3 Evaluation of In Containment Refueling Water Storage Tank Dimensions , This section implements the results of the top-down and bottom-up scaling analyses to determine the physical dimensions of the IRW3T. f3

                         ]

368Iw.7.non:Ib-060897 7-15

FACILrrY SCALING REPORT l In-Containment Refueling Water Storage Tank Liquid Surface Area The drain rate time constant given by Equation 7-47 can be written in terms of a model to prototype ratio as follows: l

                                                                         . Pt3h ts                                      (7-57)

(TIRWST)R NI IRWST. o.R Substituting Equation 7-5 into Equation 7-57 yields the following: [tIRwsT)g "%

                                                                              . O. o.R Because all of the time ratios have been set to 0.5, all of the length sale ratios have been set to 0.25 and fluid property similitude exists during IRWST injection, Equation 7-57 simplifies as follows:

O [a,]R nus, to achieve properly scaled IRWST draining rates, the liquid surface area should be scaled in accordance with Equation 7-59. The AP600 IRWST liquid surface area is 2528.5 ft2 (234.9 m 2), Therefore the liquid surface area for the APEX model should be 52.68 ft2 (4.89 m2 ). For a cylindrical tank, this corresponds to a tank diameter of 8.19 ft (2.5 m). In Contalmnent Refueling Water Storage Tank Liquid Volume The required IRWST volume scaling ratio is 1:192. The liquid volume of the full scale IRWST is 70,798 ft3(2,005 m 3). Herefore the ideal liquid volume of the APEX model IRWST is 368.74 ft 3 3 (10.44 m ). In. Containment Refueling Water Storage Tank Water Level The normal water depth in the AP600 IRWST is 28 ft (8.53 m) and the overflow depth is 29 ft (8.84 m). Applying the 0.25 length scaling ratio yields a water depth of 7 ft (2.13 m) and an overflow depth of 7.25 ft (2.21 m) for the APEX model. f 368Iw.7.non:1t>060897 7-] 6

r I ( Summary of Model In Contair. ment Refueling Water Storage Tank Dimensions i Table 7-10 summarizes the critical attributes of the model IRWST based on the scaling analysis. The j following sections examine scaling distortions which arise as a result of implementing the selected IRWST design.  ! l 7.4.4 Evaluation of IRWST Processes and Scaling Distortions l l i The purpose of this section is to evaluate the IRWST fluid residence time constants, the injection and { liquid heat up process-specific frequencies, the characteristic time ratios, and the scaling distortion l factors. Table 7-11 presents the numerical values for these parameters, as defined in Table 7-9, for the 1 AP600 and APEX. These estimates are based on NOTRUMP calculations for an AP600 two-inch cold , leg break.W The initial conditions for IRWST injection are determined for the AP600 at a downcomer pressure of 27 psia (186.1 kPa). ) 7.5 Safety Injection Line Scaling Analysis This section presents the scaling analysis for the CMT, ACC, and IRWST safety injection lines. A top-down scaling analysis has already been performed in the previous section for accumulator and q IRWST draining and in Section 6.0 for CMT draining. The pertinent scaling requirements for V injection line mass flow rate have been established and are given by Equations 7 3 through 7-5. A bottom-up scaling analysis is performed to determine the dimensions and pressure drop requirements for each line. The parameters that have been scaled to address the injection line mass flow rate requirement are as follows:

• CMT injection line diameters
• CMT injection line vertical lengths

        *   . CMT injection line resistance

• ACC injection lice diameters
• ACC injection line vertical lengths
• ACC injection line resistance
• IRWST injection line diameters
• IRWST injection line vertical lengths l
• IRWST injection line resistance
• DVI line diameter
• DVI line length
• DVI line resistance O

36si.-7.rm.b-060897 7 17 t - . - _ _ . - - _ _ _ _ _ _ _ _ _ _

FACILTIT SCALING REPORT 7.5.1 Bottom Up Scaling Analysis for Safety Injection Lines Equations 7-3 through 7-5 present the mass flow rate scaling ratios that must be satisfied to properly simulate the depressurization process and event sequence timing. By carefully selecting the injection line diameters, length and resistances, the mass flow rate scaling ratios can be satisfied. A general model for safety injection mass flow rate can be developed using the governing equations presented in Table 7-12 as applied to a section of the safety injection line control volume presented in Figure 7-4. The following assumptions are applied to the i* section goveming equations presented in Table 7-12:

1. Mass and energy will not accumulate in the injection line during the injection process.

2. The liquid is incompressible.

3. There is no heat transfer across the control volume boundaries (that is, qe = 0).

4. There is no work done by the liquid (that is, we = 0).

During the injection process, the liquid temperature (Tg ) the injection line piping temperature (T,) and the ambient temperature (Tamb) are essentially identical. Therefore, the solid structure energy equations are not implemented in this analysis. Applying the second assumption to Equation 7-60 yields the following: (thin)i =(rhout) (7-62) I where: th = mass flow rate; equals pQ (rhin); = mass flow rate inte the i* section (rhout)i = mass flow rate out of the i* section Applying Assumptions 1,3, and 4 to the fluid energy Equation 7-61 yields the following: (rhin), e, +_P' = (thou,) e, +_P ' . (7-63) P .in.1 P.out.i O l l 368Iw.7.non:Ib.o72297 7.] 8

FACILITY SCALING REPORT ( Substituting Equations 7-20 and 7-62 into Equation 7-63 yields the following: eint + gz+1 = eing + 2

                                                                                                                                                  +gz +1                               (7-64)   j I. in,i     ,

pl.outi Rearranging this equation yields the following: 2- 2

                                                                                                             "'                                                                        (7 65) i" g Az + $ +Aeint                 =0 2                       p, l

where the liquid density is assumed to be constant. The change in liquid internal energy (Aeint) is directly attributed to the friction and form losses within the injection line control volume. Therefore: I 2 [% Ae;ng = 'f bd +fH y y dH

                                                                                                                                             +Kro ,    12                              (7-66)
                                                                                                                 ,                                  ,i 1

l l Substituting Equation 7-66 into Equation 7-65 yields the following: 2 2 ' ' 2

                                                                                          ""'~"i"
                                                                                                  + gAz +    **         I" + f y                        + Kro,                     =0  ( -67)
                                                                                                                                             + fu
                                                                                                                                                                     ,          ,i Defining the acceleration pressure drop for the i* section as follows:

2 2 2 u gog -u in = K A 1 (7-68) 2 2 rg - - l 36a:w-7.non:1b-(r72297 7-19 {' _

FACILITY SCALING REPORT and substituting into Equation 7-67 yields the following: f 2 P f g h ggh +K ronn+KA gAz+ out in . =0 pt , vd y du ,2 i l i Summing the equation from z = 0 to z = [Ly +1q (t)] yields the following: P TANK , g h ,f 1 [L, +LT (t)]+ oy, f Hd +Kronn +K4 =0 (7-70) p, i .i , v d, g , 2 For incompressible flow within the different sections of pipe that form the injection line: i"I (7-71) ui = uinj ai Therefore, Equation 7 70 becomes the following: 2 P 1 (L v+Lpt)]+ DVI TANK , p , nj r0 (7-72) where:

                                                    ,                               ,    ,     ,3 F7 ,in) = bfb+fnb +Kronn 3+K                        ini                                                                         (7-73) il< y d,         du               ,,,a, Rearranging Equation 7-72 and solving for the injection velocity:

H

                                             =     g[LT(t)+L ]+2 y     [PTANK -Poygyp,                                                                                   <f,74) u*>

FT,inj 368 Iw-7.nortIb-060897 7 20

FACILITY SCALING REPORT r's j] _ Equation 7 74 can be written in terms of scaling ratio as follow,: H E T( ) v) TANK-Poygyp, (uinjhg " (7-75) W < TANK,R l Equation 7-75 can be written in tenns of an injection mass flow rate ratio as follows: j [EiNHANK,R "[injPf)R FT,inj TANK,R Substituting Equations 7-76 into Equation 7-1 and rearranging yields the following: 2 [T,injhANK,R' %nj ,8[bT (t) {y]+ (PTANK -Poyjfpl TANK,R rx pTPo. TANK,R

        \               ;

During CMT injection, when the balance lines are filled with steam, the injection mass flow rate is primarily driven by the liquid level in the CMTs because the CMT balance lines serve to equalize  ! PTANK and Poyi. Thus, for the CMTs, Equation 7 77 becomes the following: I 2 T,injkMT.R ) YbMT,R

                                                                                                     .    . CMT,R 1

i ! O E

       'd 1

3681w-7.non:lN97 7 21 t

FACILITY SCALING REPORT For the accumulator injection lines, the mass flow rate is primarily driven by the static pressure difference between the accumulator tank and the DVI line. Liquid density similitude is maintained. Thus, for the accumulator injection line, Equation 7-77 becomes the following: 2 (7 @ ) [T,id)ACC,R ACC -Poyi), pTPo ACC,R During steady IRWST injection, the mass flow rate is dominated by the elevation difference and fluid property similitude exists in the system. Thus, Equation 7-77 becomes the following: [T,inj)IRWST,R d RWST,R T(I) +b vhRWST,R Table 7-13 summarizes the line resistance scaling ratios for each of the injection systems. > l To evaluate each of the scaling ratios presented in Table 7-13, detailed calculations of the friction and form loss coefficients (F 7,inj) and of the relationship between tank pressure and DVI line pressure are needed. Friction and Form Loss l By using orifices in the model injection lines, the value of F ,in; 7 can be adjusted to achieve the desired injection mass flow rates provided that the line diameters are not too restrictive. In contrast, the line diameters are not made so large as to significantly distort the fluid volume requirements of the injection line. Because PTANK can be controlled in the APEX model accumulator and DVI line pressure is controlled by system depressurization (which has been scaled using the criteria developed in Section 5.0), a scaling relationship for AP can be developed for the accumulator. O 368Iw.7.non:Ib-060897 7 22

FACILITY SCALING REPORT \ !. Accumulator AP l l l^ . From Equation 5-103 it is observed that the ratio of model to AP600 accumulator tank pressure is given by the following: (PAcc)R O'N Similarly, during the accumulator injection process, the DVI pressure ratio is given by the following: (Payj), = 1:3.085 (7-85) Thus, the model to AP600 pressure difference ratio is given by the following: [PAcc +Dvi)g " WM (M) 7.5.2 Evaluation of Safety Injection Line Dimensions Ideally, all of the vertical lengths should satisfy the length scaling requirement R(L = 0.25): the diameter ratio should be Dg = 0.1443; the volume ratio should be VR = 1/192; and the total friction and form loss ratio should be (F7,3,;)g = 1. However, because standard pipe and tubing dimensions do not exactly match the ideal dimensions and some flexibility in varying line pressure drop is prudent, a slightly larger line diameter was generally selected and (FT,inj)R was adjusted through the use of an orifice to satisfy the scaling requirements. The scaling ratios presented in this section implemented to size the safety injection lines for the maximum injection flow conditions. Core Makeup Tank Injection Line Dimensions By imposing the requirement that LR = 0.25 on Equation 7-78 and assuming fluid property similitude, the following scaling ratio esn be obtained for the CMT injection lines: (FT,inj)CMT,F. " ( inj MT.R i 36sIw.7.nortib-060s97 7 23

I ACILITY SCALING REPORT l Expressing Equation 7-87 in terms of injection line diameter yields: l I fT,inj)CMT.R id C M T,R The AP600 CMT injection line has a 6.81 in (17.3 cm) inside diameter. Ideally the model CMT injection line should have an inside diameter of 0.983 in. (2.50 cm). A standard diameter of 1.06 in. (2.69 cm) was selected. This yields a diameter ratio, (Dinj)CMT,R equal to 0.1557. Substituting this value into Equation (7-88) yields (F 7,;,))CMT,R equal to 1.354. Table 7-14 presents the AP600 and APEX model scaling ratios and dimensions for the CMT safety injection lines. Accumulator Injection Line Dimensions Substituting Equation 7-84 into 7-82 and assuming fluid property similitude yields: [T,inj)ACC,R ' IN ACC,R The AP600 accumulator injection line has a 7.98 in. (20.27 cm) inside diameter. A standard diameter O of i.195 in. (3.04 cm) was selected for the model accumulator injection line. This yields a diameter ratio of 0.1497. Substituting this value into Equation 7-89 yields (F ,ig)ACC.R 7 equal to 1.50. Table 7-14 presents the accumulator injection line dimensions. In-Containment Refueling Water Storage Tank Injection Line Dimensions Requiring Lato equal 0.25 in Equation 7-83 and assuming fluid property similitude yields the following: fT,iN)RWST,R id WST,R Each of the two AP600 IRWST injection lines consist of a section of pipe that connects the IRWST to the sump tee and a section that connects the sump tee to the DVI tee. Because the APEX model IRWST geometry and location is different than the AP600, additional horizontal lengths of injection line piping are needed to connect the variou.c 1RWST injection line components. This required l increasing the injection line diameters to compensate for the additional friction and form losses introduced by the atypical horizontal lengths of the IRWST injection line in the APEX model. j Table 7-14 summarizes the dimensions for the IRWST injection lines. 368Iw-7.non.Ib-060897 " 24

FACILITY SCALING REPORT 7.5.3 Direct Vessel Injection Line Scaling ' Analysis The AP600 uses two direct vessel injection (DVI) lines to provide emergency coolant to the core. Coolant from the CMTs, ACCs, and IRWST is fed to each DVI line into a diffuser nozzle located

                                                                                      ~ inside the reactor vessel downcomer. 'Ihe intent of direct vessel injection is to enhance mixing in the downcomer and reduce coolant' bypass out of the break ~during the initial phase of a large break LOCA (LBLOCA).

Using the same approach presented in Section 7.5.1 for safety injection lines leads to an expression similar to Equation 7-77 for the DVI line. That is: r ,2 [T,DV1)R (9 (gly +AP/p,(y,,, MO s pTP,% DVI,R where: Ly,pyi = verticallength of DVIline JAPoyi = pressure drop from the DVI tee to the inlet of the DVI diffuser nozzle

    . . .                                                                           . an y,           = cross-sectional flow area of the DVI line 1 -                                                                                     FT,Dv1 = total friction and form loss coefficients The total function and form loss coefficients are given by the following:

F7,pyi = fbD +K% (7-92) .

                                                                                                                                                                         . Dv1 Equation 7 9) is a' simpler form of Equation 7-77 because the DVI line has a constant diameter, thus piessure losses due to acceleration need not be considered. Lpyj represents the total length of DVI line, and Doy; is the DVI line inside diameter.
                                                                                     ' Ideally, the following conditions should be met for the DVI line:

(APy3]R o " D [FDVI)R" T I ( u usiw.7.non:ib472297 7-25 f ,

FACILITY SCALING REPORT [aDv1)g "RM G@ [Lv,Dvi)g *M @% Detailed calculations of DVI friction and form coefficients and pressure drop have been performed using Equations 7-91 through 7-96 as a guideline. The following section presents the results of these calculations. DVI Line Dimensions Requiring [Ly,9y3]g and [AP vi ao l to equal 0.25 and assuming fluid property similitude in Equation 7-91 yields the following: [FT,DVl)R o vi)4 2304(D R Table 7-15 demonstrates how Equation 7-97 is satisfied in the model. 7.6 Balance and Vent Line Scaling Analysis This sectior' presents the scaling analysis for the CMT balance lines, the ADS 1-3 train and the - sparger, and the fourth-stage ADS line. A cold leg-CMT balance line is used to maintain each CMT at system pressure. The ADS 1-3 vent lines route two-phase fluid from the ADS valve headers to the spargers located in the IRWST. The sparger breaks up the steam jet into bubbles to enhance steam condensation and to reduce the pot-ntial for condensation induced water hammer. The fourth-stage ADS vent line routes +wo-phase fluid from the hot leg to the containment sump. Fourth-stage ADS operations is necesaj to assure that primary system pressure remains below the setpoint required for gravity driven injection. In general, the phenomena of interest to CMi balance and ADS line performance are as follows:

• Two-phase mass flow rate
• Choked flow under break conditions
• Two-phase flow pattern transitions during CMT operations
• ADS sparger steam-liquid interaction 3681w-7.non:Ib-060897 7-26

FACILITY SCALING REPORT Figure 7-5 is a flow diagram for the balance and vent line scaling analysis. A top-down scaling analysis was performed in Section 5.0 to obtain the characteristic time ratio (Il group) that governs balance and vent line mass flow rate. The pertinent scaling requirements for vent and balance line mass flow rate are given by Equations 7-14 through 7-17. A bottom-up scaling analysis is performed to determine pressure drop characteristics for each line and to evaluate two-phase flow pattem transitions. The parameters that have been scaled to address balance and vent line mass flow rate are:

• Cold leg-CMT balance line diameters a Cold leg-CMT balance line length
• Cold leg-CMT balance line resistance

         =   ADS 1-3 vent line diameter

• ADS 1-3 vent line length
• ADS 1-3 vent line resistance

        =    ADS 1-3 vent line sparger ADS 4 vent line diameter

• ADS 4 vent line length

        =

ADS 4 vent line resistance 7.6.1 Bottom-Up Scaling Analysis for Balance and Vent Line Mass Flow Rate Equations 7-14 through 7-17 present the mass flow rate scaling ratios that must be satisfied to properly simulate the depressurization process and event sequence timing. The mass flow rate scaling ratios can be satisfied if the vent line and balance line diameters, lengths, and resistances are carefully selected. A general model for vent line mass flow rate can be developed using the goveming equations presented in Table 7-16 as applied to a section of vent line presented in Figure 7-6. The following assumptions are applied to the i* section goveming equations presented in Table 7-16:

1. Mass and momentum (that is, inertia is neglected) will not accumulate in the vent line during the venting process.

2. The fluid consists of a homogeneous two-phase mixture.

3. The cross-sectional flow area in the i* section is constant.

4. The mixture density in the i* section is constant.

Prior to initiating the venting process, the vent line piping is at the ambient temperature (Tamb)- Shortly after venting begins. the vent line piping reaches the saturation temperature at approximately I 368 w-7.non:1b.060897 7-27

FACILITY SCALING REPORT atmospheric conditions. Because these lines are insulated, heat loss from these lines subsequent to l l initial heat up is neglected in this analysis. In Equation 7-98, pf si the two-phase fluid mixture density defined by Equation 4-90. In Equation 7-99, M,; represents the frictional pressure drop in the section and Mu represents the local pressure drop due to form losses. Applying the first assumption to Equation 7-98 yields the following: (rh)=(thout), io (7-100) where: th = pQ (rhin)i = mass flow rate into the i* section (thoot)3 = mass flow rate out of the i* section Applying Assumptions 1 through 4 to the fluid momentum Equation 7-99 and rearranging yields the followSg: O JAPf ) = jpfLg,), +[M, + Mk); (7-101) The first term on the right-hand side of Equation 7-101 can be written as follows: (PfL g,) =(pf Lyg) (7-102) where (Ly); represents the vertical elevation difference of the i* section of vent line. The frictional pressure drop term can be written as follows:

                                                                                                             '                    2' (7-103)
                                                                                     *i   _f",L'PFi
                                                                                          ~

d, 2 ,- O 3681w-7.non:lb-o72297 7 28

FACILITY SCALING REPORT f - where:

fTP = two-phase friction factor di . = pipe diameter

           ' L; = total length ' of the i* section The local pressure drop (Mg) can be written as follows:

El i (7-104) My = 2  ; _ Substituting Equations 7-102 through 7-104 into Equation 7-101 yields the following: 2 ' JAP)=(9,Lg)+ +K (7-105) f f y TP , r sn 1 Summing the pressure drops for all of the vent line sections yields the following: N. N N 2 r '

                                                                                    -E 14 (AP                  f L yg), + E
                                                                                                                   ) =1-1-E (pf                                                     2 frpbd +K                     (7-106) 1-1                                                                      ,          ,,i l                                                                                                                                                                                                                                                          4 l             The integrated pressure drop can be related to the differential pressure driving the vent flow as follows:

N AP ,,, = -E(AP p. y )8 (7-107) l 14 p Because' the mass flow rate is constant throughout the vent line, the mass flow rate at the exit point can be related to the mass flow rate along any point in the vent line. Thus: thy ,,, = pfiuia; (7-108) I [N - LU i

           ' 3681w-7.nonilb-060897                                                                                                 7-29

1 l l FACILITY SCALING REPORT Substituting Equations 7-107,7-108, and 4-93 into Equation 7-106 yields the following: N g "2 ' APy ,,, = -E(pf Ly g) + F T, vent (7-109)

                                              ' "I 2 p ,say ,,,

where: 2 FT, vent " E ITPbd +K 1+ * "' (7-11% i =1 , ,i , psg s a; 2 Rearranging Equation 7-109 to solve for myent yields the following: 2p,3APvent + 2gpesE(pfv) L (7-111) Vent Vent . T, vent For venting processes, AP yent dominates over the gravitational term; therefore, Equation 7-111 can be simplified as follows: 2 p,3 AP vent (7 112) g vent ,a vent F T, vent Substituting Equation 7-14 into Equation 7-112 and rearranging yields the following:

                                                                     '2

[F,T vent)R = 96f [p,3APyeni]R

                                                         , PTe,o, g A more detailed analysis reveals that FT, vent can be modified to include the effects of vapor compressibility and acceleration pressure drops due to phase change.

O 3681w 7.non:lb-060897 7-30 l

FACILITY SCALING REPORT = 1 To evaluate Equation 7-113, detailed calculations of the two-phase friction and form loss coefficients are needed. APy,nt must also be evaluated under maximum flow conditions. Two-phase fluid flow

             . pattem transitions and ADS sparger performance must also be evaluated.

Two-Phase Friction and Form Loss Coefficients l: l ' A typical approach to estimating two-phase friction factors (f7p) and form loss coefficients (K) is to

use a two-phase friction multiplier based on the saturation properties of each phase. Equation 7-110 can be written as follows:(2) l l

F7,y,,, " l = +K 1+ "' j .i ' ' x pss . i a;2 (7-114) d I + , Ap 3

                                                           .-   Ess ,                  ,i l-where fhis the Darcy friction factor evaluated at the saturated liquid conditions.

By using orifices in the model balance and vent lines, the value of F , vent T can be adjusted to achieve O scaled vent flow rates. - The results of the calculations performed to optimize the balance and vent line D ~ diameters are presented in the following text. 1 Choked Flow in Balance and Vent Lines l The maximum attainable mass flow rates in the balance or vent lines are achievea under choked flow conditions. For ADS 1-3, this occurs.whenever the ADS sequence initiates at high pressure. For the I CMT balance lines, this only occurs if the line is broken. The length to diameter ratio (UD) for the lines being considered generally exceeds a value of 100. For flashing flows with IJD greater than 40, Fauske reports that a homogeneous equilibrium model l (HEM)'or equilibrium rate model (ERM) provides accurate predictions of choked flow mass flux.W Under choked flow conditions, the maximum pressure drop through the line that can be achieved is l given by APCRIT. Thus, Equation 7-113 is written as follows for choked flow conditions:

                                                                                   ,2

[Tvent)R =(96) - [phAPCRIT)R s oR s

j. . Fauske suggests that for two-phase flow at a given choked flow rate and quality, the pressure gradient

'V' at a given location reaches a maximum and finite value.W Therefore, it is desirable to match the 368Iw-7.non:1b-o72597 7 31 Rev.I

FACILITY SCALING REPORT pressure gradient in the model lines to those in the AP600. This is particularly important in lines that may exhit.it multiple choke points. As a minimum, the following criteria should be met: APcRrr

                                                                                                                                                                                ,y                                                              (7-116) b                           sR Since all length scale ratios are 0.25, the scaling criterion given by Equation 7116 reduces to the following:

(APCRIT)R " ~ This criterion is consistent with the non-choked flow criterion and is used to evaluate balance and vent line diameters. Two-Phase Flow Pattern Transitions in the Core Makeup Tank Balance Line The flow pattem transition of interest with respect to balance line mass flow rate behavior (under no-break conditions) is the transition from stratified flow to intermittent-slug flow or annular-dispersed flow. This transition was examined in Section 4.0 for the hot and cold legs. Using the dimensionless group developed by Taitel and Dukler,W a diameter scaling ratio of 0.1613 was found to adequately simulate the flow pattem transition. Thus: (Dggt),= W WQ WW This criterion is compared to the result obtained by implementing Equation 7-117 under choked flow conditions. ADS 13 Combined Vent Line Scaling In order to reduce the complexity of the test facility, a single ADS 1-3 vent line is used to simulate the two ADS 1-3 vent lines in the AP600. The combined ADS 1-3 vent line is scaled by requiring that: (PTP,o)R (7-119) (MADSI-3)R 48 O l 3681w-7.non:ii>.o60897 7-32 l I

l ( FACILITY SCALING REPORT I

                                                                                                                                                                                                                                                                                                                                   )
           )               Substituting Equation 7-119 and Equation 7-117 into Equation 7-115 determines the ADS 1-3 flow                                                                                                                                                                                                          '

l area needed to simulate combined ADS 1-3 flow rates. These criteria have been implemented to evaluate the ADS 1-3 vent line diameter By adjusting F , TADS 1-3, either a single train or both trains of ADS 1-3 can be simulated in the test facility. Table 7-17 summarizes the scaling ratios for the CMT balance lines and the ADS vent lines. 7.6.2 Evaluation of Core Makeup Tank Balance Line and Automatic Depressurization System Vent Line Dimensions I i l This section implements the scaling ratios presented in Table 7-17 to obtain specific dimensions for l the balance and vent lines. Ideally, all of the lengths should satisfy the length scaling requirement (L R= 0.25): the volume ratio should be Vg = 1/192; the maximurn pressure drop ratio should be i (APCRIT)R = 0.25; and the total friction and form loss ratio should be (FT,vem)R = 1. However, f because standard pipe and tubing dimensions do not exactly match the ideal dimensions, a slightly j larger line diameter was selected and (F T,vem)R was adjusted through the use of an orifice to satisfy I scaling requirements. Tables 7-18,7-19, and 7-20 present the AP600 and model scaling ratios and dimensions for the CMT balance lines and ADS vent lines. 7.6.3 ADS 13 Sparger Steam-Liquid Interaction n The ADS 1-3 sparger consists of perforated tubes connected to the end of the ADS 1-3 sparger line. The ADS 1-3 spargers are located below the water level inside the IRWST. Upon actuation of the ADS, a steam-water mixture flows from the pressurizer and vent into the IRWST where the steam is condensed. During ADS operation, choked flow occurs at the throat of each of the ADS 1-3 valves. Because of the significant expansion of the two-phase mixture entering the sparger line, choke flow is also likely to occur at the ADS 1-3 sparger during peak flow conditions-although at significantly lower pressure than at the valve throat. In APEX, a single ADS 1-3 sparger is used to model two spargers in the AP600. Thus, the scaling requirement given by Equation 7-119 are implemented. Rewriting Equation 7-119 in terms of mass flux yields the following for combined ADS 1-3 flow: _ (PTP,o)R (7 124)

(GCRIT asparger)R '

48 7

 !         l Q,/

368Iw.7.non:Ib-060897 7-33

FACILITY SCALING REPORT Rearranging this equation to solve for the sparger flow area yields the following: f 9 1 PTP.o (7-125) (#5Parger)R 48' G CRIT, R where the critical mass flux ratio is determined from the depressurization scaling analysis preented in Section 5.0. To model the steam-liquid interaction, the size of the holes on the sparger branch tubes must be selected carefully. 'Ihe fluid jet interaction length has been shown to be proportional to the hole diameter.W That is: L j,, or. Dhole O~@ Therefore, in terms of a scaling rado: Jet

                                                                                                                                                                                                                            =1                                          (7-127)
                                                                                                                                                                 , Dhole, a Ideally, to simulate the mixing behavior in the IRWST, the (Lj,t)g should be set equal to 0.25. This criterion need not be preserved in APEX because exact simulation of the sparger mixing behavior is not essential to overall transient behavior. To assure that the steam jet interaction length is shorter than the distance from the sparger arm to the IRWST liquid surface, (L),)g should be no greater than 0.5. This prevents jet penetration at the IRWST liquid surface.

The larger value of 0.5 greatly reduces the number of holes, thus reducing the cost of fabrication. Substituting the later value into Equation 7-123 yields the following: (Dhole)g ' M @M 3681w-7.non:Ib 060897 7-34

I. ! FACILITY SCA LING REPORT Rewriting Equation 7-125 in terms of a hole diameter yields the following: i l

                                                                                                                                                                                 'YS 6

ND[g*t R 48 =I G (7-129) CRITsR Substituting Equation 7-128 into Equation 7-129 yields the scaling criterion for the sparger hole number ratio (Ng): I PTP-Ng = 12 (7-130)

                                                                                                                                                                  ,G CRIT, R l                                             For purposes of determining the critical mass flux ratio (GCRrr)R, it is assumed that the pressure is L                                             approximately equal to 250 psia and the vapor quality is approximately equal to 20 percent in the l

AP600 sparger arms. 'Ihe equivalent conditions in the APEX sparger arms would be approximately L equal to 80 psia and approximately equal to 6 percent vapor quality. Rather than using the HEM for

                                      . critical flow which is better ' suited for higher pressures, the low-pressure experimental data of Faletti,
   ' \                                . Zaloudek, Fauske, and Moy as reported by Levy (7) has been used to determine the critical mass flux l                                             ratio for these conditions. The critical mass flux ratio was found to be 0.68 and (pTP,o)R was 1.12.

Substituting these values into Equation 7-130 yields the following: ( L ( Ng = 0.137 (7-131) 1- l l  ; Each AP600 sparger has four branch arms each having [ ]*C diameter holes. The total L number of holes for two spargers is [ ]" C. Using Equation 7-131, the number of holes in the model sparger branch lines should be the following: a.c (7-132) l j. The model sparger hole diameter should be the following: I-hhole)model 368iw.7.non:1b.060897 7-35

FACIIJIT SCALING REPORT Geometric similarity is preserved in the APEX ADS 1-3 sparger lines and sparger. This includes the sparger hole wall thickness-to-diameter ratio, the sparger line, sparger hub, and sparger arm internal flow area ratios. That is: Iwa!! 1 (7-134)

                                                                    , Dhole, R
                                                                     /
                                                                             \

aline 1 (7-135)

                                                                     ,a b,R hu ab hu 1                                  (7-136) am,R The ADS 1-3 sparger arm orientation is also preserved. Table 7-21 lists the ADS 1-3 sparger                  ,

dimensions for the AP600 and APEX. 7.7 Bottom-Up Scaling Analysis for Upper Core Support Plate Draining l During normal operation, the reactor vessel upper head is completely filled with liquid. During a depressurization transient some of the water in the upper head flashes to steam. The remaining liquid eventually drains into the core through perforations in the upper core support plate. The upper head dmining process involves counter-current steam-liquid flow through the perforated plate. This section develops the scaling criteria for sizing the diameters of the upper core support plate drain holes. Counter-current steam-water flow through perforated plates has been the subject of much study.(8 9* In general, it has been found that the onset of flooding through multiple tube paths shows similar characteristics to those of single-path flow. It was found that the likelihood of flooding increases with the increasing number of holes. This, however, was a secondary effect. The results of the studies indicated that flooding could be bounded by the following: (7-137)

                                                               ]{. + m (JQ.%      = C, O

3681w-7.non:lt>-060897 7-36

FACILITY SCALING REPORT where Jj is the Wallis parameter for the upward flowing steam. The Wallis parameter physically represents a ratio ofinertia forces to buoyancy forces as given by: u I. (7-138) g " (gDo Ap/pss) where: f Ap = ph - Pgs Do = core plate perforation diameter cp = core plate

              'J*                              = Wallis parameter for the downward flowing liquid reaching the core Jj is de. fined as follows:

u, I. " t (7-139) (gD, Ap/ph) s

             - In Equation 7-137, the values for the coefficient m were found to lie between 0.9 and 1.8, and those for the constant C, were found to lie between 0.9 and.1.7.

A criterion for the onset of upper head draining, in the presence of upward flowing steam from the core, can be found by setting Jj equal to 0. Thus, Equation 7-137 becomes the following: 7 1% iJ*1 =C* (7-140) i 84 2

             ' When the value of Jj drops below C , then draining begins. Equatio's 7-140 can be written as a scaling ratio as follows:

1 7 i J*l =1 (7-141) 4 8/ R 36stw.7.non: b-060897 7-37

FACILITY SCALING REPORT ) Substituting Equation 7-138 into Equation 7-141 yields the following: l u s =1 (7-142)

                                                                                                                                                                                            .bo Ap/pss)"     cP,R Upper head draining is most important under rapid depressurization transients, (that is, LBLOCA).

Since low-pressure conditions are achieved quite rapidly, fluid property similitude exists. Thus, assuming fluid property similarity and squering both sides of Equation (7-142) yields the following: ug

                                                                                                                                                                                                              =1                             (7-143)
                                                                                                                                                                                                .bo)     CP,R From Table 4-18, when fluid property similitude exists, then the core flow areas and fluid velocities are scaled as follows:

aR. = 0.0208 (1:48) (7-144) l 1 un = 0.5 (1:2) (7-145) Substituting Equation 7-145 into 7-143 and solving for the perforation diameter ratio yields the following: (Do )C , 6 M (m) The number of holes in the model upper core plate can be found by requiring that Equation 7-144 be satisfied as follows: 1 1 a fo o(P, APEX 48 )CP.AP600 0 368:w-7.non: b-060897 7-38

FACILITY SCALING REPORT l !'g ~ Rearranging to solve for the ratio of the hole number yields the following: O CP,R 0R li 1 or in 'erms'of hole diameter (Do): O CP.R 0R Substituting Equation 7-146 into Equation 7-149 yields the following: (No (p , = m M @W 1' l Table 7-22 summarizes the upper core suppon plate perforation scaling ratios and dimensions. ft . ( V.8 Reactor Vessel Downcomer Scaling Analysis He AP600 reactor vessel downcomer serves as an annular mixing region for the emergency coolant . ,  : injected by the CMTs, ACCs, and IRWST. During system depressurization, the downcomer rejects I heat'to the emergency coolant. For a rapid depressurization transient, of the type postulated for a LBLOCA, the downcomer walls would be superheated with respect to the coolant, resulting in significant vapor generation. His effect, coupled with reverse steam flow from the core (that is, cold leg break), would delay emergency coolant penetration into the lower plenum. Numerous studies of the phenomenon, known as emergency core cooling (ECC) bypass, have been performed.#W ne j majority of these studies are documented in NUREG-1230. The major conclusion derived from the L studies is that full-scale reactor vessels exhibit little or no emergency core cooling (ECC) bypass behavior. For slow depressurizations, of the type postulated for a SBLOCA, emergency coolant mixing in the downcomer is important. . During safety injection, it is desirable that the cold injected liquid mix thoroughly with the hot primary liquid to minimize the potential for thermal shock. For very low

                ' injection flow rates, the hot primary liquid may backflow into the DVI line.
                ' This section presents the scaling analysis for the reactor vessel downcomer and the DVI diffuser nozzle. The phenomena of interest to this analysis are as follows:

[0 O *1 Downcomer fluid heat up during depressurization events

• Downcomer fluid mixing during safety injection

[- 368Iw-7.non:Im97 7-39 c _ _ _ _ _ _ _ _ _ _ - _ - - _ -

FACILITY SCALING REPORT

                         =   Buoyant fluid backflow in safety injection lines

• Two-phase flow regime transitions Figurc 7-7 is a flow diagram for the downcomer scaling analysis. First, a top-down scaling analysis is performed to obtain the residence times, characteristic time ratios, and specific frequencies important to fluid heat up in the downcomer. Next, a bottom-up scaling analysis is performed to scale specific l processes including fluid mixing, buoyant fluid backflow, and two-phase flow regime transitions.

The following parameters have been scaled to address the downcomer phenomena of interest.

• Downcomer length a Downcomer gap width
• Downcomer stored energy to fluid volume ratio
• DVI diffuser geometry 7.8.1 Top-Down Scaling Analysis of Downcomer Fluid Heat Up During Depressurization The governing equations for the transport process in the downcomer are as follows:
• Liquid Mass O

(p,V,) = A [pQ) (7-151)

• Liquid Energy I (p,e,V,) = A pQ e +[ + goc +* (7-152) dt pm

                          . Solid Structure Energy PsV,Cvs                                                                      (7-153) fs) " 90c Equation 7-153 has already been evaluated in Section 5.0 in terms of an average stored energy release rate during depressurization. Therefore, the liquid mass and energy equations will be the focus of this analysis.

3681w-7.non:lb-o6o897 7-40

r FACILITY SCALING REPORT s The following assumptions will be applied to the governing equations:

1. There is no work done by the liquid.

2. 'Ihe change in kinetic and potential energy internal to the liquid inside the control volume is small compared to its internal energy.

The liquid mass balance equation can be writte.n as follows: i d g (p,V y = rhoyg + scL - SLP g (7-154) l l where:  ! rhoyg = DVI mass flow rate into the downcomer rhcL . = mass flow rate entering the downcomer through the cold leg thtp ' = mass flow rate leaving the downcomer to the lower plenum

     'N                                     Applying the simplifying assumptions to the liquid energy equation yields the following:

(b i (PfeintVs)Dc =(rhh)Dyl +(rhh)ct -(rhh)tp + qoc (7-155) Equations 7-154 and 7-155 are summarized in Table 7-23. 4 As in previous analyses, Equations 7-154 and 7-155 can be expressed in terms of dimensionless E quantities using initial and boundary conditions. Dividing the liquid mass conservation equation by (rhoyg), and the liquid equation by (thh)Dv1,o yields the non-dimensionalized balance equations listed in Table 7-24. Also presented are the characteristic time ratios and specific frequencies for the downcomer processes. 7.8.2 Evaluation of Downcomer Dimensions Because of the physical constraints imposed by the American Society of Mechanical Engineers

                                           '(ASME) pressure vessel code, the wall thickness of the model vessel is set by the operating conditions of the system. Furthermore, the extemal diameter of the core barrel is fixed by the scaling of core flow area and core reflector geometry as presented in Section 4.0. These factors limited the range of downcomer gap sizes that can be selected.

I 4 3681w.7.non:1b-072297 7-41 l

FACILITY SCALING REPORT Section 5.0 presents an evaluation of the average downcomer stored energy release rate during depressurization. Based on that analysis, assuming a minimum reactor vessel wall thickness of 0.5 inches (ASME code), the following result was obtained: (qoc)g" N M b M b Under two-phase fluid natural circulation conditions in the loop, the scaling requirement given by Equation 4-138 must be satisfied. Therefore: (9Dc)g =(ape), poc)% g [g ND where Ygis the fluid property group ratio defined in Section 5.0. The Section 5.0 analysis found Yg to be 2.70 for the APEX operating conditions. Thus, the cross-sectional flow area scaling ratio can be obtained by substituting the result given by Equation 7-156, Yg = 2.70 and (f oc)R = 0.25 into Equation 7-157. Thus: (anc)g = m WW (w g Or in terms of the downcomer annulus gap (soc') (soc)R (*DC)g " N @% The downcomer annulus gap is defined as follows: soc = (D, - DB) (7-170) where: D, = inside diameter of the pressure vessel Da = outside diameter of the core barrel 368Iw-7.non:Ib4)60897 7 42

FACILITY SCALING REPORT p t . The average circumference of the annulus (woc)is given by the following: wDC = (D y +D) g (7-171) Because the outside diameter of the core barrel (D3 ) has already been fixed in the test facility by the core geometry, Equation 7-171 fixes the value for D y. Thus, the downcomer dimensions are fixed. Table 7-25 summarizes the downcomer scaling ratios and downcomer dimensions. 7.8.3 Bottom-up Scaling Analysis of Downcomer Phenomena This section presents scaling analyses for two phenomena:

                                                                  .            Fluid mixing and buoyant backflow

• Two-phase flow pattern transitions Downcomer Fluid Mixing and Buoyant Backflow

(~~T Experimental studies of fluid mixing and buoyant backflow in injection lines have been performed at

         \d                                         various facilities.(15,16) Theofanous and Yan have gathered this experimental data for test facilities

{ ranging from 1/5th scale to full scale.(15) The result of this research is that the onset of buoyant i backflow and fluid mixing at the injector nozzle can be scaled by requiring that (fl ,,9yg)g p equal one in the following equation: {uinj)R b'bVI,R H (7-172) Do,9yi (Pin] ~ PH) Pinj R where: pin; = density of the injected liquid pg. ., density of the hot primary fluid Do,py, " hydraulic diameter of the diffuser I ( \

     ' Q ,Y.

3681w-7.nou:1b.072297 43

FACILITY SCALING REPORT For the rectangular diffuser selected, the hydraul: diameter is defined as follows: DD.Dyl = 5 *} (7-173)

                                                                                              .q l

where: I f, = diffuser length along the axis of the downcomer gap (sDC) f, = diffuser length along the axis of the annulus circumference (woc) Expressing Equation 7-173 in terms of DVI mass flow rate yields the following:

                                                                                               /

(E DVl apyj), ( (UFr)DVI.R  % Do,9yipio)(pinj - PH). R The density of the injected liquid in the model and in the AP600 are identical. Therefore, Equation 7-174 becomes the following:

                                                                                               /

(E DVl apyj)R (U)DV1,R Fr ( n DD,DVI (Pinj ' PH). R Setting (Up,)pyg,g equal to 1 and (spyi)g equal to (pTP,o)R / 96 in Equation 7-175 and requiring geometric similarity fixes the dimensions for f, and f,. Table 7-26 lists the ideal diffuser dimensions as a function of model pressure and the actual DVI diffuser dimensions selected. Two Phase Flow Transition For most slow depressurization transients it is expected that the downcomer region (from DVI center line to the top of lower core support plate) will be filled with saturated or subcooled liquid. During depressurization events, however, some subcooled nucleate boiling may occur on the downcomer wall surface. Because of DVI injection, bubbles detaching from the surface will collapse in the bulk fluid. O 3681w-7.non:Ib-060897 7-44

FACILITY SCALING REPORT

 .,e %

(] For rapid depressurizations with DVI actuation, fluid property similitude exists. Under these conditions, the following requirements need to be met: l l Venical Flow: Bubbly to slug flow:(17) UR*I (7~I7b) j l t Slug to churn flow':(17)  ! oc (7-177) 1 s oc, Slug / churn to annular flow:(17) { v UR*I (7"I70) l For the downcomer, both the bubbly to slug flow transition and slug to churn flow transitions will be reasonably modeled. Forced Convection Heat Transfer.

       - During CMT and accumulator injection, a large quantity of subcooled water is injected into the downcomer. The hot downcomer walls are cooled by convective heat transfer to the subcooled liquid.

A' convective heat transfer coefficient (HDc) can be calculated using the Dittus-Boelter correlation for forced convection heat transfer.(18) That is: t Hoc = 0.023 Ref.8 Pr[A (7-179) 2sDc f3 L) ' t 368iw.7.non:1b-o60897 7-45 I

FACILITY SCALING REPORT Substituting into Newton's law of cooling yields the following: qDC = 0.023 Aoc (I w-T,) Re,"8 Pr[A (7-180) j 2s oc where: t A Dc = downcomer surface area T, = downcomer wall temperature k, = liquid thermal conductivity T, = liquid temperature Soc = downcomer gap and where the characteristic dimension is twice the downcomer gap (Soc)- i The Reynolds number (Re,) is defined as follows: rhoyl Re,= (7-181) WDCMt The Prandtl number (Pr,) is defined as follows: Cp ,p' ' Pr, = (7-182) kg l where: l l th ey = injection mass flow rate p, = liquid viscosity C,p = liqu'd specific heat wDC = average annulus circumference O 3681w-7.non:1b-072297 7 46

FACILITY SCALING REPORT Substituting Equation 7-180 into' Equations 7-163 and 7-167 yields expressions for the heat source

                                 - ratio and the heat transfer specific frequency:

l 0.0115 k,Re['8Prf'# ADC(Iw-T,) h DC , (7-183) s oc (rhh)pyj and 0.0115 k,Ref'8Pr,U'# ADc(Tw -T,) (7-184)

                                                                  @q.DC "

sDCPtV,ha vi 1 7.8.4 Evaluation of Downcomer Processes and Scaling Distortions

                                - To assess the relative importance of the downcomer processes, the fluid residence time constants,
                                , specific frequencies, and characteristics ratios presented in Table 7-24 must be evaluated for both the
        -(p)                    :

APEX model and the full-scale AP600. A comparison of these values reveals which of the processes are dominant and to what extent the proposed APEX geo:netry and operating conditions introduce scaling distortions.

                                 ' Table 7-27 presents the numerical' values for the downcomer fluid residence time, the specific frequencies, the characteristic time ratios and the distortions factors for the downcomer processes.
                                - De maximum injection flow rate occurs during accumulator injection. His occurs at 715 psia in the AP600 and at 230 psia in the APEX facility. During accumulator injection, the flow through the DVI (rhoyg) is approximately equal the flow through the lower plenum (thtp). The balance of the flow goes towards refilling the primary system. Derefore, the cold leg flow rate (rhet)is very low The following results:

LP

                                                                                  =1      and         " =0                             (7-185) my o                   m ayi l

De mass flow rate ratio and energy flow rate ratio are as follows: 4,0c.1 (7-186) O

                                - 368tw.7.non:lh .'50897                                 7-47

FACILITY SCALING REPORT and S.Dc " 1 (7-187) The dominant downcomer transport process during accumulator injection is characterized by the heat source ratio. He DVI fluid mixing time constant is approximately equal in APEX and the AP600. This is a deviation from the one-half time scale requirement applied to the other processes in APEX. The time scale ratio for the downcomer is as follows: l l (toc)g " I @@ Since the time scale ratio is as defined in Equation 7-188 and the downcomer length scale ratio is 0.25, then the fluid velocity ratio is given by the following: (uoc)R " Table 7-27 reveals that the downcomer processes will be reasonably modeled, although on a different time scale. l 7.9 Conclusions This section implemented the H2TS methodology to obtain the scaling criteria for the venting, draining, aad injection processes. The fluid residence time constants, characteristic time ratios, and j specific frequencies for each of the processes for each of the passive safety systems have been l obtained and evaluated at the end of each section. Scaling distortions introduced by the APEX geometry or operating conditions are also highlighted at the end of each section. l Because the safety injection line and vent line resistances can be controlled, the desired scaled mass flow rates can be obtained for all of the passive safety injection and venting systems. The CMT, ACC, and IRWST volumes are shown to be properly scaled. The ADS 1-3 sparger is scaled using low-pressure critical flow data and geometric similarity is preserved. The upper core support plate has been sized by considering counter-current steam-water flow through perforated plates. De reactor vessel downcomer has been sized such that the downcomer stored energy release rate to

 . fluid volume is reasonably scaled. This resulted in increasing the downcomer gap such that the l

1 368Iw 7.non:Ib-060897 7 48

FACILITY SCALLNG REPORT downcomer length to gap ratio is preserved in APEX. Thus, rather than a 1:2 time scale for fluid transport, a 1:1 fluid transport time scale is established in the downcomer. The distortion introduced by increasing the downcomer fluid volume is minimal. The downcomer diffuser nozzle has been sized taking into account the potential for buoyant fluid j backflow at very low injection flow rates. l 1 In conclusion, the scaling analysis presented in this section indicates that the passive safety system venting, draining, and injection process are reasonably scaled and that the data obtained in the APEX facility should be of sufficient quality to validate advanced thermal-hydraulic computer codes. 7.10 References

1. AP600 NOTRUMP Calculations, Westinghouse Electric Corporation, May 26,1992.

1

2. Todreas, N.E., and M.S. Kazimi, Nuclear Systems 1: Thermal Hydraulic Fundamentals, Hemisphere Publishing Corporation,1990. _

3. . Fauske, H.K., " Flashing Flows or: Some Practical Guidelines for Emergency Releases,"

Plant / Operations Progress, Vol. 4, No. 3, July 1985. l [ ' 4. Hsu,' Yih-Yun, and R.W. Graham, " Transport Processes in Boiling and Two-Phase Systems," Section 11.4.2.1, Fauske's Theory of Critical Flow, Hemisphere Publishing,1976.

5. . Taitel. Y., and A.E. Dukler, "A Model for Predicting Flow Regime Transitions in Horizontal and Near Horizontal Gas-Liquid Flow," AIChE Journal, Vol. 22, No.1, pp 47-54, January 1976.

6. Encyclopedia of Fluid Mixing.

7. Levy, S., " Prediction of Two-Phase Critical Flow Rate," ASME Journal of Heat Transfer, 64-HT-8,1964.

             ' 8.       Dilber, J., Counter-Current Steam / Water Flow Above a Perforated Plate-Vertical injection of Water, U.S. Nuclear Regulator Commission, Washington, DC 20555, NUREG/CR-4415, October 1985.

9. Dilber, I., S.G. Bankoff, R.S. Tankin, and M.C. Yuen, Counter Current Steam /MSter Flow Above a Perfurated Plate VerticalInjection of Water, U.S. Nuclear Regulator Commission, Washington, DC 20555, NUREG/CR-2323, September 1981.

10. Liu, C.P., C.L. Tien, and G.E. McCarthy, Flooding in Vertical Gas-Liquid Countercurrent Flow Through Parallel Paths EPRI NP-2262, Electric Power Research Institute, Palo Alto, CA 94304, February 1982.

11. Block, J.A., and G.B. Wallis, Effects of Hot Walls on Flow in a Simulated PWR Downcomer V . During a LOCA, U.S. Atomic Energy Commission (USNRC), Washington, DC 20555, CREARE-TN-188, May 1974.

3681. 7.non:tb-060897 7-49

FACILITY SCALING REPORT

12. " Upper Plenum Test Facility Downcomer ECC Bypass Tests," MPR Associates presentation to the USNRC Downcomer Bypass Meeting, Rockville, MD, July 13,1987.

13. Rothe, P.H., Technical Summary Attachment to ECC Bypass RIL, Volume I; Review of Findings, NUREG/CR-0885, Vol.1, July 1979.

14. Crowley, C.M., et al., Technical Summary Attachment to ECC Bypass RIL, Volume 11; Technical Appendices, NUREG/CR-0885, Vol. 2., July 1979.

15. Theofanous, T.G., and H. Yan, A Unified Interpretation of One-fifth to Full Scale Thermal Mixing Experiments Related to Pressurized Thermal Shock, U.S. Nuclear Regulatory Commission, Washington, DC 20555, NUREG/CR-5677, April 1991.

16. King. J.B., Jr., and J.N. Reyes, Jr., " Buoyant Backflow in Vertical Injection Lines," Proceedings of the 1991 International Meeting of Safety of Thermal Reactors, Portland, OR, July 1991.

17. Schwartzbeck, R.K., and G. Kocamustafaogullari, " Similarity Requirements for Two-Phase Flow Pattern Transitions," Nuclear Engineering Design,116, pp 135-147,1989.

18. Welty, Jr., C.E. Wicks, and R.E. Wilson, fundamentals of Momentum Heat and Mass Transfer, John Wiley and Sons, New York,1984.

O O

O TABLE 71 TOP DOWN SUBSYSTEM LEVEL SCALING BALANCE EQUATIONS l FOR SAFETY INJECTION SYSTFMS l i Liquid Mass: (p,V,) = A[pQ] (7-18) Liquid Energy: (p,(e,)V,) = A pQ e, + + q, - w, (7-19) where: w, = shaft work rate done by the liquid e, = average liquid energy per unit mass O TABLE 7-2 TOP.DOWN SUBSYSTEM LEVEL SCALING ANALYSIS: CONTROL VOLUME BALANCE EQUATIONS FOR SAFETY INJECTION TANK DRAINING (WITII SIMPLIFYING ASSUMPTIONS) Liquid Mass: PtDANK (bt ) " b~l $2 (7-24) Liquid Energy: p,aTANK (Cygfg)L )g = E(sh)3 -h2 Eb 2+9: (7-25) I O 3681w-7.non:Ib-060897 7 51

TABLE 7-3 SET OF INITIAL AND BOUNDARY CONDITIONS l l USED TO NON.DIMENSIONALIZE THE SAFETY INJECTION TANK BALANCE EQUATIONS I Eth; =(Eth g)0 I Ethy = {Eth 2)0 2 L=LL* o Cy ,f,) =(Cy,f,))gC,*,f,)* E(thh)g = E(thh)i,oE(thh)* E(rhh)2 = E(thh)2.0E(thh)2 91 " 91,0 91 0 368Iw-7.non:ltM97 7-52

s FACILITY SCALING REPORT TABLE 7-4 NON DIMENSIONALIZED BALANCE EQUATIONS FOR SAFETY INJECTION TANK DRAINING Liquid Mass: To L, = I( Erh 3 *-Es + 2 (7-26) Liquid Energy: Ley

                                                                   *C *,f,) = q E(rhh)3 -E(sh)2

• Ik9t (7-27)

TIME CONSTANT, SOLID SPECIFIC FREQUENCY, AND CHARACTERISTIC TIME RATIOS Drain Rate Time Constant: Pf8TANKL, tg- _ N 2 ,0 Characteristic Time Ratios: O Ik = gE+,' (7-29) r 2>0 Mass Flow Rate Ratio:

                                                                                  'E(sh)3 Il' =                                          (7-30) s E(sh)2 so Energy Flow Rate Ratio:

U= q (7-31) s E(sh)2 JO Heat Loss Ratio: Property Ratio: h2 Ye = (7-32) , Cy ,f,)_ l Specific Heat Ratio: O 368Iw-7.non:Ib-060897 7 53

FACILITY SCALING REPORT l i TABLE 7-5 NON.DIMENSIONALIZED BALANCE EQUATIONS FOR ACCUMULATOR INJECTION Liquid Mass: , T ACC " ~b tACC (Ye (7-34) Liquid Energy: ACC y,+Cy *, T,' = -(thh)[ACC 7ACC TIME CONSTANT AND CHARACTERISTIC TBIE RATIO Injecdon Rate Time Constant: PfVf (7-36) O TACC " r b < ACC,o Property Ratio: I r h' SPecific Heat Ratio 7 ACC * (7-37) 4 CVI fI). ACC,o l

                                                                                                                                                        )

O 3681w-7.non:Ib-060897 7 54

FACILITY SCALING REPORT 1 l v TABLE 7-6 IdODEL ACCUMULATOR SCALING RATIOS AND DIMENSIONS l g APEX Model Parameter Ratio AP600 Ideal Actual (TBM) Internal Volume 2000, 3 3 1:192 ft 10.42 ft 10.42 ft 3 3 56.64 m 0.295m 3 0.295m 3 < ( Initial N2Volume 1:192 300. ft 3 1.56 ft 3 1.56 ft 3 3 3 8.49 m 0.044m 0.044m 3

                                                                                          ,       Initial Liquid Volume      1:192   1700.      ft 3       8.85 ft 3                                      8.85 ft 3

3 3 48.15 m 0.251m 0.251m 3 Initial Liquid -- 120. *F 120. *F 100. 'F Temperature 48.9 *C 48.9 *C 37.8 *C Initial N2Charging 1:3.086 714.69 psia 231.7 psia 231.7 psia Pressure 4.93 MPa 1.60 MPa 1.60 MPa N 2Pressure at Liquid 1:3.086 50.19 psia 16.27 psia 16.27 psia Empty Condition 0.35 MPa 0.11 MPa 0.11 MPa t"N g Mass of N2 in 1:192 40.0 lbm 0.2081bm 0.1081bm Solution 18.18 kg 0.095kg 0.049kg Mass of N2Injected 1:192 566.6 lbm 2.95 lbm 0.140lbm 257.5 kg 1.34 kg 0.063kg Supplemental N2 Mass --- --- --- 2.91 lbm Injected From Flask 1.32 kg

• Total N2 Mass Injected 1:192 606.6 lbm 3.16 lbm 3.16 lbm into Primary 275.7 kg 1.44 kg 1.44 kg i

TBM- to be measured. Numerical values may not reflect as-built conditions. Refer to the APEX Facility Description Report.

• Sum of N2 injected, N2in solution and supplemental N2-j n

3681w.7.non:lt460897 7-55

FACILTlY SCALING REPORT TABLE 7-7 MODEL ACCUMULATOR SCALING RATIOS AND DBIENSIONS TIIAT SATISFY THE TRANSITION PRESSURE REQUIREMENT APEX Model Parameter Ratio AP600 Ideal Actual (TBM) Intemal Volume 1:192 2000. ft 3 10.42 ft 3 10.42 ft3 56.64 m 3 0.295 m 3 0.295m 3 Initial N2 Volume 1:192 300. ft3 1.56 ft 3 4.47 ft3 8.49 m 3 0.044 m 3 0.127m 3 Initial Liquid Volume 1:192 1700. ft3 8.85 ft 3 8.85 ft 3 48.15 m 3 0.251 m 3 0.251m 3 Initial Liquid -- 120. *F 120. *F 100. *F Temperature 48.9 *C 48.9 *C 37.8 *C Initial N2Charging 1:2.625 714.69 psia 231.7 psia 231.7 psia Pressure 4.93 MPa 1.60 MPa 1.60 MPa N2Pressure at Liquid 1:1 50.19 psia 19.12 psia 50.19 psia Empty Condition 0.35 MPa 0.13 MPa 0.35 MPa Mass of N2 in 1:192 40.0 lbm 0.208 lbm 0.1081bm Solution 18.18 kg 0.095 kg 0.049kg Mass of N2 Injected 1:192 566.6 lbm 2.95 lbm 3.33 lbm 257.5 kg 1.34 kg 1.51 kg l

• Total N2 Mass Into 1:192 606.6 lbm 3.16 lbm 3.44 lbm i

Primary 275.7 kg 1.44 kg 1.56 kg TBM- to be measured. Numerical values may not reflect as-built conditions. Refer to the APEX Facility Description Report. !

• Sum of N2 injected and N2 in solution.

l TABLE 7-8 I ACCUMULATOR TIME CONSTANTS, RESIDENCE TIME RATIOS, AND PROPERTY RATIOS AP600 APEX Fluid Residence Time Constant (ta ce): 352.5 s 176.2 s Specific Heat Ratio (ygcc); 1.0 1.0 Residence Time Ratio (TAcc/Tsys); 2-inch break 0.08 0.08 j DEDVI break 1.9 1.9 l 3681w-7.non:1b-o60897 7-56

i FACILTTY SCALING REPORT 1 r l l' TABLE 7-9 l NON DmfENSIONALIZED BALANCE EQUATIONS FOR IRWST INJECTION Liquid Mass: T IRWST bl "kIRWST (BADS 1-3 ec +b ) ~ b1RWST l Liquid Energy: e 8 TIRWST d YIRWST g hi *C +/ vf(T,A[ = %IRWST ((sh) ADS 1-3 4sh)cc]+ (7-46) 1 N)IRWST h.!RWST9 PRHR TIME CONSTANT AND CHARACTERISTIC TBIE RATIOS l Drain Rate Time Constant: i I TIRWST " (7-47) m. IRWST. o

   ,,3 ty.      Characteristic Time Ratios:

b ADS 1-3 N ec . k!RWST = Mass Flow Rate Ratio (7-48) b lRWST

                                                  .o (sh) ADS I-3 +(sh)cc Uh.IRWST =                                                                                                                             Energy Flow Rate Ratio                    (7-49)

(sh)IRWST .o h.IRWST = Heat Source Ratio (7-50) (sh)IRWST o Property Ratio: h, YrRWST " SPecific Heat Ratio (7-51) l

        -                        VI   I).IRWST.o u

l l 368Iw-7.non:1b@0897 7 57 a

i I FACILITY SCALING REPORT TABLE 7-9 (continued) IRWST PROCESS SPECIFIC FREQUENCIES ( 0ADSl-3 + b ec j Om,IRWST " (7-52) [ pvt t .o (sh)ADSl-3 +(sh)cc (7-53) klRWST - h1RwsT P V, ,9 1 9FRHR S q.IRWST * (7-54) h!RWST PtV, / 1 l O 368Iw.7.non:Ib-060897 7,

FACILITY SCALING REPORT TABLE 7-10 MODEL IRWST SCALING RATIO AND DIMENSIONS ) i Scaling Parameter Ratio AP600 Ideal Actual (TBM) , a,C i Normal Liquid 1:192 Volume Liquid Surface Area 1 '.8 Normal Water Depth 1:4 Model Tank Diameter --- Initial Water --- Temperature Tank Pressure * --- f'h.

 .Q                                                                       Condensate Retum Rate

• TBM- to be measured. Numerical values may not reflect as-built conditions. Refer to the APEX Facility Descr;ption Report.
• Varied parametrically

(")h

 \_

l 368Iw-7.non:Ib-072297 7-59

FACILITY SCALING REPORT TABLE 711 IRWST TBIE CONSTANTS, PROCESS SPECIFIC FREQUENCIES, CIIARACTERISTIC Tm1E RATIOS, AND DISTORTION FACTORS AP600 APEX Residence Time Constants (t1RWST): 2.29x104 s 4 1.145x10 s Process Specific Frequencies: 4 -1 Com.!RWST: 1.77x10 4 s'I 3.88x10 s 5.45x10 4 s'I 4 ~l Wh.IRWST: 8.83x10 s (ca.IRWST: 4.59x10-5 s 9.17x10 4 s'l Characteristic Time Ratios: U m.IRWST: 4.06 4.45 U h.IRWST: 12.51 10.12 H a.IRWsT: 1.05 1.05 Specific Heat Ratio: YIRWST: 1.0 1.0 Distortion Factors (DF): (DF)m.IRWST: -9.5% (DF)h.IRWST: 19.1 % (DF)a 1RWST: 0% 9 368Iw-7.non.Ib.060897 7 60 _ _ _ _ _ _ _ _ _ _ - - _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _1

FACILITY SCALING REPORT ) i l TABLE 7-12 CONTROL VOLUME BALANCE EQUATIONS FOR THE i* SECTION OF A SAFETY INJECTION LINE j l Liquid Mass: (p,V,); = A[pQ]; (7-60) Liquid Energy: 1 (Pset Vg )i " A PQ e+ + 9e - ws (7-61) ( )

 'V                                                                                                                                              )

TABLE 7-13 I SCALING RATIOS FOR SAFETY INJECTION LINE RESISTANCE CMT Injection Lines:

                                                                     '2

[T,iNhM T.R (9 L T(t)+L V,CMT.R 4 TP,o , CMT,R ACC Injection Lines: r '2 l [T,idkCC,R "(9 ACC EDVI)R ( 0 s ACC,R IRWST Injection Lines: fT,idhRWST,R . inj L t) +Ly

                                                                                    .IRWST R O

3681w-7.non:1b-060897 7-61

FACILITY SCALING REPORT TABLE 7-14 Ol l MODEL SAFETY INJECTION LINE SCALING RATIOS AND DIMENSIONS Ideal APEX Model Scaling Parameter Ratio AP600 Ideal Actual (TBM) CMT Injection Line 0.1443 6.81 in 0.983 in 1.06 in Inside Diameter 17.3 cm 2.50 cm 2.59 cm

                                                                                                                                                                                                -                                                                                                                                          - a,c CMT Injection Line                                                                                                                  1:192 Volume                                                                                                                                                    _

CMT Vertical Elev. 1:4 28.86 ft 7.215 ft 7.215 ft [Ly + L 7(t=0)] 8.80 m 2.20 m 2.20 m _ - a,c CMT Maximum 1:96 Injection Flow Rate - - l CMT (F 7,;n;) 1:1 24.18 24.18 32.71** l a,c CMT Injection Line 1:4 O Loss =(F 7 ,io;)p."ini 2 l ACC Injection Line 0.1443 l Inside Diameter ACC Injection Line 1:192 Volume j ACC Maximum 1:96 Injection Flow Rate ACC (F 7,ini) 1:1 IRWST #1-Sump Tee 0.1443 ) Inside Diameter IRWST #1-Sump Tee 1:192 ) Volume IRWST #1 Vertical 1:4 j Elev. [Ly +L7(t=0)] _ _ l l O I 3681w-7.non:Ib472297 7 62

- _ _ _ _ _ _ _ _ _ _ _ _ .   . _ _ _ _ _ _ _ _ _ _ _ _ = _ _ _ _ _

l FACILITY SCALING REPORT O TABLE 714 (continued) MODEL SAFETY INJECTION LINE SCALING l RATIOS AND DIMENSIONS Ideal APEX Model Scaling Parameter Ratio AP600 - Ideal Actual (TBM) l

                                                                                                                                                                          -                a,e i IRWST #1 Maximum                          1:96 Injection Flow Rate IRWST #1-Sump Tee                         1:1 (F7,in;)

i IRWST #1-Sump Tee 1:4 2 APtos,=(FT ,inj)P! I"I 2 IRWST #1-Sump-DVI 0.1443 Inside Diameter

                  ,, .s IRWST #1-Sump-DVI                         1:192
           .g                 Volume IRWST #1-Sump-DVI                         1:1 (F73n;)                                                                                                                                                          !

IRWST #1-Sump-DVI 1:4 2 l APtos,=(F7,ioj)p3Sni 2 IRWST #2-Sump Tee 0,1443 J Inside Diameter I l IRWST #2-Sump Tee 1:192 Volume IRWST #2 Vertical 1:4 Elev. [Ly+LT (t=0)] IRWST #2 Maximum 1:96' Injection Flow Rate . 1 IRWST #2-Sump Tee 1:1 (F7,in;) - - 3681w-7.non:1b-072297 7-63 f

FACILITY SCALLNG REPORT O TABLE 7-14 (continued) MODEL SAFETY INJECTION LINE SCALING RATIOS AND DIMENSIONS Ideal APEX Model Scaling Parameter Ratio AP600 Ideal Actual (TBM) r- - a,c IRWST #2 Sump-DVI 1:4 2 APtos,=(F7,;,j)p,"ini 2 IRWST #2-Sump-DVI 0.1443 Inside Diameter IRWST #2-Sump-DVI 1:192 Volume IRWST #2-Sump-DVI 1:1

                                               @T,ini)

IRWST #2-Sump-DVI 1:4 APto,,=(FT,inj)P! i"I 2

                                                                                                                                                            =

9 TBM- to be measured. Numerical values may not reflect as-built conditions. Refer to the APEX Facility Description Report.

• Based on NOTRUMP calculation.W

                                      ** Adjusted using orifice to satisfy Equations 7-88 through 7-90.

l O 36stw.7.nomm om97 7-64

7 i FACILITY SCALING REPORT l h 4

 \)

TABLE 7-15 DVI LINE SCALING RATIOS AND DIMENSIONS Ideal APEX Model Scaling Parameter Ratio AP600 Ideal Actual (TBM) DVI Line Inside 0.1443 - a,e Diameter DVI Line Volume 1:192 DVI Maximum 1:96 Injection Flow Rate (F7,oy,) 1:1 DVI Line 1:4 U.2, AP to3,=(F 7 ,9yi)p, { _ i fq TBM- to be measured. Numerical values may not reflect as-built conditions. Refer to the APEX

 'Q        Facility Description Report.

• Based on NOTRUMP calculation.(U

           ** Adjusted using orifice to satisfy Equation 7-97.

TABLE 716 8 CONTROL VOLUME BALANCE EQUATIONS FOR THE 1 ' SECTION OF A VENT LINE (TWO-PHASE HOMOGENEOUS MIXTURE) Fluid Mass: i (PfV g), = A[pQ), (7-98) l Fluid Momentum: l h . dt (Pr f g)8u V . = A[puQ) n +f a (APV) -(pf fgz), (7-99)

 .p                                                                                                                                          + (M,i + Mgn]a v

l 3681w-7.non:Ib-072297 7-65 L_.___._

FACILITY SCALING REPORT TABLE 717 SCALING RATIOS FOR CMT BALANCE LINES AND ADS VENT LINES CMT Balance Lines:

                                                                                                                                                                                                          ,2
                                                                                                                                                                                                      ^

[T,BAL)R (96 [PhAPCRIThAL,R s PTP,o, ADS 1-3 Vent Lines:

                                                                                                                                                                                                                ,2
                                                                                                                                                                                                       ^

Single: [FT, ADS 14)R (9b fPb APCRITkDS14,R

                                                                                                                                                                                                                   ,2 AD I Combined: fT, ADS 14)R =(48f                                   (PhAPCRITkDSid,R p

ADS 4 Vent Lines:

                                                                                                                                                                                                          ,2 ADS 4

[FT. ADS 4)p =(96) ( 3pCRITkDS4,R

                                                                                                                                                                                               , P n . ,. ,   -

0 368Iw.7.non-Ib-060897 7,66

FACILITY SCALING REPORT TABLE 718 BA. LANCE LINE SCALING RATIOS AND DIMENSIONS { l Ideal APEX Scaling Parameter Ratio AP600 Ideal Actual (TBM) { CUCMT Balance Lines 0.1443 a,c Inside Diameter CUCMT Balance Lines 1:192 I Volume CUCMT Maximum 1:96 Flow Rate l CUCMT (F 7,3,3) 1:1 CUCMT (APCRIT) 1:4 TBM- to be measured. Numerical values may not reflect as-built conditions. Refer to the APEX Facility Description Report. f N

• Based on NOTRUMP calculation.II)

                                           ** Adjusted using orifice to satisfy Equation 7-120.

I 1 1 l l l l l l O 3681w-7.non:lb472297 7-67

FACILITY SCALING REPORT O TABLE 7-19 O ADS 13 VENT LINE SINGLE AND COMBINED TRAIN DIMENSIONS AP600 APEX (TBM) Parameter Single Tier Single Tier Combined Tier 2 ,c PZR to ADS Header: Inside Diameter PZR to ADS Header: Maximum Flow Rate PZR to ADS Header: F7(Two-Phase) PZR to ADS Header: l (APcarr) _ l 1" Stage ADS Line: Inside Diameter 1" Stage ADS Line: Maximum Flow Rate 1" Stage ADS Line: l F7(Two-Phase) l 1" Stage ADS Line: l (APcarr) 2"d or 3'd Stage ADS Line: Inside Diameter 2"d or 3'd Stage ADS Line: Max. Flow Rate 2"d or 3'd Stage ADS Line: F7(Two-Phase) 2"d or 3'd Stage ADS l Line: (APcRrT) l ADS Sparger Line Inside Diameter l ADS Sparger Line l Maximum Flow Rate l ADS Sparger Line F7(Two-Phase) ADS Sparger Line (APcRrr) _ _. EBM- to be measured. Numerical values may not reflect as-built conditions. Refer to the APEX Facility Description Report.

• Based on NOTRUMP calculation.m (1) Adjusted using orifice to satisfy Equation 7-121. .

1 (2) Adjusted using orifice to satisfy Equation 7-122. 3681w.7.non:lb472297 7 68 l

FACILITY SCALING REPORT A 4 i i

   \_./                                                                                                                                                                     )

TABLE 7-20 ADS 4 LINE SCALING RATIOS AND DIMENSIONS 4 l Ideal APEX 1 Scaling Parameter Ratio AP600 Ideal Actual (TBM)

                                                                                    ~
                                                                                                                                                                    ~

ADS 4*-Stage Line 1:96 a,c Maximum Flow Rate l ( ADS 4* Section 1 0.1443 Inside Diameter f ADS 4* Section 1 1:1 F7(Two-Phase) ADS 4* Section 1 1:4 (A Pcgi7) ADS 4* Section 2 0.1443 Inside Diameter ADS 4* Section 2 1:1 F7(Two-Phase)

  ?

V ADS 4* Section 2 1:4 (A PcRrr) l ADS 46Section 3 0.1443 Inside Diameter l ADS 4* Section 3 1:1 F7(Two-Phase) ADS 4* Section 3 1:4 (A PcRrr) _ TBM- to be measured. Numerical values may not reflect as-built conditions. Refer to the APEX Facility Description Report. l

• Based on NOTRUMP calculation.(D

         ** Adjusted using orifice to satisfy Equation 7-123.

i > - (N U 36siw.7mn:1b-072297 7-69 t-L-

FACILITY SCALING REPORT O TABLE 7-21 ADS 13 SPARGER SCALING RATIOS AND DIMENSIONS Scaling AP600 APEX Model Parameters Ratio (2 Spargers) (Combined Sparger) a,c Hole Diameter 1:2 Number of Holes Per Arm 1:3.65 Sparger Arm Wall Thickness Hole Wall Thickness / Diameter Ratio 1:1 Number of Arms 1:2 Total number of Holes 1:7.3 Branch Orientation 1:1 Single Sparger Line I.D. Single Sparger Hub I.D. Single Sparger Arm I.D. Sparger Line/ Hub Area Ratio 1:1

• Sparger Hub / Arm Area Ratio 1:1 ._ _

Note: Numerical values may not reflect as-built cr.nditions. Refer to the APEX Facility Description Report. l-I O 368iw.7.non:1b-072297 7 70

FACILITY SCALING REPORT j O b

                                                                                                                                                                     =

TABLE 7-22 UPPER CORE SUPPORT PLATE PERFORATION SCALING RATIO AND DIMENSIONS Ideal APEX Model Scaling Parameter Ratio AP600 Ideal Actual (TBM) a,c

                                                                                                                                                                   ~
                                                                                                                                                                                                                                                       ~

Hole Diameter 1:4 Number of Drain Holes 1:3 1 Drain Hole Flow Area 1:48 '

• Additional Flow Area 1:48 Total Flow Area 1:48 )

l

 ,rx                                                                                                                                                                                                                                                          \

TBM- to be measured. Numerical values may not reflect as-built conditions. Refer to the APEX Facility Description Report.

                                                   '

• Additional flow area is modeled using four 0.38-in ,(2.11-cm) diameter holes.

TABLE 7 23 CONTROL VOLUME BALANCE EQUATIONS FOR THE REACTOR VESSEL DOWNCOMER Liquid Mass: (p,V,y = rhovi +$cL -bLP

                                                                                                                                                                                                                                                     ~
                                                                                                                                                                                                                                                        )

Liquid Energy: i (Pt e:V,y = (rhh)gyg frbh)ct drhh(P Moc @N C\ L) 368Iw.7.non:1b-072297 7-71

l FACILITY SCALLNG REPORT TABLE 7-24 NON DIMENSIONALIZED BALANCE EQUATIONS FOR DOWNCOMER LIQUID TRANSPORT PROCESSES Liquid Mass: f T Pr*V,1=sDv1

                                                                                                          *                             + U ,DC m                                          b CL                             -                     (7-158)

DC LP, Liquid Energy:

                                                                                                'e PsintV+1=

t (sh)py, + D h.Dc (sh)* -(sh)*p TDC (7-159)

                                                                                                                                   +1   Dc 9 TIME CONSTANT AND CHARACTERISTIC TIME RATIOS DVI Liquid Mixing Time Constant:

El 8 (7-160) Tg= m ovl. Dc.o Characteristic Time Ratios: ( Dm.Dc = Mass Flow Rate Ratio (7-161) mDVI DC,o l (sh)ct -(sh)tp II,Dc = Energy Flow Rate Ratio (7-162) (sh)Dv1 oco l Heat Source Ratio (7-163) l IkDc = l (sh)Dvt .o Property Ratio: h

                                        '/DC "                                                                                      SPecific Heat Energy Ratio                                                                                   (7-164)     >

C int i i DC,o I 3681w 7.non:lt4)60897 7-72

FACILITY SCALING REPORT I

    'O                                                                                                                                                                  l TABLE 7 24 (continued)                                                                                       '

NON DIMENSIONALIZED BALANCE EQUATIONS FOR DOWNCOMER LIQUID TRANSPORT PROCESSES Process Specific Frequencies: co - CL ~ LP (7 165) PV  % (sh)ct -(sh[p , P V ehovt oc,o

                                                                    =      9DC co                                                                                               (7-167)

, EfV,hovi, o l i i r . u% 4 l i

    ,rr 3681w.7.nortibM897                               7-73 E______._____.________________ __ __        _

FACILITY SCALING REPORT TABLE 7-25 DOWNCOMER SCALING RATIOS AND DIMENSIONS Ideal APEX Model Scaling Parameter Ratio AP600 Ideal Actual (TBM) Outside Diameter of 0.1443 137.75 in 19.88 in 20.0 in Core Barrel (Dg) 349.89 cm 50.49 cm 50.8 cm Inside Diameter of RPV *0.1637 157.00 in 25.70 in 25.0 in (Dy) 398.78 cm 65.28 cm 63.5 cm Average Annulus - 463.00 in 71.59 in 70.69 in Circumference (woc) 1,176.0 cm 181.86 cm 179.54 cm Annulus Gap (sDc) -- 9.625in 2.59 in 2.50 in 24.45 cm 6.57 cm 6.35 cm Downcomer Length 0.25 208.8 in 52.2 in 52.2 in 530.4 cm 132.6 cm 132.6 cm (foc) Downcomer Flow Area *0.046 4,456.3 in 2 204.9 in 2 176.7 in 2 (aoc)g 28,750.3 cm 2 1,322.5 cm 2 1,140.1 cm 2 Downcomer Length to 1:1 21.69 21.69 20.88 Gap Ratio (t/s)pc l l TBM- to be measured. Numerical values may not reflect as-built conditions. Refer to the APEX Facility Description Report.

• Based on two-phase natural circulation requirements.

1 i e 368Iw.7.non:Ib-060897 7 74

                                                                                                                                                                                                                                                     .-                _-___a

FACILITY SCALING REPORT O I TABLE 7-26 DVI DIFFUSER DIMENSIONS APEX APEX Model - (Ideal Range) Model Parameters AP600 (1) 400 psia (2) 230 psia (3) 50 psia Actual Length (f,) at Diffuser a,c Exit L.ength (e,) at Diffuser Exit DVI Line I.D. at Diffuser Inlet (Dn vi) Hydraulic Diameter at Diffuser Exit (Dg .oyi) Flow Area at Diffuser Exit (agvi) O Diffuser Angle Outward from Vertical - - Note: Numerical values may not irflect as-built conditions. Refer to the APEX Facility Description Report. (1) CMT injection starts. (2) Accumulator injection starts. (3) Fluid property similitude. O . 3681w-7.non:lb-o72397 7-75

FACILITY SCALING REPORT i TABLE 7 27 EVALUATION OF DOWNCOMER FLUID RESIDENCE TIMES, CHARACTERISTICS TIME RATIOS, AND SPECIFIC FREQUENCIES APEX AP600 Downcomer Fluid Residence Time: Toc,o 53.0 s 55.6 s Downcomer Process Specific Frequency: to,pc* q 18.2 s'3 14.5 s*I

                                                                                                                       %,DC                                                                                      0.019 s'I             0.019 s'I (oh.DC                                                                                    0.019 s'I             0.019 s'I                            i Characteristic Time Ratio:

U q.Dc 966.8 805.4 Um ~1 ~1 U h.DC ~I 'l Distortion Factors (DF): Heat Source Ratio (Da nc): -16.7%

• Dominant transport process. .

1 I l I i O l 3681w-7.non:lb-060897 7 76 l

FACILITY SCALING REPORT 4

• 3 '

H1 F b 1 r7 I I

                                                   ,___                          =______i           i_____________,

l l I 1 SAFETY TANK CONTROL VOLUME 8 l 'l 8 l ' I I

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t_____________., ~___~____~____~__J c l i L_J w h lf M 2 Le Figure 7-1 Control Volume Boundaries for Safety Tank Draining Analysis N l l 3681w.7.non:lW97 7 77 i

FACILITY SCALING REPORT e ACCUMULATOR INJECTION PHENOMENA I 1 1 Bottom Up Top Down Scaling Analysis Scaling Analysis

                                                                                   . Charging Gas Carryover

• ACC !sjection Rate iP iP Accums1 or Accumulator Injection g'N' , , ---* IIOroups and Scaling S milarity Criteria Evaluate Scaling Distortions Accumulator Design Specifications Figure 7-2 Flow Diagram for the Accumulator Scaling Analysis O

f 1 I 7,73 36si 7.non:ts.o60897 l 1

FACILITY SCALING REPORT O [ IRWST DRAINING PHENOMENA l 1 1 Top Down Bottom Up Scaling Analysis Scaling Analysis IRWST Draining Rate Containment Condensate Return Rate

• Containment Back Pressure IR9rST Liquid lieat-up Rate IRWST IRWST Draining injection - II Grc,sps and une Scalleg Similarity Criteria v

Eralante Scaling Distortions v IRWST Design Specifications Figure 7 3 Scaling Analysis Flow Diagram for the IRWST Draining Prr. css O 3681w 7.non:Ib 060897 7-79

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FACILITY SCALING REPORT e P,, N h 7- ~

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                                                                   ~~              L, . vertest teactas or sapeten tm, L, . ,eormentsi tengt%. or impcten Leo
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INSC7IGN

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                              ,         - TANK                     ..               P = Stetc PressW at DV1 tme t::::::::      ~:1---         ::::::::::::::a M, ,P,
                                                !, i!
                                                !!                                 SAFCTY 9
                                                    ,l i

L* g'-___...._... g INJECTION '

                                                                            -      LINE i

i-L ,, n,l, j I I, U Hw ,Pw_ f DVI LINE j Figure 7-4 Control Volume for Safety Injection Line (Actual piping will have various geometries and fittings) O 3681w-7.non:lb-060897 7 80

l l l CMT BALANCE AND ADS VENT LINE FLOW PHENOMENA i 1 1 Top.Down B ottom-Up Scaling Analysis Scaling Analysis

• Two-Phase Finid Mass Flow
• Choked Flow Rate
• ADS Sparger Steam Uquid lateraction
• Two-Phase Flow Pattern Transitions Chapter 5 Depressurization Balance and Vent une

                        $1milarity     ----.             IIGroups and Criteria                      Similarity Criteria Evaluate Scaling Distortions CMT Balance une and ADS Vent Line Design Specincations Figure 7 5 Scaling Analysis Flow Diagram for CMT Balance Lines and ADS Vent Lines 3681w.7.non:lt4)60897                                     7 81 l

FACILITY SCALING REPORT O l l l l High Pressure Vent Line System ,, r-

                                                                                                                         --- --- g,i 6___.---    _-        'ith section, li     Control Volume ll kI~_~_ _ _ _ _] +3 ~

Two-Phase a p,, > Mix ture s Figure 7-6 Control Volume for Sections of a Vent Line or Balance Line (Actual geometries will j vary) 368Iw-7.rmItM)60897 7 82

1 FACILITY SCALING REPORT . I

  /~T                                                                                                             I l

REACTOR VESSEL DOWNCOMER PHENOMENA l 1 1 Top-Down/Sebsystem Bcttom-up/ Process Scaling Scallag

                      .       Downcomer Fisid Heatap

• Downcomer Fleid Mixing During Depressurization a Buoyant Fluid Backflow
• Two-Phase Flow Regimes N i, 4, Chapter 5

                        . Depressurization    %

DowncomerII0roup j Scaling Analysis and Similarity Criteria r 1P Evalsate Scaling J Distortions l Downcomer and DVI Diffuser Design Specifications t l w Figure 7-7 Scaling Analysis Flow Diagram for Downcomer Phenomena I 3681w-7.nostb-060897 7-83 l

I FACILITY SCALING REPORT

                                            ~

8.0 LCS RECIRCULATION COOLING SCALING ANALYSIS His section presents the similarity criteria that must be satisfied to best simulate'the phenomena I

identified by the PPIRT as being important to lower containment sump (LCS) recirculation and heat L-transfer. The analyses presente}}