ML20093N069
| ML20093N069 | |
| Person / Time | |
|---|---|
| Site: | 05200003 |
| Issue date: | 09/30/1995 |
| From: | WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML20093N029 | List: |
| References | |
| WCAP-14293, WCAP-14293-R01, WCAP-14293-R1, NUDOCS 9510270126 | |
| Download: ML20093N069 (703) | |
Text
{{#Wiki_filter:.-. , . - - - .. .. _ _ - . . . -. WESTINGHOUSE NoN PROPRIETARY CLASS 3 1 f V WCAP-14293 AP600 LOW-PRESSURE INTEGRAL SYSTEMS TEST AT OREGON STATE UNIVERSITY TEST ANALYSIS REPORT SEPTEMBER 1995 i a 4 i l l WESTINGHOUSE ELECTRIC CORPORATION Energy Systems Business Unit Advanced Technology Business Area i P.O. Box 355 )
- Pittsburgh, Pennsylvania 15230-0355
]
01995 Westinghouse Electric Corporation j All Rights Reserved 1 l l O m:hp60023W.non:lbl00395 REVISION: 1 9510270126 951020 PDR A ADOCK 05200003
.PDR
COPYRIGIIT NOTICE The reports transmitted herewith each bear a Westinghouse copyright notice. The NRC is permitted to make the number of copies of the information contained in these reports which are necessary for its internal use in connection with generic and plant-specific reviews and approvals as well as the issuance, denial, amendment, transfer, renewal, modification, suspension, revocation, or violation of a license, permit, order, or regulation subject to the requirements of 10 CFR 2.790 regarding restrictions on public disclosure to the extent such information has been identified as proprietary by Westinghouse, copyright protection notwithstanding. With respect to the non-proprietary versions of these repons, the NRC is permitted to make the number of copies beyond those necessary for its internal use which are necessary in order to have one copy available for public viewing in the appropriate docket files in the public document room in Washington, D.C. and in local public document rooms as may be required by NRC regulations if the number of copies submitted is insufficient for this purpose. Copies made by the NRC must include the copyright notice in all instances and the proprietary notice if the original was identitied as proprietary. O O. mAapgxn344w.non:lb-100395 REVislON: 1
.O Q TABLE OF CONTENTS Section Title P_ age
< < < VOLUME 1 > > >
ACKNOWLEDGMENTS xvi
SUMMARY
1
1.0 INTRODUCTION
1-1 1.1 Background 1.1-1 1.2 Test Objectives 1,2-1 1.3 Important Small-Break Loss-of-Coolant Accident and Long-Term Cooling 1.3-1 Phenomena Identification and Ranking Table 1.3.1 Small-Break Loss-of-Coolant Accident 1.3-1 1.3.2 Long-Term Cooling Transient 1.3-3 1.4 Test Facility Scaling 1.4-1 1.5 Test Scaling Assessment and Dimensions 1.5-1 2.0 FACILITY DESCRIPTION
SUMMARY
2-1 q 2.1 Overall Facility Description 2.1-1 2.2 Facility Instrumentation 2.2-1 2.2.1 Differential Pressure Transmitters (FDP, LDP, DP) 2.2-1 2.2.2 Pressure Transmitters 2.2-1 2.2.3 Magnetic Flow Meters 2.2-1 2.2.4 Ileated Phase Switches 2.2-2 2.2.5 lleat Flux Meters 2.2-2 2.2.6 Load Cells 2.2-2 2.2.7 'Ihermocouples 2.2-2 3.0 TEST
SUMMARY
3-1 3.1 Test Validation 3.1-1 3.2 Test Matrix 3.2-1 4.0 DATA REDUCTION METHODOLOGY 4-1 4.0.1 Nomenclature 4-1 4.0.2 Energy Equation Approximation 4-3 4.0.3 Amnient Conditions 4-4
- 4.1 LDP Compensation Function 4.1-1 4.2 Selected Level Compensations 4.2-1 4.3 Accumulators 4.3-1 4.3.1 Fluid Mass Conservation Equations 4.3-1
[ 4.3.2 Fluid Energy Conservation Equations 4.3-5
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TABLE OF CONTENTS (Continued) Section Title Page 4.4 Core Makeup Tanks and Cold-Leg Balance Lines 4.4-1 4.4.1 Core Makeup Tank Fluid Mass Conservation Equations 4.4-1 4.4.2 Core Makeup Tank Fluid Energy Conservation Equations 4.4-7 4.4.3 Core Makeup Tank Metal Energy Conservation Equations 4.4-10 4.4.4 Cold-Leg Balance Line Fluid Mass Conservation Equations 4.4-14 4.4.5 Cold-Leg Balance Line Fluid Energy Conservation Equations 4.4-18 4.4.6 Cold-Leg Balance Line Metal Energy Conservation Equations 4.4-20 4.5 In-Containment Refueling Water Storage Tank (IRWST) 4.5-1 4.5.1 General Mass and Energy Balance Formulation 4.5-1 4.5.2 Case 1 4.5-3 4.5.3 Case 2 4.5-6 4.5.4 Case 3 4.5-7 4.5.5 Case ~ 4 4.5-9 4.5.6 Direct Vessel Injection Line Flow Reversal 4.5-10 4.5.7 Energy Loss due to Ambient Heat Transfer Rate 4.5-11 4.5.8 Energy Loss to Metal 4.5-13 4.5.9 Fluid Stored Energy 4.5-13 4.6 Automatic Depressurization System 1-3 Separator 4.6-1 4.6.1 Automatic Depressurization System 1-3 Separator Liquid Inventory 4.6-2 4.6.2 Steam Flow Rates 4.64 4.6.3 Liquid Flow Rates 4.6-4 4.6.4 Total Flow Rate 4.6-5 4.6.5 Energy Balance 4.6-6 4.7 Automatic Depressurization System-4 Separators 4.7-1 4.7.1 Automatic Depressurization System-4 Separator Liquid Inventory 4.7-2 4.7.2 Steam Flow Rates 4.7-4 4.7.3 Liquid Flow Rates 4.7-5 4.7.4 Total Flow Rate 4.7-5 4.7.5 Energy Balance 4.7-7 4.8 Break Separator 4.8-1 4.8.1 Break Separator Liquid Inventory 4.8-2 4.8.2 Steam Flow Rates 4.8-4 4.8.3 Liquid Flow Rates 4.8-5 4.8.4 Total Flow Rate 4.8-5 4.8.5 Energy Balance 4.8-6 4.9 Sumps 4.9-1 4.9.1 Sump Liquid Inventory 4.9-2 4.9.2 Sump Steam Exhaust Flow 4.9-4 m:\ap0A2344w.non:lb.100395 jv REVISION: 1
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1 TABLE OF CONTENTS (Continued) l Section Title Page I l 4.9.3 Sump Injection 4.9-5 ) 4.9.4 Total Flon Rate Out of the Sump 4.9-6 4.9.5 Energy Balance 4.9-6 4.10 Passive Residual Heat Removal 4.10-1 4.10.1 Fluid Mass Conservation Equation 4.10-1 4.10.2 Fluid Energy Conservation Equation 4.10-6
- 4.10.3 Tube Metal Energy Conservation Equation 4.10 10 4.11 Reactor Pressure Vessel 4.11-1 4.11.1 Core Vessel Model 4.11-2 4.11.2 Core Power and Flow Model 4.11-3 4.11.3 Energy Balance 4.11-7 4.12 Downcomer 4.12-1 4.12.1 Downcomer Level and Mass 4.12-1 4.12.2 Fluid Stored Energy 4.12-1 4.13 Steam Generator Primary Side 4.13-1 4.13.1 Inlet Plenum 4.13-1 4.13.2 Steam Generator Tubes 4.13-6 4.13.3 Outlet Plenum 4.13-11 4.14 Steam Generator Secondary Side 4.14-1 4.14.1 Inputs and Assumptions 4.14-1 4.14.2 Mass Balance Calculations 4.14-2 4.15 Pressurizer 4.15-1 4.15.1 Inputs and Assumptions 4.15-1 4.15.2 Mass Balance Calculation 4.15 2 4.15.3 Energy Balance 4.15-5 4.16 Pressurizer Surge Line 4.16-1 4.16.1 Inputs and Assumptions 4.16-1 4.16.2 Mass Balance 4.16-2 3
4.16.3 Energy Balance 4.16-3 l 4.17 Cold Legs 4.17-1 1 4.17.1 Cold Leg with Core Makeup Tank Balance Lines 4 (CL-1 and CL-3) 4.17-1 4.17.2 Cold Leg without Core Makeup Tank Balance Lines 5 (CL-2 and CL-4) 4.17-7 4.18 Hot Legs 4.18-1 4.18.1 Mass Storage in the Hot Legs 4.18-2 4.18.2 Energy Terms 4.18-3 . O l l m:\ap6000344w.non:ltr100395 y REVISION: 1 l I
p;
\
s TABLE OF CONTENTS (Continued) i
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Section Title Page 4.9.3 Sump Injection 4.9-5 4.9.4 Total Flow Rate Out of the Sump 4.9-6 4.9.5 Energy Balance 4.9-6 4.10 Passive Residual Heat Removal 4.10-1 4.10.1 Fluid Mass Conservation Equation 4.10-1 4.10.2 Fluid Energy Conservation Equation 4.10-6 4.10.3 Tube Metal Energy Conservation Equation 4.10-10 4.11 Reactor Pressure Vessel 4.11-1 4.11.1 Core Vessel Model 4.11-2 4.11.2 Core Power and Flow Model 4.11-3 4.11.3 Energy Balance 4.11-7 4.12 Downcomer 4.12-1 l 4.12.1 Downcomer Level and Mass 4.12-1 4.12.2 Fluid Stored Energy 4.12-1 4.13 Ster;s Generator Primary Side 4.13 1 4.1.i.1 Inlet Plenum 4.13-1 4.13.2 Steam Generator Tubes 4.13-6 ,U 4.13.3 Outlet Plenum 4.13-11 l 4.14 Steam Generator Secondary Side 4.14-1 4.14.1 Inputs and Assumptions 4.14-1 , 4.14.2 Mass Balance Calculations 4.14-2 4.15 Pressurizer 4.15 1 4.15.1 Inputs and Assumptions 4.15-1 4.15.2 Mass Balance Calculation 4.15-2 , 4.15.3 Energy Balance 4.15-5 4.16 Pressurizer Surge Line 4.16-1 4.16.1 Inputs and Assumptions 4.16-1 4.16.2 Mass Balance 4.16-2 4.16.3 Energy Balance 4.16-3 i 4.17 Cold Legs 4.17 1 4.17.1 Cold Leg with Core Makeup Tank Balance Lines (CL-1 and CL-3) 4.17-1 1 4.17.2 Cold Leg without Core Makeup Tank Balance Lines (CL-2 and CL-4) 4.17-7 . 4.18 Hot Legs 4.18-1 I 4.18.1 Mass Storage in the Hot Legs 4.18-2 4.18.2 Energy Terms 4.18-3 j 1 mAaptaA2344w.non:Ib-100395 v REVISION: 1 l 1
TABLE OF CONTENTS (Continued) l l Section Title h I 1 4.19 Pressure Conversions 4.19-1 4.20 Adjusted Data 4.20-1 4.21 System Mass Analysis 4.21-1 4.21.1 Total System Mass Inventory 4.21-1 4.21.2 Primary System Mass Balance 4.21-2 4.21.3 Sump Mass Balance 4.21-4 4.21.4 In-Containment Refueling Water Storage Tank Mass Balance 4.21-4 4.21.5 Variations in Mass Balance Models with Break Location 4.21-5 4.22 Overall System Energy Balance 4.22-1 5.0 ANALYSIS OF OSU TEST DATA 5-1 5.1 Analysis of Matrix Test SB01 5.1-1 5.1.1 Facility Performance 5.1.1-1 5.1.2 Short-Tena Transient 5.1.2-1 5.1.3 Long-Term Transient 5.1.3-1 5.2 Analysis of Matrix Test SB18 5.2-1 5.2.1 Facility Performance 5.2.1-1 5.2.2 Short-Term Transient 5.2.2-1 5.2.3 Long-Term Transient 5.2.3-1 5.3 Analysis of Matrix Test SB06 5.3-1 5.3.1 Facility Performance 5.3.1-1 5.3.2 Short-Term Transient 5.3.2-1 5.3.3 Long-Term Transient 5.3.3-1 5.4 Analysis of Matrix Test SB09 5.4-1 5.4.1 Facility Performance 5.4.1-1 5.4.2 Short-Term Transient 5.4.2-1 5.4.3 Long-Term Transient 5.4.3-1 5.5 Analysis of Matrix Test SB10 5.5-1 5.5.1 Facility Performance 5.5.1-1 5.5.2 Short-Term Transient 5.5.2-1 5.5.3 Long-Term Transient 5.5.3-1 5.6 Analysis of Matrix Test SB12 5.6-1 5.6.1 Facility Performance 5.6.1-1 5.6.2 Short-Term Transient 5.6.2-1 5.6.3 Long-Term Transient 5.6.3-1 O mAap6(Xh2344w.non:lb-100395 yi REVISION: 1
t b TABLE OF CONTENTS (Continued)
- Sution Title g
< < < VOLUME 2 > > >
4 5.7 Analysis of Matrix Test SB13 5.7-1 5.7.1 Facility Performance 5.7.1-1 5.7.2 Shon-Term Transient 5.7.2-1 5.7.3 Long Term Transient 5.7.3 1 5.8 Analysis of Matrix Test SB14 5.8-1 Facility Performance 5.8.1 5.8.1-1 5.8.2 Short Term Transient 5.8.2-1 1 5.8.3 Long-Term Transient 5.8.3-1 5.9 Analysis of Matrix Test SB15 5.9-1 l 5.9.1 Facility Performance 5.9.1-1 Short-Term Transient 5.9.2 5.9.2-1 5.9.3 Long-Term Transient 5.9.3-1 5.10 Analysis of Matrix Test SB19 5.10-1 5.10.1 Facility Performance 5.10.1-1 5.10.2 Short-Term Transient 5.10.2-1
\~ 5.10.3 Long-Term Transient 5.10.3-1 5.11 Analysis of Matrix Test SB21 5.11-1 5.11.1 Facility Performance 5.11.1-1 5.11.2 Short-Term Transient 5.11.2-1 5.11.3 Long-Term Transient 5.11.3 1 5.12 Analysis of Matrix Test SB23 5.12-1 5.12.1 Facility Performance 5.12.1-1 5.12.2 Short Term Transient 5.12.2-1 5.12.3 Long-Term Transient- 5.12.3-1 6.0 'EST FACILITY PERFORMANCE 61 6.1 Observed Thermal-liydraulic Phenomena 6.1-1 6.1.1 Core Makeup Tank Reflood Response 6.1.1-1 6.1.2 Passive Residual Heat Removal System Performance 6.1.2-1 6.1.3 Flow Oscillations During Long-Term Cooling 6.1.3-1 6.1.4 Effects of Accumulator Nitrogen 6.1.4-1 6.2 Data Evaluation 6.2.1 Core Energy 6.2.1 1 6.2.2 Mass Balance 6.2.2-1 6.2.3 Overall Energy Balance 6.2.3-1 map 60mmw.wa:n>.ioo395 vii REVISION: 1
TABLE OF CONTENTS (Continued) Section Title h 7.0 SYSTEM ANALYSIS FOR SMALL-BREAK LOSS-OF-COOLANT ACCIDENTS AND LONG TERM COOLING 7.0-1 7.1 Variations in Break Size 7.1-1 7.1.1 Passive Residual Heat Removal Behavior 7.1.1-1 7.1.2 Event Timing Discussicas 7.1.2-1 7.13 Downcomer Condensadan Phenomena 7.13-1 7.1.4 Break Flow and Flow Integrals 7.1.4-1 7.2 Variations in Break Location 7.2-1 7.2.1 Event Timing / Phenomena 7.2.1-1 7.2.2 Core Makeup Tank Drain / Refill Behavior 7.2.2-1 7.3 Closure on the Phenomena Identification and Ranking Table for AP600 Small-Break Loss-of-Cooiant Accident and Long-Term Cooling for the OSU Tests 7.3-1
8.0 CONCLUSION
S 8-1
9.0 REFERENCES
9-1 m:%60th2344w.non:lt>100395 yjjj REVISION: 1 I l l
Lj LIST OF TAllLES Table Title Page 1,3-1 Phenomena Identification Ranking Table for AP600 SBLOCA i and LTC Transient 1.3-5 1.4 1 General System Hierarchy: OSU/APo00 Scaling Analysis 1.4-7 1.5-1 initial Conditions for OSU Test Facility to Model a 2-in. Cold-Leg Break 1.5-3 1.5-2 Scale Factors to Relate the AP600 Plant to OSU NOTRUMP Calculations 1.5-4 1.5-3 Distortion Factors for the AP600 Dominant Processes identifled Using the li2TS Methodology 1.5-5 3.1 1 Overall Acceptance Criteria 3.1-2 3.2-1 OSU Matrix Test Summary 3.2-3 4.2-1 Pressures and Temperatures for Compensated LDPs 4.2-2 4.3-1 Instrumentation Employed for Accumulator Fluid Calculations 4.3-8 4.4-1 Instrumentation Employed for CMT Fluid Calculations 4.4-25 4.4-2 Volume Versus Height Tables for CMT Fluid Volume Calculations 4.4-26 4.4-3 CMT Metal Wall Thermocouple Instrumentation 4.4-27 p 4.4-4 Data for CMT Metal Energy Calculations 4.4-28 V 4.4-5 Specific Heat Capacity Versus Temperature Table for 4.4-29 CMT Metal Energy Calculations 4.4-6 Instrumentation Employed for Cold-Leg Balance Line Fluid Calculations 4.4-29 4.4-7 Volume Versus Height Tables for Cold-Leg Balance Line Fluid Volume Calculations 4.4-30 4.4-8 Data for Cold-Leg Balance Line Metal Energy Calculations (per segment) 4.4-30 4.4-9 Data for Cold-Leg Balance Line Metal Energy Calculations 4.4-31 4.4-10 Specific Heat Capacity Versus Temperature Table for Cold-Leg Balance 4.4-31 Line Metal Energy Calculations 4.5- 1 IRWST Mass and Energy Calculations identification of Fluid 4.5-15 Thermocouples and Elevation 4.5-2 Volume Versus Height Table for IRWST Fluid Volume Calculations 4.5-15 4.5-3 Data for IRWST Metal Energy Calculations (per segment) 4.5-16 4.5-4 Data for IRWST Metal Energy Calculations 4.5-16 4.5.5 Data for IRWST Energy Loss Due to Ambient Heat Transfer 4.5-14 4.6-1 Instrumentation to be Used for ADS 1-3 Levels Instrument Correction 4.6-9 4.6-2 Volume Versus Height for ADS 1-3 Volume Calculations 4.6-9 4.6-3 ADS 1-3 Separator Steam and Liquid Pressure and Temperature 4.6-9 Instrument Channels A I t
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m:\ap600\2344w.noo:lb-100395 ix REVislON: 1
LIST OF TAllt.ES (Continued) O i Table Title Pace 4.7-1 Instrumentation to be Used for ADS-4 Separator Levels Instrument Correction 4.7-12 4.7-2 ADS-4 Separator Steam and Liquid Pressure and Temperature Instrument Channels 4.7.12 4.7-3 Instruments to be Used in Calculation of Local Flow Qualities 4.7-13 4.7-4 Volume Versus lielght for ADS 4-1 Fluid Volume Calculations 4.7 14 4.7-5 Volume Versus Height for ADS 4-2 Fluid Volume Calculations 4.7 14 4.8-1 Instrumentation to be Used for Break Separator Mass and Energy Balance 4.8-9 4.8-2 Break Separator Steam Exhaust and Liquid Pressure and Temperature Instrument Channels 4.8-9 4.8-3 Volume Versus Height for Break Separator Fluid Volume Calculations 4.8-10 4.9-1 Irbtrumentation to be Used for Sump Mass and Energy Balance 4.9-10 4.9 2 Sump Steam Exhaust and Injection Pressure and Temperature 4.9-10 Instrument Channels 4.9-3 Data for Sumps Metal Energy Calculations (per segment) 4.9-11 > 4.9-4 Data for Sumps Metal Energy Calculations 4.9-11 4.9-5 Specific IIcat Capacity versus Temperature Table for 4.9-12 Sumps Metal Energy Calculations 4.10-1 Instrumentation Employed for PRHR Fluid Calculations 4.10-12 4.10-2 Volume Versus Height Tables for PRHR Fluid Volume Calculations 4.10-12 4.10-3 Data for PRHR Tube Metal Energy Calculations (per segment) 4.10-13 4.10-4 Specific Heat Capacity Versus Temperature Table for PRHR Tube 4.10-13 Metal Energy Calculations 4.11-1 Core Vessel Model Geometry 4.11-10 4.11-2 Mass Methodology Effects 4.11-10 4.11-3 Heater Rod Instrumentation 4.11-10 4.11-4 Power Distribution 4.11-11 4.11-5 Constant Multipliers for the Free Convection Heat Transfer Coefficient 4.11-11 4.11-6 OSU Test Analysis Plot Package for Section 4.11 4.11-12 4.12-1 Volume Versus Height Table for Downcomer Fluid Volume Calculations 4.12-3 4.13-1 Data Channel ID for SG Inlet Plenum Mass and Energy Calculations 4.13-13 4.13-2 Volume Versus Height Table for Steam Generator Inlet Plenum 4.13-14 4.13-3 Volume Versus Height Table for Steam Generator Tubes (Down-Hill Side) 4.13-15 4.13-4 Volume Versus Height Table for Steam Generator Tubes (Up-Hill Side) 4.13-15 4.13-5 Volume Versus Height Table for Steam Generator Outlet Plenum 4.13-16 4.14-1 Instrument Channel ids for SG-1 Secondary-Side Mass and Energy 4.14-7 Calculations O manp600\2344w.non:Ib-100395 x REVISION: 1
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4 i --s'( - ' lV 1 LIST OF TAllLES (Continued) Table Title - ,P_ age 4.14-2 Instrument Channel ids for SG.2 Secondary-Side Mass and 4.14-8 . Energy Calculations
- 4.14-3. Instrument Channel ids for SG System Secondary-Side Mass 4.14-9 and Energy Calculations 4.14-4 Fluid Height Versus Volume for SG Secondary-Side Mass and 4.14-9 Energy Calculations ,
j 4.15-1 Instrument Channel ids for Pressurizer Mass and Energy Calculations '4.15-11 4.15-2 _ Fluid Height Versus Volume for Pressurizer Mass Calculations 4.15-12 4.15-3 Metal Data for Pressurizer Metal Energy Calculations . 4.15-12
. 4.15-4 Temperature Versus Heat Capacity for Pressurizer Metal Energy Calculations 4.15-13 3
4.16-1 Instrument Channel ids for Pressurizer Surge Line Mass and 4.16-8 . Energy Calculations 4.16-2 Fluid Height Versus Volume for Pressurizer Surge Line Mass Calculations 4.16-8 i 4.16-3 Metal Data for Pressurizer Surge Line Metal Energy Calculations 4.16-9 4.16-4 Temperature Versus Heat Capacity for Pressurizer Surge Line Metal 4.16-9 Energy Calculations ; j 4.17-1 Data Channel ids Used to Calculate Local Fluid Properties for Flow Meters 4.17-13 l
, 4.17-2 Data Channel ids Used to Calculate Fluid Properties for Levels Transducers 4.17 13 ]
5 4.17 3- Data Channel ids Used to Calculata Local Fluid Properties for Flow Meters 4.17-14 , 4.17-4 Data Channel ids Used to Calculate Fluid Properties for Levels Transducers 4.17-14 l !- 4.18-1 Data Channel ids Used in Hot-Leg Mass and Energy Calculations 4.18-5 4.19-1 Pressure Conversions 4.19-2 . 4.20-1 Channels for Data Smoothing 4.20-2 . 4.20-2 OSU Test Analysis' Plot Package for Section 4.20 4.20-6 ) j 4.21 1 Data Channel ids Used for Flow Meter Calculations 4.21-10 e 5.1.11 OSU Test Analysis Plot Package for Subsection 5.1.1 5.1.1-8 5.1.2-1 OSU Test Analysis Standard Plot Package for Subsection 5.1.2 5.1.2-10 5.1.3-1 OSU Test Analysis Standard Plot Package for Subsection 5.1.3 Long-Term 5.1.3-6 , Transient 5.2.1 -1 OSU Test Analysis Plot Package for Subsection 5.2.1 5.2.1-6 . 5.2.2-1 OSU Test Analysis Standard Plot Package for Subsection 5.2.2 5.2.2-10 4 5.2.3-1 OSU Test Analysis Standard Plot Package for Subsection 5.2.3 5.2.3-4 Long-Term Transient l l' 5.3.1-1 OSU Test Analysis Plot Package for Subsection 5.3.1 5.3.1-4 ! j 5.3.2-1 OSU Test Analysis Standard Plot Package for Subsection 5.3.2 5.3.2-5 I O m:\np600Q344w.non:ltrl00395 xi REVISION: I I
LIST OF TAllLES (Continued) Table Title Page 5.3.3-1 OSU Test Analysis Standard Plot Package for Subsection 5.3.3 5.3.3-4 Long-Term Transient 5.4.1 -1 OSU Test Analysis Plot Package for Subsection 5.4.1 5.4.1-4 5.4.2- 1 OSU Test Analysis Standard Plot Package for Subsection 5.4.2 5.4.2-5 5.4.3-1 OSU Test Analysis Standard Plot Package for Subsection 5.4.3 5.4.3-4 Long-Term Transient 5.5.11 OSU Test Analysis Plot Package for Subsection 5.5.1 5.5.1-4 5.5.2-1 OSU Test Analysis Standard Plot Package for Subsection 5.5.2 5.5.2-5 5.5.3-1 OSU Test Analysis Standard Plot Package for Subsection 5.5.3 5.5.3-4 Long-Term Transient 5.6.1-1 OSU Test Analysis Plot Package for Subsection 5.6.1 5.6.1-4 5.6.2-1 OSU Test Analysis Standard Plot Package for Subsection 5.6.2 5.6.2-5 5.6.3-1 OSU Test Analysis Standard Plot Package for Subsection 5.6.3 5.6.3-4 Locg-Term Transient 5.7.1-1 OSU Test Analysis Plot Package for Subsection 5.7.1 5.7.1-4 5.7.2-1 OSU Test Analysis Standard Plot Package for Subsection 5.7.2 5.7.2-5 5.7.3-1 OSU Test Analysis Standard Plot Package for Subsection 5.7.3 5.7.3-4 Long-Term Transient 5.8.1-1 OSU Test Analysis Plot Package for Subsection 5.3.1 5.8.1-4 5.8.2-1 OSU Test Analysis Standard Plot Package for Subsection 5.8.2 5.8.2-5 5.8.3-1 OSU Test Analysis Standard Plot Package for Subsection 5.8.3 5.8.3-4 Long-Term Transient 5.9.1-1 OSU Test Analysis Plot Package for Subsection 5.9.1 5.9.1-4 5.9.2-1 OSU Test Analysis Standard Plot Package for Subsection 5.9.2 5.9.2-5 5.9.3-1 OSU Test Analysis Standard Plot Package for Subsection 5.9.3 5.9.3-4 Long-Term Transient 5.10-1-1 OSU Test Analysis Plot Package for Subsection 5.10.1 5.10.1-4 5.10.2-1 OSU Test Analysis Standard Plot Package for Subsection 5.10.2 5.10.2-5 5.10.3-1 OSU Test Analysis Standard Plot Package for Subsection 5.10.3 5.10.3-4 Long-Term Transient 5.11.1-1 OSU Test Analysis Plot Package for Subsection 5.11.1 5.11.1-4 5.11.2-1 OSU Test Analysis Standard Plot Package for Subsection 5.11.2 5.11.2-5 5.11.3-1 OSU Test Analysis Standard Plot Package for Subsection 5.11.3 5.11.3-5 Long-Term Transient 5.12.1 1 OSU Test Analysis Plot Package for Subsection 5.12.1 5.12.1-4 5.12.2-1 OSU Test Analysis Standard Plot Package for Subsection 5.12.2 5.12.2-5 5.12.3-1 OSU Test Analysis Standard Plot Package for Subsection 5.12.3 5.12.3-4 Long-Term Transient m:\ap600\2344w.non:lb-100395 xil REVISION: 1
c' ()/ LIST OF TAHLES (Continued) Table , Title _P,aage 6.1.1-1 OSU Test Analysis Plot Package for Subsection 6.1.1 6.1.1-3 6.1.2-1 Instrumentation for Calculating the PRHR/IRWST Heat Balance 6.1.2-6 6.1.2-2 Key Parameters for Calculating the PRHR/IRWST Heat Balance 6.1.2-6 6.1.3-1 OSU Test Analysis Plot Package for Subsection 6.1.2 6.1.2-7 6.1.3-1 Summary of Flow Oscillation Data 6.1.3-31 6.1.3-2 OSU Test Analysis Plot Packags for Subsection 6.1.3 6.1.3-32 6.1.4-1 Summary of Accumulator Behavior for Test SB01 6.1.4-3 6.1.4-2 OSU Test Analysis Plot Package for Subsection 6.1.4 6.1.4-3 6.2.1-1 Saturated Water Properties 6.2.1-8 6.2.1-2 OSU Test Analysis Plot Package for Subsection 6.2.1 6.2.1-9 6.2.2-1 OSU Test Analysis Plot Package for Subsection 6.2.2 6.2.2-6 6.2.2-2 Steam Flow during Short and Long-Term Transients 6.2.2-8 6.2.3-1 OSU Test Analysis Plot Package for Subsection 6.2.3 6.2.3-6 7-1 Sequence of Events Comparison for Matrix Tests 7-3 7.1.1-1 PRHR Behavior for Various Cold-Leg Break Sizes 7.1.1-2
-7.1.4-1 Subsection 7.1.4 Plot Package 7.1.4-2 A 7.2.2- 1 OSU Test Analysis Plot Package for Subsection 7.2.2 7.2.2-5 >) %/
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LIST OF FIGURES Figure Title Page 1.4-1 Decomposition Paradigm and liierarchy 1.4-8 1.4-2 AP600 SBLOCA Scenario 1.4-9 1.4-3 Scaling Analysis Flow Diagram for System Depressurization 1.4-10 1.51 Normalized Pressure Comparisons between AP600 and OSU Facility 1.5-6 1.5-2 Normalized CMT-1 Level for AP600 and OSU Facility 1.5-7 i 1.5-3 Normalized CMT-2 Level for AP600 and OSU Facility 1.5-8 1.5-4 Normalized ACC-1 Level for AP600 and OSU Facility 1.5-9 1.5-5 Normalized ACC-2 Level for AP600 and OSU Facility 1.5-10 1.5-6 Normalized ADS 1-3 Flows for AP600 and OSU Facility 1.5-11 1.5-7 Normalized Break Flow for AP600 and OSU Facility 1.5-12 1.5-8 Normalized System Mass for AP600 and OSU Facility 1.5-13 1.5-9 Comparison of OSU and SPES-2 CMT-1 Injection Flow Rate 1.5-14 1.5-10 Comparison of OSU and SPES-2 2-in. Break Pressure Histories 1.5-15 1.5-11 Comparison of OSU and SPES-2 CMT-1 Liquid Level Histories 1.5-16 1.5-12 Comparison of OSU and SPES-2 ACC-1 Liquid Level Histories 1.5-17 1.5-13 Comparison of OSU and SPES-2 ACC-1 Injection Flow Rate 1.5-18 1.5-14 Comparison of OSU and SPES-2 IRWST 1 Flow Rate 1.5-19 2.1-1 Isometric Drawing of the OSU Test Facility 2.1-4 2.1-2 Simplified Flow Diagram of the OSU Test Facility 2.1-5 O m:\ap600\2344w.non:Ib-100395 xiv REVISION: 1
i- ) i 4 i ! . . ) i '/ !b ACRONYMS'
)
ADS automatic depressurization system APEX advanced plant test facility at OSU j ASME American Society of Mechanical Engineers
; BAMS break and ADS measurement system j CCT condensate collection tank J
CD ROM . compact disk read-only memory CMT. core makeup tank j. CL cold leg CLBL cold leg balance line CRP condensate return pump 1 CVS chemical and volume control system j DAS data acquisition system DBE design basis event DEG double-ended guillotine . DP differential pressure transmitter
. DVI direct vessel injection FMM magnetic flow meter j GSM general scaling methodology H2TS hierarchical two-tiered scaling analysis HFM heat flux meters HPS heated phase switch HX heat exchanger
-(q HL hot leg l I bj 1RWST in-containment refueling water storage tank 1 . LAN- local area network l , LCS lower containment sump l
- LDP level differential pressure ;
- LOCA loss-of-coolant acciden LRGMS large main steam ,
3 LTC long-term cooling ' MSS main steam system i NSS nonsafety systems OSU Oregon State University , PC personal computer i PIRT phenomena identification ranking table ' PPIRT plausible phenomena identification ranking table l
- i. PQP project quality plan
, PRHR passive residual heat removal l i PT pressure transducer PWR l pressurized water reactor l PXS passive core cooling system QLR quick look report 1 RCP reactor coolant pump
- RCS reactor coolant system l
1 , RNS normal residual heat removal system RPV reactor pressure vessel ,.. S signal safety signal l ! mAap600\2344w. son:lb-100395 xV REVISION: 1
i i l ACRONYMS (Continued) O1 SASM severe accident scaling methodology , SBLOCA small-break loss-of-coolant accident l SCR silicon-controlled rectifier l SG steam generator l SGS steam generator system l VI virtual instrumentation l l l a I 4 O I I I i O m:\apMXA2344w.non:lb-100395 XVI REVISION: 1
t 4 4
. ACKNOWLEDGMENTS The authors express their appreciation for the extensive discussions and inputs obtained from Mr. L.-
K. (Louis) Lau, lead designer of the OSU test facility, and Mr. C. L. (Carl) Dumsday, Manager of
] Test Operations for Westinghouse at the OSU test facility.
To W. (Bill) Brown, who performed the technical review of the equations ' developed for analysis of the data, and to M. (Michael) Parks, team leader, I. (Igor) Halijasmaa, R. (Rain) Veinjary and P. (Peter) Ward, who performed the verification and validation check calculations for the coding of those
; equations, the authors extend their thanks and appreciation for an arduous task that was done well.
. We also gratefully' acknowledge M. H. (Mike) Mankowski for his efforts in generating composite
, figures of safety system flows included in the transient descriptions included in Section 5 of this report and D. J. (Don) Longo for his efforts in preparing and performing heat transfer calculations with the CMT data. The thorough and insightful review of the report performed by E. H. (Earl) Novendstern, Manager, I Plant Safety Analysis, E. J. (Gene) Piplica, Manager, Test Engineering, and B. A. (Brian) McIntyre, [ Manager, Advanced Plant Safety and Licensing, is gratefully acknowledged; their comments significantly improved the final report. i t Our thanks to the technical editors, Mariane Cox, Leslie McSwain, and Tom Hendrick for the efforts and understanding they put forth to transform the report from the initial drafts written in phrases and _p jargon to its current form and format. To the Word Processing team members, a special thanks is extended, not only for their skill, patience and cooperation in working with the large volume of material that formed the basis of this report, but also for their willing cooperation to work with the editors and authors, the long hours they put in, and
- the enthusiasm they displayed through it all.
. Finally, but not in the least, the authors express their gratitude to R. B. (Bob) Tupper and P. (Paul) Wardman for the assistance in initially scheduling and tracking the tasks necessary to complete this report, and especially to B. E. (Bruce) Rarig for his diligence and determination in tracking tasks, i attention to detail, and promoting communications among the following up with the authors, reviewers,
- .' technical editors, word processing team, and reproduction office during the critical final stages of the report generation effort.
J_ ^ nap 6ao23m. on:isioo395 xvii REVISION: 1 i :
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SUMMARY
'Ihe Oregon State University (OSU) test facility is a 1/4-height, reduced-pressure simulation of tiie -
. AP600 nuclear steam supply system and the AP600 passive safety features. A series of design-basis events were simulated at OSU to obtain data for verification and validation of the computer models used for the safety analysis of AP600,
- The purpose of this report is to describe the analysis of the test data and the thermal-hydraulic behavior of the test facility, to identify the phenomena observed in the tests and tle relationship to the phenomena identification ranking table (PIRT), and to show the applicability of the OSU tests for computer model verification and validation through mass and energy balances.
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4 X ([ ' t,0 INTRODUCTION ; His report describes the analysis of the Oregon State University (OSU) test data used to validate certain AP600 safety analysis computer codes.' The test data report for the OSU tests is given in the OSU Final Data Report,m which describes the test facility, the valid instrumentation, and the test - facility performance for the different tests. His report will examine, in additional detail, the thermal-hydraulic behavior of the test facility and the phenomenon observed in the tests, as identified in the phenomena identification ranking table (PIRT), Table 1.3-1. This analysis will aid computer code validation activities. . De OSU test facility is a 1/4-height, reduced-pressure model of the AP600 and its passive emergency core cooling systems. The test facility located at the Radiation Center at the University in Corvallis, Oregon includes the reactor coolant system (RCS), steam generators (SGs), passive core cooling system (PXS), automatic depressurization system (ADS), and nonsafety-related injection systems, such as the normal residual heat removal system (RNS) and the chemical and volume control system (CVS). The test facility, fabricated from austenitic stainless steel designed for normal operation at 450*F and 400 psig, was scaled using the hierarchical, two-tiered scaling analysis (H2TS) method developed by the U.S. Nuclear Regulatory Commission (NRC). Simulated piping breaks were tested . In the hot leg (HL), cold leg (CL), pressure balance line between the cold leg and the core makeup tank (CMT), and the direct vessel injection (DVI) line. Decay heat that scaled to 3 percent of the full O V power (about 2 minutes after shutdown) was supplied by electrically heated rods in the reactor vessel. Simulated transients were programmed by the control system to proceed automatically. About 850 data channels were recorded by the data acquisition system (DAS) and downloaded to compact disks ) for subsequent data reduction and plotting. De OSU test facility was specifically designed to examine l the small-break loss-of-coolant transient (SBLOCA) periods as well as the long-term-cooling (LTC) aspects of the AP600 passive safety systems. , 1 1 d f-( l mAap600G344w. son:Ib 100395 11 REVIslON: I l l
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- 1.1 Background t
AP600 is a 600-MWe Westinghouse advanced reactor designed to enhance plant safety with accident ~
' mitigation features that, once actuated, depend only on natural forces, such as gravity and natural ..
circulation, to perform all required safety functions.
' he AP600 primary system is a two-loop design. Each loop contains one hot leg, two cold legs, and ' !
one SG with two canned motor reactor coolant pumps (RCPs) attached directly to the SG outlet. 'i I
- channel head. ;The passive safety systems comprise the following:
i Two full-pressure CMTs that provide borated makeup water to the primary system at full , system pressure.
. .Two accumulators (ACCs) that provide borated water to the reactor vessel if the primary .
pressure s 700 psia. t
- A passive residual heat removal (PRHR) heat exchanger (HX), comprised of a C-shaped tube bundle submerged in the in-containment refueling water storage tank (IRWST), that can J remove heat from the primary system at full system pressure. I h
- The ADS, which is comprised of a set of valves connected to the RCS at the pressurizer steam space and the two hot legs. The valves connected to the pressurizer vent to the IRWST through a sparger. The valves connected to the hot leg vent to the containment. Dese valves are opened sequentially to provide controlled depressurization of the primary system. )
l 1
. An IRWST that provides a large source of core cooling water, which drains by gravity after j the ADS has actuated.
A passive containment cooling system (PCS) that utilizes the AP600 steel containment shell to transfer heat to the environment (ultimate heat sink). The PCS was not directly included in , the OSU tests, however, the containment circulation of condensed liquid back to the IRWST or I sump was simulated for selective tests. In reviews of the AP600, the U.S. NRC identified several concerns regarding the performance of the AP600 passive safety systems. Those concerns include the following: I l Possible high-pressure passive safety system interactions that could retard cooling of the core. Possible active system / passive system interaction that could retard cooling of the core. p
- The dependence on small temperature differences resulting in small density differences, which V then are responsible for driving heads for recirculating flows.
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- The effects of code accuracy in predicting long transients in which the driving heads for How in the system are small.
+ The behavior of the primary system and the containment during the LTC phase of the transient; specifically, the reduced driving heads for flow from the sump and the resulting pressure drops in the primary system that could reduce venting and increase steam binding of the system at low pressure.
He OSU test facility was specifically designed to obtain experimental thermal hydraulic data that would address NRC concerns. He OSU test facility was constructed specifically to investigate the AP600 passive system characteristics. De facility design models the detail of the AP600 geometry, including the primary system, pipe routings, and layout for the passive safety systems. He primary system consists of one hot leg and two cold legs, with two active pumps and an SG for each of the two loops. Bere are two CMTs, each connected to a cold leg of one primary loop. The pressurizer is connected to the other primary loop, as in the AP600 plant design. Gas-driven accumulators are connected to the DVI lines. He discharge 1 nes from a CMT and one of the two IRWST and reactor sump lines are connected to each DVI line. The two independent lines of each stage of ADS 1,2, and 3 are modeled by one line containing an orifice. Two-phase flow from ADS 1-3 is separated in a swirl-vane separator, and liquid and vapor flows are measured to obtain the total flow rate. The separated flow streams are then recombined and discharged into the IRWST through a sparger. Thus, mass and energy flow from the ADS into the IRWST are preserved. De period for simulation included not only IRWST injection, but also IRWST draining and sump injection to simulate the LTC mode of the AP600. De time scale for the OSU test facility is about l one-half; that is, the sequence of events occurred about twice as fast in the test facility as in the AP660. To model the LTC aspects of the transients, two-phase flow from the break was separated in a swirl-vane separator, and the liquid and vapor portions of the total flow were measured. He liquid fraction of the flow was discharged to the reactor sump, as in the AP600 plant. He vapor was discharged to the atmosphere, and the equivalent liquid flow was capable of being added to the IRWST and sump to simulate the condensate return from passive containment. A similar approach was also used for the l two ADS-4 valves on the hot legs. He two-phase flow was separated in a swirl-vane separator, the I two stream flows were measured, the liquid phase was discharged into the reactor sump while the vapor phase was discharged to the atmosphere, and the liquid equivalent was capable of being added to the IRWST and sump. He IRWST, reactor sump, and separators could be pressurized to simulate containment pressurization following a postulated loss-of-coolant accident (LOCA). l Prior to the performance of matrix tests, a series of cold, low-pressure and hot, high-pressure l pre-operational tests were performed to characterize the OSU facility to show proper operation of the l m:\ap600Q344w.non:Ib-100395 1,].2 REVISION: 1 l
. ym facility and to verify that piping / component parameters properly matched the AP600 plant, and to provide benclunark data on facility behavior for the computer code analysis. The tests that were analyzed in Sections 5,6, and 7 were selected to examine the AP600 passive safety system s performance in mitigating the effects of design-basis events (DBEs). Events that were evaluated include LOCAs ranging from 0.5-in, diameter equivalent to the double -ended guillotine (DEG) break
.of an 8-in. DVI line. A larger break was also simulated for LTC perf, rmance via 4-in, equivalent diameter breaks on both the top and bottom of a cold leg.
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- 1.2 . Test Objectives -
De OSU facility was designed and constructed to specifically examine the LTC performance of the
- - AP600 passive safety-related systems and their interaction with the nonsafety-related active systems.
4 he range and types of tests investigated in the OSU test facility covered the ranges and phenomena expected for the SBLOCA and LTC transients. Tests were performed to examine the different break ~ sizes and locations to cover the range of phenomena of interest. The data from the tests are used to i validate the safety analysis computer codes used to analyze the AP600. < To cover the range, of conditions for the LTC transient, various tests were performed. One test modeled as large a break as possible to rapidly depressurize the facility so that decay power would be at a high value when LTC began. Other tests were performed with conditions that would result in a J hot IRWST and sump when the primary system transitioned into LTC. Both conditions maximized the production of core steam to be vented through the ADS-4 valves to maintain sump injection. j A detailed scaling analysis was developed for the OSU tests to relate the scaled-pressure and reduced-height facility to the AP600 plant. De OSU Facility Scaling Report
- specifled the facility ,
i dimensions, resulting flow areas for the breaks, and pressure drops needed to preserve the phenomena expected for the AP600 SBLOCA and LTC transients. De scaling study provides the bridge to relate the AP600 SPES-2 Test Analysis Report
- to the similar OSU tests, and to relate the OSU tests to the h
Ls AP600 design. De following are the specific test objectives of the OSU program: (
- To provide data to establish the pedigree of the passive safety-related systems for LTC
- To provide overlap with the full-pressure, full-height SPES-2 tests
- so that an assessment of the scaling effects of the OSU tests could be made. Derefore, similar break locations and sizes (scaled) were examined in both facilities and comparisons were made.
4 To cover the range of phenomena expected for the AP600 LOCA in addition to the LTC period. d m:W344w.noa:Ib-100395 1.2-1 REVISION: 1
l b' _13 Important Small Hreak Loss-of-Coolant Accident and Long-Term Cooling Phenomena Identification and Ranking Table he OSU test matrix was developed to simulate the thermal-hydraulic phenomena expected during SBLOCA and LTC transients. IJ.1 Small Hreak Loss-of-Coolant Accident j The SBLOCA can be divided into the following four periods that characterize thermal-hydraulic phenomena:
- Blowdown - Initial depressurization from plant operating pressure to the SG secondary-side pressure, after which pressure stabilizes.
- Natural Circulation - De period from the stabilization of primary pressure with secondary-side pressure until ADS-1 is activated. De primary reactor system is cooled by different modes of heat transfer. Each cooling mode is dependent on the system mars inventory. As the mass is lost through the break, cooling proceeds from single-phase natural circulation, to two-phase natural circulation, to reflux condcasstion cooling. l l
h
- ADS 1-4 Blowdown - Once the CMTs drain to their setpoint, the ADS-1 valve opens and the l reactor system is depress, rized through the ADS flow path in addition to the break. As the CMT continues to drain into the reactor vessel, additional valves are opened on the pressurizer l and RCS hot legs to enhance blowdown of the system.
IRWST Injection - Stable injection from the IRWST indicates the complete depressurization of l the primary system down to containment pressure. Also, injection from the IRWST indicates the end of the small-break transient and the beginning of the LTC transient. l l Using these different periods, the imponant thermal-hydraulic phenomena have been identified and ; ranked in a PIRT (Table 13-1). His PIRT has been updated from that which was provided in the Applicability of the NOTRUMP Computer Code to the AP600 SSAR Sma!I Break LOCA Analysis
- and reflects small changes from the .W., COBRA / TRAC LTC Preliminary Validation Report.CD Re only changes other than footnotes were,1) ADS subsonic flow was increased in importance to "M" since this contributes to the system depressurization and the mass redistribution and the mass redistribution into the pressurizer and 2) in the aatural circulation phase, the wall stored energy to "L" from "N/A" due to the heat transfer to the CMT walls. Individual phenomena were emphasized for the ADS, and other components have been added to the PIRT. The phenomena for each identified phase of the small-break transient relative to the AP600 small-break performance is discussed in the following paragraphs.
m m:\ap600Q344w.non:1b-100395 1J-1 REVISION: 1
l l l l The reactor is assumed to be operating at normal full-power, steady-state conditions at the start of the O l i blowdown. The break opens at time zero, and pressurizer pressure begins to fall as mass is lost out the break. This depressurization is largely defined by critical flow through the break. With the break located at the bottom of the cold leg, a mixture flow exits the break for the majority of the transient, since the mixture level stays high in the reactor vessel. Pressurizer pressure falls below the safety signal setpoint, causing the reactor to trip. The safety systems actuation signal (S) follows and results in the opening of the CMT isolation valves. Once the residual fissions decrease, core power is defined by the decay heat model. The RCPs trip after a short delay. Pump performance, both before and after the trip, is modeled according to the pump characteristic curves. After the pumps coast down, the primary RCS is cooled by natural circulation, with energy removed from the primary system by the SGs via their safety valves and the break. Stored energy from the metal in the reactor vessel and pressurizer is transferred to the coolant. These phenomena are essentially the same for AP600 as for conventional pressurized water reactors (PWRs). Liquid in the upper plenum and upper head (depending on the temperature) will flash, and the upper head will start to drain. Blowdown phase phenomena unique - W600 are those associated with CMT delivery. Once the CMT isolation valves open, the CMT injects borated water by gravity-driven recirculation into the RCS through the DVI lines. The CMT injected volume is replaced with hot liquid via the cold-leg balance line (CLBL); this hot liquid collects at the top of the CMT. The downcomer fluid stays subcooled through the initial blowdown phase. For the natural circulation phase of the transient, the primary system exists in a quasi-steady-state O condition with the secondary side, with decay energy being removed by the SG secondary side as the primary system drains. He SG in the AP600 plays a more limited role in the natural circulation cooling phase than for conventional plants because the SGs drain relatively early in the transient. Since PRHR is activated on an S signal during a SBLOCA, the IRWST becomes the primary heat sink for the RCS carly in the transient. He PRHR will remove energy from the primary system, causing it to depressurize. The SG secondary side becomes a potential heat source once PRHR reduces primary pressure to that of the secondary side. PRHR is ranked high in the PIRT since it becomes a sigraficant heat removal path, particularly after primary pressure is less than SG pressure. Therefore, condersation in the SG tubes during a SBLOCA ceases early. The requirements for detailed models for condensation heat transfer in the SG tubes are not as significant for AP600 as for a conventional plant. %c importance shifts to the PRHR performance and the IRWST heat-sink behavior. The reverse heat transfer path due to secondary heating of the RCS primary system continues until the SGs drain. He CMT continues to deliver in the recirculation mode, but eventually a vapor region forms at the top of the CMT volume, and CMT draindown begins. As the CMT drains while injecting, its level falls to the ADS actuation setpoint, initiating the third phase of the AP600 SBLOCA transient, ADS blowdown. He downcomer and lower plenum are ranked as medium importance in the PIRT since they provide the driving head for natural circulation. He ADS blowdown phase continues through the actuation of ADS-1, ADS-2, ADS-3, and ADS-4 as the primary system depressurizes to approximately the containment pressure. He PIRT relates m w w m m :n-tmws 1.3-2 REVISION: 1
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' AP600-specific components, events, and phenomena that occur during automatic depressurization of l the RCS to achieve _ water injection by gravity from the IRWST. Since ADS-1 creates an opening at j
! the top of the pressurizer, the pressurizer two-phase fluid level increases markedly. Pressurizer tank level and surge line phenomena are significant factors in the depressurization behavior following ADS i: actuation. Flashing of fluid in the RCS occurs due to the depressurization caused by the ADS. l Following actuation of ADS 1, the next two stages of ADS, ADS-2 and ADS-3, activate via timers.
- Once the pressure drops below 700 psia, accumulator injection begins, reducing flow delivered from the CMT.. CMT flow may even be stopped temporarily due to pressurization of the DVI line by the.
- accumulator. De CMT drain rate, and DVI line and cold-leg balance line flow characteristics are
. significant because ADS 4 actuation is based on the CMT liquid level decreasing below a low-low setpoint value. _ Condensation of vapor on the CMT walls is of somewhat less importance since recirculation results in heating of the CMT.
Critical flow through the ADS stages is the major factor in determining when the RCS has l depressurized to the extent that the gravity injection of water from the IRWST can begin. ADS-4 [ performance is affected by the nature of flow in the hot legs. Successful operation of the ADS leads to the IRWST injection cooling phase of the AP600 SBLOCA event. The final stage of the SBLOCA is IRWST injection. At this point, the primary system is depressurized, and the transient continues into the LTC phase of the accident. By the time of IRWST injection, the CMT is either completely or very nearly empty. CMT phenomena have, therefore, become relatively unimportant, whereas the IRWST gravity-drain rate through the DVI line is important. He hot-leg flow phenomena, together with ADS-4 flow,is also important. Moreover, the break critical flow behavior is now less important than before because all ADS flow paths are open, providing a large area through which to vent steam. Keeping the core covered with liquid or a two-phase mixture becomes a function of the decay heat level and IRWST flow. De impact of noncondensable gas released when the accumulators empty of liquid during AP600 SBLOCAs is shown to be of low importance in the SBLOCA PIRT because of the large number of vent paths for the gas. 1.3.2 Long Term Cooling Transient LTC is a post accident phase defined as the period after IRWST injection begins until the plant is recovered. De AP600 passive safety systems are designed to provide post-accident core cooling indefinitely. Steam generated in the core is vented to containment. He steam condenses on the containment shell, and the condensate is directed into the IRWST and sump where it ficws into the core through the DVI line. De closed-circuit reflux condensation process ensures adequate cooling inventory to maintain the core in a coolable state indefinitely, mAap6000344w.non:lb 100395 1.3-3 REVISION: 1
When the reactor system is in the LTC mode, the primary system is drained to the hot-leg level. With the SG primary side being filled with stagnant steam, the pressurizer and upper head of the reactor vessel are empty. The CMTs and accumulators have already injected, and the PRHR may or may not be active, depending on whether the IRWST level covers the HX and noncondensable gas is present in the PRilR tubes. Inidal injection flow to the vessel comes from the IRWST as long as the IRWST head is larger than the containment sump. If the IRWST has drained to the sump level, there could be 1 injection from both the sump and the IRWST until the IRWST has drained. Flow from the sump or IRWST is directed to the reactor vessel thiough the DVI line into the downcomer. The delivery of injection flow is gravity-driven from the elevated sump water level into the reactor vessel. The driving force for core cooling is the level in the reactor downcomer, which provides the elevation head to drive flow through the core and out the hot leg. This gravity-flooding behavior of the core is no different than that in operating plants with the exception that injection flow is driven by a pump and not gravity as in the AP600. The inclusion of a large vent path on the top of the hot leg through the ADS 4 valves provides a low-pressure drop vent path so that ample flow through the core can occur. In operating plant cold-leg breaks, the downcomer must drive core flow through the SG primary side, superheating the primary fluid, and creating a backpressure that reduces core inlet flow (steam binding). This situatiore is avoided in the AP600 by using the large vent areas on the top of the hot legs so that very little, if any, flow goes through the SGs. Also, once the IRWST drains, the ADS 1-3 vent path is also available to vent core-generated steam. Since the primary system is at containment pressure, only the driving heads in the downcomer and the two-phase pressure drop in the core, hot leg, and ADS 4 determine the resulting core flow. If the PIRT is examined, only those items related to the core, downcomer, upper plenum, hot leg, and ADS 4 are highly ranked. While most items indicated on the PIRT for LTC are directly measured in the tests, the hot-leg flow regime must be inferred from the data. Using the OSU data, the importance of these phenomena can be assessed and used for guidance in validating the AP600 safety analysis computer codes for the LTC period. O maarnosuw.non:n, toows 1.34 REVISION: 1
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5 ~ A TA!!LE 151 h PIIENOMENA IDENTIFICATION RANKING TABLE FOR AP600 SBLOCA AND LTC TRANSIENT t
?
g e SBLOCA TRANSIENT LTC Transient y Component Initial Natural ' ADS , IRWST Sump N Phenomenon Blowdow n Circula5m Blowdown injectica Injection 5 Break Critical flow H H H M L Subsonic flow N/A N/A M M L ADS Stages I to 3 H* (inadvertent ADS) H* (madvertent ADS) H -M L Critical flow Two-phase pressure drop H+ H* H M L Valve loss coefficients H+ H* H M L Single-phase pressure drop H* H+ N/A L L ! Vessel / Core Decay heat H H H H H
,- Forced convection M N/A N/A N/A N/A Y3 Flashing M N/A M L N/A Wall stored energy M L M M M Natural circulation flow and heat transfer M M M M M Mixture level mass inventory H H H H H RCP RCP performance M N/A N/A N/A N/A Pressurizer '
Pressurizer fluid level M M M L L_ Wall stored heat M M M L L Pressurizer Surge Line Pressure drop / flow regime L L M L L Downcomer/lmwer Plenum L M M M M Upper Head / Upper Plenum L M M M H Cold Legs L M M M L The ADS is not normally opened during these phases unless the transient is an inadvettent ADS; for that case, the ADS phenomena would be ranked as high (II). W m b i z! i l
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s g TABLE 1.3-1 (Cont.) g PIIENOMENA IDENTIFICATION RANKING TABLE FOR AP600 SBLOCA AND LTC TRANSIENT t
? SHLOCA TRANSIENT LTC Trandent ! Component Initial Natural ADS IRWST Sump
{ Phenomenon Blowdown Circulation Blowdown Injection Injection
@ Steam Gener tor 3 20 - natural circulation L M L L L Steam generator heat transfa L M L L L Secondary conditions L M L L L Hot Leg Flow pattern transition L H H H H ADS-4 Critical flow N/A N/A H H L Subsonic flow N/A N/A L H H CMT Recirculation injection M M L L L Gravity draining injection N/A M H L L y Vapor condensation rate N/A M M L L Y CMT Halance Lines
- Pressure drop M H H L L Flow compositn>n M H H L L Accumulators injection flow rate N/A M H N/A N/A Noncondensible gas entrainment N/A N/A L L L 1RWST Gravity draining injection N/A N/A N/A H M Vapor condensation rate N/A N/A M L L DVI Line Pressure drop M M M M M PRHR Natural circulation flow and heat transfer L H M L L Sump Gravity draining injection N/A N/A N/A N/A 11 Level N/A N/A N/A N/A H p Temperature N/A l N/A N/A N/A H m
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.g 1.4 Test Facility Scaling A detailed component and system scaling analysis was performed for the OSU test facility and is given in the OSU Facility Scaling Report.* De results of the scaling analysis were used to specify.
j the design of the facility. The primary objective of the scaling analysis was to design a worldng scale - '~ model capable of producing the same types of flow behavior encountered in the AP600 during the SBLOCA transient, and LTC. i . . Various scaling techniques can be applied to the design of a small-scale thermal-hydraulic test facility, j . The traditional approach 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 Final Report.* Re optimum condition for this scaling approach occurs when the scale model implements ,
- - the same worldng fluid as the full-scale system,is built at full height using similar materials, and is
[ operated at full pressure. This generally results in constructing a very tall and thin scale model. Unfortunately, the hydrodynamic behavior in the plenum regions may not be fully represented in the l full-height model. A reduced-height, power-to-volume scaled model gives a better representation of
- j. multidimensional effects in the plenum and downcomer regions.*
I The hierarchical two-tiered scaling analysis (H2TS) method has been used to develop the similarity l- criteria necessary to scale the systems and processes of importance to AP600 integral system and LTC.
, The H2TS method, developed by the NRC, is fully described in Appendix D of An Integrated Structure and Scaling Methodologyfor Severe Accident Technical issue Resolution
- and is referred to as the SASM methodology. There are four basic elements of the H2TS analysis method. The first l- element consists of system decomposition. Each system can be subdivided into interacting subsystems (or modules), further subdivided into interacting constituents (materials), and further subdivided into l interacting phases (liquid, vapor, or solid). Each phase can be characterized by one or more geometric i' configurations, and each geometric configuration can be described by three field equations (mass, energy, and momentum conservation). Each field equation can be characterized by several processes.
His is depicted in Figure 1.4-1. After identifying the system of interest and decomposing it as in Figure 1,4-1, the next step is to identify the scaling level at which the similarity criteria should be developed. This is determined by the phenomena being considered. i For example, if the phenomenon being considered involves mass, momentum, or energy transport i between materials such as water and solid particles, then the scaling analysis should be performed at the constituent level. If the phenomenon ofinterest involves mass, momentum, or energy transport
, between vapor and liquid, then the scaling analysis should be performed at the phase level. Therefore, l- : Identifying the scaling level will depend on the phenomenon being addressed. Table 1.4-1 presents the j- system hierarchy implemented in the OSU Facility Scaling Report.m 4
[ mwaxn23m.m:lb loo 395 1.4-1 REVISION: 1 f
, . , , +- . - , . , . - - . .
f Thermal-hydraulic phenomena involving integral system interactions, such as primary system depressurization or loop natural circulation, are examined at the " system" level. Dermal-hydraulic phenomena-such as PRHR decay heat removal, CMT, accumulator, and IRWST paasive safety injection, automatic depressurization and LCS recirculation cooling-are examined at the subsystem level. Dermal-hydraulic phenomena important to individual components-such as the reactor core, pressurizer, SGs, hot legs, cold legs, coolant pumps, and interconnecting piping-are examined at the module level. Specific interaction between the steam-liquid mixture and the stainless steel structure are examined at the constituent level. The OSU scaling study presents scaling analysis performed at different levels. The thermal-hydraulic phenomena of interest, the system level at which the analysis was performed, the control volume for the analysis (i.e., the geometric configuration), the applicable balance equations, and the processes important to the thermal-hydraulic phenomena of interest are discussed and analyzed for the simulated reactor system as well as the major components in the system. e ne third element of the II2TS method requires the performance of a top-down (system) scaling analysis. The top-down scaling analysis examined the synergistic effects on the system caused by complex interactions between the constituents deemed important by the plausible phenomena identification ranking table (PPIRT). This has been modified as discussed in Section 1.3, and a revised PIRT is presented in Table 1.3-1. The top-down scaling approach used the conservation equations at a given scaling level to obtain characteristic time ratios and similarity criteria, and identified important processes to be addressed 11 the bottom-up scaling analysis. The fourth element of the II2TS method required the performance of a bottom-up (process) scaling analysis, which developed the similarity criteria for specific processes such as flow-pattern transitions, and geometry. and flow-dependent heat transfer. De focus of the bottom-up scaling analysis was to develop similarity criteria to scale individual processes of importance to system behavior identifled by the PIRT and develop the design information for the test facility. The basic objective of the ll2TS method was to develop sets of characteristic time ratios for the transfer processes of interest. This can be done by writing the control volume balance equations for each constituent, k, as follows: dV"y" = AlQ,ValtI),,A,, 1.4-1 dt Defining AlQ,y,l: AlQ.Vil = (QaV,1, - [Q Val.,,, i 1.4-2 O mAap600\2344w.non:IM00395 ],4 2 REVISION: 1
{}.
'v where:
l
% = Conserved property; p, pu, or pc (mass, momentum, or energy per unit volume) l V, = Control volume .
Q, = Volumetric flow rate l j,, = Flux of property % transferred from constituent k to m across the transfer area A., i 1
% transferred from constituent k to m across the transfer area A... Hence, A[Q,y,] represents the i usual mass, mornentum, or energy convection terms, and Ej,,A,, represents transport process terms such as condensation. ,
Equation 1.4-1 can be put in dimensionless form by specifying the following dimensionless groups in terms of the constant initial and boundary conditions: V[ = ",yl=Y' " Al,= " 1.4-3 k.0 Yk.0 Ql= h o. j[,= km""0 k km0 Substituting these groups into Equation 1.4-1 yields: A Q V,,og,o dVlyl
= Q,,og,oA [Qlyl]tI(j,,,oA,,,o)j[,Al, 1.4-4 dt Dividing both sides of this equation by Q,,o y,,o yields:
l dVly tk g, = @,NowR,r,A ', g W l 1 where the residence time of constituent k is: l v"- 1.4-6 t, = Oz.o and the characteristic time ratio for a transfer process between constituents k and m is given by: , J [I,,= lao ^ tao h 1,4 7 O t. ova.o >
/'
, mW344w.non:Ib too395 1.4 3 REVISION: 1 j
- y
l It is the il (pi) ratio of the proposed test facility to the plant that are ofinterest. Important processes can be replicated in the model by fixing the variables that control the process, such as geometry, so that the following criteria is met:
=1 1.4-8 TIE A deviation from unity indicates the possible deviation of the proposed test design from the plant.
The transients modeled at the OSU facility were SBLOCA transients that transition into the LTC mode for the AP600 design. Since the operating pressure for the OSU facility was chosen as 400 psia, a scaling approach was needed to develop the test design so that the most important parameters identified in the PIRT would be preserved. The SBLOCA and LTC scenario shown in Figure 1.4-2 indicates the five periods ofinterest. After the initial blowdown phase, there are extended periods of single- and two-phase natural circulation as the reactor system drains. Eventually, the ADS valves will open, creating a larger break, which will depressurize the primary system down to containment pressure. This is the ADS operational period. Once IRWST injection begins, there will be a two-phase natural circulation cooling mode with injection from the IRWST and venting from the ADS-4 valves located on the hot legs. This is the IRWST injection period. The LTC period begins as IRWST and sump injection continues for extended times. The LTC mode is with injection from the IRWST or sump and venting through the ADS-4 valves. A top-down scaling analysis was performed using the SASM methodology for both single- and two-phase natural circulation. The objective of the scaling analysis was to scale the steady-state single-and two-phase natural circulation flow rates and the natural circulation heat transfer. A bottom-up scaling was then performed to develop the similarity criteria to specifically scale the core and SG heat transfer regimes, flow regimes and transitions, frictional and form pressure losses, and critical heat ilux. To maintain similarity, the Il values developed from the dimensionless conservation equations should be preserved or the ratio of the groups should be unity. The scaling study requires the user to choose a length scale and an area or diameter scale for the facility to satisfy the system of equations, power requirements, and geometric representation of the facility rciative to the plant. Other scaling considerations such as flow regimes in the loop piping must also be considered. Small diameters distort the flow regime and have different transitions between the flow regime compared with the prototype. Small-diameter pipes also have different two-phase counter-flow behavior compared with the prototype. Reduced size can also cause manufacturing problems for the core heater simulators. An evaluation determined that a 1/4-length scale was the most appropriate for the OSU facility since it minimized the power requirements while maximizing the height. A l/4-scaled facility also had rnAap600c344w.non:Ib 100395 1,4-4 REVis!ON: 1
, sufficient volume and size to correctly model the plant pressure drop and possible three-dimensional
- flow behavior that could occur in the simulated reactor vessel, plenums, and downcomer To chose a l - consistent diameter scale, a simple relationship was derived from the one-dimensional momentum
- equation to relate the length ratio to the diameter ratio. The choice of the diameter ratio was further j verilled with a bottom-up scaling approach in which the two-phase flow regimes and transitions between flow regimes were examined using the work of Taitel and Dukler.* Re possible distortions -
in the flow regimes and their transitions was also examined for the horizontal piping following the j- approach of Schwartzbeck and Kocamustafaogullari.* Re flooding review by Bankoff and Lee" was also used to verify that the chosen diameter ratio would have minimum surface tension effects if i flooding occurred. Using this approach, the facility dimensions could be specified with confidence l that the key parameters and phenomena identified in the PIRT would be preserved in the OSU facility so that the resulting data could be used for AP600 safety analysis code validation. i . The OSU tests were designed to start in an all liquid recirculation mode with the simulated RCPs I operating with system pressure at about 400 psia. When a break is initiated, the system begins to , depressurize. To preserve the depressurization behavior of the OSU facility, a reference pressure was selected and a scaling rationale developed to relate the lower pressure OSU tests to the higher-pressure . AP600 transient for the depressurization transients. He results of the scaling approach were used to develop the relationships that led to selection of the OSU facility break areas, ADS valve areas, J 3 accumulator gas pressure, and SG secondary-side safety pressures to preserve the scaling relationships between the facility at its reduced pressure and the AP600 plant at its higher pressure. The scaling process used is shown in Figure 1.4-3, in which a top-down scaling approach was used to develop the systems scaling analysis for a simplified control volume of the reactor primary system. A bottom-up approach was then used to develop the fluid property relationships for the depressurization transients. , i The approach, originally developed by Kocamustafaogullari and Ishii" and expanded on by Moskal," was extended to relate the OSU fluid property conditions to the AP600 plant conditions. Moskal defines the property relationship: Y= M 1.49 ) P ,P htq 4 as the key property group to be preserved. His particular grouping also appears in the coefficients for y the core velocity from the two-phase natural circulation loop scaling analysis previously described.
! De fluid properties and depressurization approach is to select a reference pressure for both the OSU i
} facility and the AP600 that will capture the important parameters identified in the PIRT, Examining
- Figure 1,4-2, the AP600 primary system pressure will stabilize, after the initial subcooled blowdown, J to a near constant value, slightly above the safety valve setpoint for the SG secondary side De )
primary pressure will remain at this value for a relatively long period, depending upon the break size l i and when the ADS activates, which will depressurize the primary system to the containment pressure. During this time period, the passive safety-related systems of the AP600 will be in operation and the phenomena ofimportance, which are identified in the PIRT, will be present. Herefore, the m:Whnon:lb 100395 ],4 5 REVISION: 1 1
secondary-side SG safety valve setpoint pressure was chosen as the reference pressure for the AP600 plant. Note that this ignores the subcooled depressurization portion of the transient which, for a SI3LOCA, is a short period compared with the total transient length. A similar reference pressure can be chosen for the OSU facility where the primary pressure stabilizes above the SG pressure, so that: _ = _ l.4-10
.Yo_, , Yo, g where:
Yo], = OSU reference pressure Yo], = AP600 reference pressure The top-down and bottom-up pressure scaling must also be consistent with the natural circulation scaling, which establishes the facility volume, time, and velocity scaled ratios given the selection of the length and diameter for the facility. 'The bottom-up pressure scaling examined the critical flow through the break and the ADS valves, and ' developed the relationships for the break areas and the valve areas that were consistent with the fluid property scaling given above. Therefore, given a break size in the AP600, a corresponding break size can be calculated for the OSU facility that will maintain the time, velocity, and volume scaling for a selected length and diameter scale which was chosen from the two-phase natural circulation scaling relationsh:ps. O m:\aguhA2344w.non:lb-100395 1,46 REVISION: 1
__ . .. - .. . = . - - . - . . .. - . .. . - . . - . .
- ; \g)
TABLE 1.41
- GENERAL SYSTEM IIIERARCHY
OSU/AP600 SCALING ANALYSIS SYSTEM: Pnmary loop SUBSYSTEMS: PRHR, CMT, IRWST, accumulator, ADS, LCS recire. system MODULES: Reactor core, pressurizer, steam generators, hot legs, cold legs, reactor coolant pumps, j interconnecting piping 1 .. CONSTITUENTS- Steam-liquid mixture, stainless steel structure h j i 1 E a i r J i 1 m:wan2344w.non:1b-too395 1,4 7 REVISION: 1
O SYSTEM (S) S SUBSYSTEM (SS) SSI SSk MODULES (M) 'M1 M2 CONSTITUENTS (C) C1 Ck n SYNTHESIS PH ASES (P) K f 5 q ANALYSIS _ _ _ _ _ _ _ i _ _._.___ __ GEOMETRICAL U CONFIGURATIONS (G) 61 62 Ok FIELDS (F) M E MM PROCESSES O O Pk Figure 1.41 Decomposition Paradigm and Ilierarchy m m:\ap60(A2344w.non:lt>100395 1.4 8 REVISION: 1
i !O 4 i i 4 i i i AP600 SBLOCA Scenario
- l SBLOCA l:: LTC :
1 I I I
- - 4 ,
Blowdown l ADS l lRWST l Sump 4 i i Blowdown Injection i Injection 1 i i I 8 , , , j i I i I 4 e- , , , , ! 5 l Natural l l l
' o "'' " ' '
.O E a- . , , 1 i I I e : i I i i 1 j i i I I I i 1 1 I I I I I I I I I e e i e i I I I I I _ I I I I I I I I I 518978.1 O ,I t 1 i 1 1 ] 4 4 1-Figure 1.4 2 AP600 SHLOCA Scenario m:Wmw.non:tbioo995 1,4 9 REVISION: 1 j - ..
O SYSTEM DEPRESSURIZATION PHENOMENA I 1 1 Top Down/ System Bottom Up/ Process Scaling Scaling
- Fluid Property Scaling
* * "I
- System Depressu:ization Rate fg ,,'"
- o hase Natural l r De ay Power
- Component Stored Energy
- team Generator ScaHag
. Net System Power
- Pressurizer Scaling
- PRHR Scaling System Depressurization IIGroups and Similarity Criteria u
Evaluate Scaling Distortions v Pressure Setpoints. Core Power and ADS / Break Size Specifications Figure 1.4-3 Scaling Analysis Flow Diagram for System Depressurization m:wisoom4w.non:tb-too995 1.4-10 REVISION: 1
o (/ 1.5 Test Scaling Assessment and Dimensions To assess the scaling of the OSU facility, both the proposed scaled facility and the AP600 were modeled using the NOTRUMP, A Nodal Transient Small Break and General Network Code." The objective of this study was to investigate whether the OSU facility response to a small-break transient would be similar to the response of the AP600. A 2-in. cold-leg break on the CMT side of the plant was selected. The OSU facility initial conditions for this break are shown in Table 1.5-1. Table 1.5-2 shows the scaling relationships between the plant and the OSU facility, which account for the time scale difference (two-to-one for OSU) and the normalization of flow, pressure, two-phase mixture levels, and total system mass. Figure 1.51 compares the normalized pressure transient for the plant and the OSU test facility and indicates that reasonably good agreement was achieved. The reference pressure chosen is based on when the primary pressure stabilizes above the secondary-side pressure at time,ot . The normalized CMT levels are shown in Figures 1.5-2 and 1.5 3, and the normalized accumulator levels are shown in 4 Figure 1.5-4 and 1.5-5. These figures are in good agreement and indicate that the scaling correctly preserves the timing of the events for the OSU facility compared with the AP600 when the time scaling logic is applied. ADS 1-3 flow is shown in Figure 1.5 6 and is in reasonable agreement between the test facility and the plant. Break flows are compared in Figure 1.5-7, and a difference between the facility and the plant is indicated. One possible explanation for the difference is that, o when scaling the critical flow area for the break, a quality of the flow must be assumed. In reality, 1 ~h the quality of the flow at the break is not a constant and will change with time. Furthermore, exact similitude cannot be simultaneously achieved for both the break energy flow rate and the break mass 1 flow rate with a reduced pressure scale. Flowever, the integrated mass inventory similitude can be preserved. Figure 1.5-8 shows the normalized mass inventory for both the test facility and the plant, j Again, agreement between the two calculations is very good, indicating that the scaling approach will yield thennal-hydraulic phenomena similar to the AP600. 1 In addition to the NOTRUMP code calculations that compared the OSU and the AP600 response, the l results from the SPES-2 and OSU tests can be compared to investigate the scaling performance of the OSU facility. Comparisons of a 2-in. cold-leg break from SPES-2 Test S00303* have been made with OSU Matrix Test SB01. The derived OSU scaling factors were applied to the SPES-2 results to compare time, pressure, and ; flow rates. The OSU time scale was multiplied by a factor of 2. The OSU pressure scale was i normalized using the reference pressure (maximum pressure on secondary side). Similvly, the SPES-2 I pressure scale was normalized using the reference pressure for the test. The flow rate normalization j factor in SPES 2 was the maximum flow rate for the process being examined. For purposes of l comparison, the flow rate normalization factor in OSU was the maximum flow rate observed for the identical process in SPES-2 multiplied by the ratio 395/96. Thus, the flow rates can be compared on a i similar basis. ;
/% i V
m:\ap600\2344w. son:lb-100995 1.5-1 REVIslON: 1
A 2-in. cold-leg break was simulated in both the OSU and SPES-2 facilities. He break location for these tests was the tmttom of a single celd leg. Each system was at its steady-state initial condition at break initiation. Subsequent depressurization behavior was recorded for each facility, and key data plots are presented for the purpose of comparison. De vertical axis of each graph has been r3ormalized as described previously. Figure 1.5-9 presents a comparison plot of the SPES-2 and OSU reactor vessel pressure histories and Figure 1.5-10 through 1.5-14 present the data comparisons for the key passive safety systems. In general, the data comparisons for the 2-in. break case indicate good agreement. De tindng of key events, such as ADS valve actuation, were preserved. One difference can be identified in Figure 1.5-14, however, where it is observed that the onset of IRWST injection was delayed in the OSU facility relative to SPES-2. He difference in the IRWST injection time is due, in part, to the oversizing used in the SPES-2 facility. Similar comparisons were perforn ed for the DEG DVI line break, and the agreement between OSU and SPES-2 was very good. Rese comparisons of a full-height, full-pressure test facility with the OSU reduced-height, reduced-pressure facility support the scaling logic used in the OSU test design. As with any scaled test, scaling distortions are unavoidable, therefore, the purpose of the PIRT is to identify the most important processes to be scaled so that the system response is most prototypical. In the OSU Facility Scaling Report,* Section 10.2 specifically discusses scaling distortion for the facility, and an evaluation is made of the scaling distortions on the primary parameters developed from the PIRT. De results are provided in Table 1.5-3. As the table indicates, all of the important pi ratios are within 20 percent, which is acceptable in terms of the uncertainties for scaling of the tests. The depressurization ratio given in the table incorporates the use of the revised depressurization scaling, which uses the system energy and volumetric scaling as a basis. This portion of the scaling was revised after the completion of the Westinghouse tests at OSU. De original scaling for the breaks, which was used for the tests in this report, used the break-dominated depressurization process described in Section 5.4.1 of the Facility Scaling Report.* nis scaling approach was used for all breaks modeled in the OSU tests given in this report as well as the OSU Final Data Report.* Dis scaling process is valid for break sizes of 2 in. or more, but was found to be inaccurate for smaller breaks. Derefore, the 1-in, cold-leg break and the 1/2-in. cold-leg breaks are oversized when considecing the system energy and volumetdc scaling methodology. He ratio of the break diameter scale factors between the two methods is given in Table 5-6 in the Facility Scaling Report,* and is 1.5. The break areas used for the Westinghouse tests were greater than the scaled values; therefore, the timir.g of the events for these 1-in. and 0.5-in. tests will be distorted from the properly scaled values, and the events and total transient will be shorter than the revised scaling would predict. Revision of the scaled break diameters is the only significant re-scaling for the OSU test facility. This distortion only affects a few tests; the remaining tests have tne properly scaled break diameter. Data for the affected tests are still suitable for the purposes of computer code validation, since the break area used in the test can be simulated in the code prediction of the test. maapsoochoon:tb-ioon5 1.5-2 REVIslON: 1
4 N TABLE LS I INITIAL CONDITIONS FOR OSU TEST FACILTn' TO MODEL A 2 IN. COLD-LEG BREAK
; Heactor Cooling System 1
Core Power 0.700 M Wt Core Flow 116.7 lb/sec. Pressurizer Pressure 400 psia Core Inlet Temperature 410.4'F Core Outlet Temperature 415.6'F Secondary SG Temperature 407.6*F Break Size Simulated 2-in. cold-leg break 4 0 .i I s (G l l l 1 m:W344w.no :1b-toons 1.5-3 REVISION: 1 l
TABLE 1.5-2 O SCALE FACTORS TO RELATE TIIE AP600 PLANT TO OSU NOTRUMP CALCULATIONS AP600 OSU Time t-tf) 2'(t-Q Pressure P/1080 P/320 Flow W/96 W Note:
% is the reference time when the primary pressure stabilizes above the secondary-side pressure.
O O m:W344w.non:Ib-loo 995 1.5-1 REVISION: 1
1 a TAlti E 1.5 3 DISTORTION f.WT2 VOtt 1DF, Al%00 DOMINANT PROCESSES IDENTIFIED NNG THE H2TS METHODOLOGY
. . w s ,. - -
Distortion Factor Characteristic (DF) Operational Time katio (%) Mode Um 0
- l$ natural circulation D, 0
- l$/2$ natural circulation l H. 0 .
2$ natural circulation with fluid
, property similitude /LCS recirculation H. (Not scaled)
- 2$ natural circulation pressured
- scaled core void fraction preserved instead of R.
Dr/C, O e Depressurization in. cold-leg break (energy dondnated) Dr.wr 17 4
- CMT draining with hot walls H. 2.8 . CMT draining with cold walls Duc n.a 6.3 = CMT draining with cold walls O.2,wn 0 + IRWST draining (property similitude)
H,; awn 0 . IRWST draining (property similitude) D.; awn -9.5 + IRWST beat up (pressure scaled) H aanwn 19.1 = IRWST beat up (pressure scaled) D,oc 16.7
- Downcomer beat transfer during Accumulator injection D, TBD* + Sump filling and recirculation AP600 data not available
*TBD - To be determined O
m:Whoon:ib-tom 95 1.5-5 REVISION: 1
O
.. ,AP800 / OSU NORtVLUZED RCS PRESSURES PFN 9 OSU NORW. PRESSUNL
_ . PFN 9 AP600 NORW. PRESSURE 1.4
~
\
1.2
~
j
;- - n N-
~
. \
~
P m * ( P,,, .
.6 L
,a
.4
.2 0 ' '
-500 0 500 1000 1500 I
Reference time (AP600 time) Figure 1.51 Normalized Pressure Comparisons between AP600 and OSU Facility mvm2mw.no.:it>1oo995 1.5-6 REVISION: I
O ape 00/ OSU NORWdRED CalT1 EVELS EWlXSFN 56 OSU CMT FILL
- .EWlXSFN 56 AP600 CNT FlLL 1 - - - - - -
1 8 3 16 - H ( aml . 6 O -
.4 s N
~
l 1 2 1 l
' 1
- n o , , , , .
-500 0 500 . 1000 1500 l
- Reference time (AP600 time)
- Figure 1.5 2 Normalized CMT-1 Level for AP600 and OSU Fadlity m:wp6000mw.non:Ib too995 1.5-7 REVISION: 1
O AP900/OSU NOIMAAllZED,CMT2 LEVELS EWlXSFN 66 OSU CMT FILL _ .EWlXSFN 66 AP600 CMT FILL 1 - _ _ _ _ _
~
i
~
l 8
~
Hg
.6
~
/j g 4
2 0 ' ' ' ' s's -500 0 500 1000 1500 Re'ference time (AP600 time) Figure 1.5-3 Normalized CMT-2 Level for AP600 and OSU Facility mWmw.non:tb-ioo995 1.5-8 REVISION: 1
f ape 00 / OSU N0pual wn ACCUuuLATDR 1 LNELS EWlXSFN 51 05U ACC FILL
. . EWlXSFN 51 AP800 ACC FlLL 1
i _a* H O _ p . 4
. .c
- 2 t
me i ' ' ' ' ' ' ' ' ' ' ' ' 0 '
-500 0 500 1000 1500 Reference time (AF500 time) i
[ t Figure 1.5-4 Normalized ACC 1 Level for AP600 and OSU Facility m:vm2344w. on:ivioo995 1.5 9 REVISION: 1 e
O N/OSU pas rysnNTOR2 LEVELS EWlXSFN 61
. EWlXSFN 05U ACC FILL
. 61 AP600 ACC FILL 1
g 4 e
*2
. = - 0 ' ' ' ' ' ' ' ' i
-500 0 500 1000 1500 Reference time (AP600 time) 4 Figure 1.5 5 Normalized ACC-2 Level for AP600 and OSU Facility m:pmw.non:ibioo995 1.5-10 RE\ GION: 1
4 i 1 O 1 l i Apeoo/OSU N0punie ads 14 nows WFL 58 OSU REFERENCE FLOW ! . .WFL 58 AP600 NORM. FLOW l l 12 i - l 10 j 3 )ll I w 6 iO 1
- l l 4-l .
e ..
, i 2 -
I o 0
-2 ' ' ' ' ' ' ' ' ' ' ' l
-500 0 500 1000 1500 l Reference time (AP600 time) 4 O Figure 1.5-6 Normalized ADS 13 Flows for Al%00 and OSU Facility i mhp600cu4w.non:1 M 00995 1.5-11 REVISION: 1
. . _ . . _ . __ __ _ - - . ~ - - . _ .,
I 1 . I l 9 I i ! AP900/OSU NORalAll2ED BREAK R.0WS WFL 80 OSU REFERENCE FLOW i ; ,WFL 80 AP800 NORM. FLOW f 8 si i - t
- i 5
^-J 4 VO W
i 4 3
- - b I O i
A 2-M -
~
k ug ver
- I
(
' ~
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' -500 0 500 1000 1500 1
Reference time (AP600 time) Figure 1.5-7 Normalized Break Flow for AP600 and OSU Facility i m:\ar6002344..no :ivia995 1.5-12 REVISION: 1
O 1 AP900 / OSU NORMAUZED SYSTEM MASS MSTONES WASS 70 OSU NORW. WASS
. .WASS 78 . . AP600 NORW. WASS 14 1.2 i u 1 Me
.8 l .6 l l
l X= 4 4 'I
- V V l j -
1 ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' '
- 2
-500 0 500 1000 1500 Reference time (AP600 time)
O Figure 1.5 8 Normalized System Mass for AP600 and OSU Facility 4 maapem2mw.noa:tb-iow95 1.5-13 REVISION: 1
i
} 134 SPES22lNCHBREAKvsOSU2iNCHBREAK 0 0 CMT A injection he (SPES) y e F_A40E g ,5 s- - - - FMM_504 119 0 0 CMT-2 Injection Flow (osu)
V R j 1 .
- u -
\ ,,
6 - l 9 ,
.! ik I X .,
au .
-l
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{ -l I 1
'+h n
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+ sg yA _i , ;
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8 - 3 - I - t i 1 i e i i e i e e i e e i t g 0 1000 8 00 3000 4000 2 Time i,si g . e - e _ __ e
i f O .U I I _ ! l . i iR
, )
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se x-- 1 _ - : l
,, / -:
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j l i u _ l I
; l I I f i e l.1L 1 I I f f I f I l i I I I I g I
.. eg me <g es o (q g
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N (' Figure 1.5-10 Comparison of OSU and SPES-2 2-In. Break Pressure Histories map 60m2344w.noa: b ioo995 1.5-15 REVISION: 1
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M [ ,- (,j ' 2.0 FACILITY DESCRIPTION
SUMMARY
*Ihe OSU low-pressure integral systems effect test facility is a scaled representation of the AP600.
The design operating conditions for the facility are 400 psig at 450*F. The facility includes all the passive safety-related injection systems that appear in the AP600 design, as well as nonsafety-related injection systems such as the normal residual heat removal system (RNS) and chemical and volume control system (CVS). The test facility operated in a steady-state fashion with a maximum electrical power of 660 kW, using four primary system recirculation pumps and two SGs. SBLOCAs are simulated using break spool pieces. The break can be located on the hot legs, cold legs, DVI lines, or cold-leg balance line. The test facility is designed to simulate the scaled decay power of AP600. .p ij .( b m:w344 21.on:isioo395 2-1 REVISION: 1
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( 2.1 Overall Facility Description De OSU test facility is a scaled model of the AP600 reactor coolant system (RCS), steam generator -
- system (SGS), passive core cooling system (PXS), automatic depressurization system (ADS), lower ')
containment sump (LCS), chemical and volume control system (CVS), and normal residual heat -) i removal system (RNS). In addition, the facility is capable of simulating the AP600 passive containment cooling system (PCS) condensate return process. Figure 2.1 1 is an isometric drawing of the test facility, and Figure 2.1 ~2 is a simplified flow diagram of the test facility. The facility reflects . the scaled AP600 geometry, including the piping routings. All components and piping are fabricated from austenitic stainless steel. - De relative locations of all tanks and vessels-such as the IRWST, } CMTs, and accumulators were maintained as determined by the scaling approach. The facility uses a unique break and ADS measurement system (BAMS) to measure two-phase break and ADS flow.
]
$ ~ The RCS is composed of a reactor vessel, which has electrically heated rods to simulate the decay heat
- lin the reactor core, and two primary loops. Each primary loop consists of two cold-leg pipes and one hot-leg pipe connecting a SG to the reactor vessel. A reactor coolant pump (RCP) on each cold leg takes suction from the SG channel head (downstream of the SG U-tubes) and discharges it into the downcomer region of the reactor vessel. A pressurizer with an electric heater is connected to one of the two hot legs through surge-line piping. The top of the pressurizer is connected to the ADS 1-3 line. An ADS-4 line is connected to each hot leg.
He reactor vessel contains two DVI nozzles that connect to the DVI lines of the PXS. A flow venturi is incorporated in each DVI nozzle to limit the loss of inventory from the reactor vessel in the event of a double-ended DVI line break. !~ i , This test facility models the primary and secondary side of the SGs with one generator per primary . loop. A simulated feedwater line is used for each loop to maintain proper secondary water level. The steam produced in each generator is measured and exhausted to the atmosphere through a common diffuser and stack. The test control logic simulates the response of the AP600 by providing an S signal at a fixed time following a break. The passive safety injectior, systems consist of two CMTs, two accumulators, one IRWST, and one , passive residual heat removal heat exchanger (PRHR HX). He test facility simulates the AP600 IRWST with a cylindrical tank with scaled water volume and height. The IRWST is located above the reactor core; two injection lines connect to the two DVI lines. Each IRWST injection line also connects to the sump tank with interconnecting piping and isolation valves. The PRHR HX is located i inside the IRWST, using IRWST water as the heat sink. The inlet of the PRHR HX is connected to the pressurizer-side hot leg via a tee at the ADS-4 line, and the outlet is connected to the SG channel head at the cold-leg side. Since the inlet is hot and the outlet is cold, water is circulated through this
- . system by natural convection. De water volume and elevation of each CMT are scaled and modeled.
i mAmp600G3h-2a.non:1b 100395 2.1-1 REVISION: 1 l l
They are elevated above the reactor vessel and the DVI lines. A line connecting the top of each CMT to its cold leg provides pressure balance between the RCS and the CMT. Therefore, the CMT injects cooling water by its own elevation head. The accumulators are also modeled with scaled volume and height. Ilowever, they are pressurized with nitrogen and, therefore, inject when RCS pressure is below the preselected scaled accumulator pressure. The AP600 uses four stages of valves to depressurize the RCS. The first three stages of the ADS are provided through connections to the pres.surizer. These three stages are arranged in parallel, with each stage containing two lines with each line containing an isolation and control valve. The fourth stage of the ADS contains four separate lines. The OSU test facility uses only one set of valves to model the ADS 1-3 stages for AP600. This is done using removable flow nozzles to match the scaled flow characteristics of either one or two lines of valves. The lines of ADS 1-3 split into parallel lines from one connection off the pressurizer in the AP600. The discharge lines from the ADS 1-3 valves are joined into one line connect < :1 to the ADS 1-3 separator. This two-phase flow is separated using a swirl-vane separator. The liquid and vapor flows are measured to obtain the ADS total flow for mass and energy balance analysis he separated flow streams are then recombined and discharged into the IRWST through a sparger. Re OSU test facility uses one ADS-4 line connected to the top of each hot leg. Each line contains a O pneumatically operated, full port ball valve acting as the ADS-4 isolation valve and a flow nozzle simulating the flow area in the AP600. Two sets of ficw nozzles are used in the test: one simulates 100-percent flow area and the other simulates 50-percent flow area. In the test facility, ADS-4 discharge flows to the ADS-4 separators, where the steam and liquid flows are separated and measured. Steam flow is measured and exhausted to the atmosphere. He lower containment sump in the AP600 consists of two volumes: normally flooded and normally nonflooded. The normally flooded volume consists of those compartments which collect liquid break flow and ADS flow. The normally flooded volume is modeled in O3U with a cylindrical tank, identified as the primary sump tank. The normally nonflooded volume includes those compartments which do not collect any liquid flow. The normally nonflooded volume is modeled in OSU with a cylindrical tank, identified as the secondary sump tank. These two tanks are connected with a hne at a level simulating the curb level in the AP600. In the AP6(X), the RNS is used to provide nonsafety-related cooling water injection to the reactor core. He RNS pump takes suction from the IRWST and discharges it into the DVI lines. The test facility RNS pump takes suction from the IRWST at the scaled location and elevation and discharges equal flow to both DVI lines at scaled heations. O m%umwaine-im395 2.1-2 REVISION: 1
1
)
BAMS is uniquely designed for the test facility to indirectly measure two-phase flow and energy leaving the break and ADS location. His system separates two-phase flow into single-phase liquid - and single-phase steam flows for direct fiow rate and temperature measurements. De BAMS consists of steam liquid separators and the interconnecting pipes and valves to the various break sources, the primary sump tank, the ADS 1-3 lines, and the main steam header. - Two-phase flow (steam and water) from the ADS 1-3 lines enters the ADS 1-3 separator, where the steam is separated from the mixture. Steam flows out of one outlet while liquid drains down the other. Dese two lines recombine the separated flows' downstream and discharge into the IRWST via the sparger located inside the IRWST, herefore, the mass and energy from ADS 1-3 is transferred to the IRWST as in the AP600. Two ADS-4 separators are used, one for each ADS-4 line. Each ADS-4 separator separ.tes two-phase flow into single-phase steam and single-phase liquid for flow rate, pressure, and temperature f measurements. The steam line connects to a common steam header, and the liquid line connects to the primary sump tank. Rese connections simulate the ADS-4 operation process in the AP600, where the steam flow rises to the containment wall and liquid drains to the sump.
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I ~ l f V. 2.2 Facility Instrumentation 1 l OSU 'used standard instrumentation to indicate mass, flow, temperature, and pressure in the system and l components. He instrumentation plan was specifically designed to provide a transient mass and energy balance on the components and for the entire system. Details of the instrumentation and its performance can be found in the OSU Final Data Report.* Re function of each of the instruments is discussed below. The use of the instruments for the mass and energy balances is provided in Section 4. 2.2.1 Differential Pressure Transmitters (FDP, LDP, DP) Differential pressure transmitters measure the flow (FDP), or the level (LDP), in the system. The only application of FDPS is to measure differential pressure across the flow orifices in the ADS 1, j ADS 2, and ADS-3 lines (FDP-604, FDP-605, and FDP-606, respectively). De level (LDP) transmitters measure levels in the facility tanks, the RCS hot-leg and cold-leg pipes, SG tubes, and PRHR H.X tubes. He LDPs were designed to measure level (mass)in a component . and were calibrated at ambient temperature. He level data recorded by the facility's data acquisition i system (DAS) were uncompensated for temperature. De resulting signals were temperature
- O compensated for the mass and energy balance calculations.
0 2.2.2 Pressure Transmitters Pressure transmitters (Frs) are identical to differential pressure transmitters except that the low-pressure side of the transmitter senses atmospheric pressure. 2.2.3 Magnetic Flow Meters Foxboro" magnetic flow meters (FMMs) were used to measure liquid flow, in the different liquid solid lines. The FMMs are not designed to accurately measure steam or two-phase flow, and the data from the transmitters are invalid when either of these are measured. Foulteen Foxboro" vortex flow meters (FVMs) were used to measure steam flow in the test facility. The FVMs measure steam flow from the ADS 1-3, ADS 4-1, ADS 4-2, and break separators. In addition, they measure steam flow from the primary sump, the IRWST, and the BAMS header. These meters are known to have a manufacturer's warranted cutoff of 141 actual cfm (acfm) which means ) they may not detect flows below 0.088 lbm/sec. at ambient conditions. l l' m:\ap600\2344w-2a.non:Ib 100395 2.2-1 REVISION: 1
1 2.2.4 Ileated Phase Switches Heated phase switches (HPSs) manufactured by Reotherm"' were used to indicate fluid phase in different components. There are 12 switches: one each on the cold and hot legs, CMT balance lines, PRIIR IIX inlet, and ADS 1-3 header, in addition, two switches are installed in the pressurizer surge line. The llPSs were usually used in conjunction with an LDP to indicate a separate level in a component. 2.2.5 IIeat Flux Meters lleat flux meters (HFMs) were used to measure heat flux through pipe or tank walls to indicate heat loss. The small, wafer-thin instruments are glued to a pipe or tank surface. Bree thermocouples are imbedded into each HFM. Two thermocouples measure temperature on either side of the HFM. The thermocouple signals are measured by the DAS, and their temperature difference is converted to a heat flux using coefficients provided by the vender. The third thermocouple measures the temperature of the surface. 2.2.6 Load Cells The mass of water in the IRWST, primary sump, and secondary sump was measured by load cells mounted under the four supports of each tank. After the transmitter load cell was calibrated, it measured only the weight of water in the tank. The transmitter also provided local indication of weight in the tank for use by test yrsonnel. The load cells were found to be sensitive to variations in , ambient conditions, and thus, are not used for absolute measurements. Their readings are valid for l tracing short time scale variations in mass. 2.2.7 Thermocouples Thermocouples are assigned one of four instrument designations, depending on the thermocouple's application. A TF thermocouple inserted through the wall of a pipe or tank is mounted on a thermocouple rod. TW thermocouples are mounted on the inside or outside walls of a tank or pipe. TR thermocouples, unique to the reactor vessel, are mounted on vertical thermocouple rods installed in , the reactor vessel. TH thermocouples are mounted on the heaters for the reactor vessel and the l pressurizer. Thermocouple type and diameter are specified in Appendix C of the OSU Final Data Report.") The reactor vessel contains TH thermocouples to measure temperatures of selected heaters. Selected heater thermocouples are used as inputs to the safety shutdown of the reactor heaters to detect abnormally high temperatures. Thermocouples mounted in hollow rods (TR thermocouples) are unique to the reactor vessel. Five thermocouple rods are installed in the reactor vessel to provide radial and axial fluid temperature mAap600\2344w-2a.non:1b 100395 2.2-2 REVISION: 1
l
/ - distributions in the heated section of the reactor vessel. Each rod contains thermocouples mounted
, along its entire length. Thermocouples protrude from the hollow rod and are sealed from the outside wit:i silver solder.
The CMTV are instmmented with numerous fluid and wall thermocouples to measure the CMT wall heat flux and temperature, as indicated in Appendix G, Dwg. OSU 600501 and 600502 of the Final Data Report.m Each CMT contains one long and two short thermocouple rods instrumented with thermocouples along its entire length. In addition, inside and outside wall thermocouples, fluid thermocouples installed 1 in. from the inside wall, and tank centerline thermocouples are installed at the same elevation to measure the temperature of the fluid and walls at that elevation.
'Ihe IRWST also contains two thermocouple rods to measure the energy gain in the IRWST as shown
'in Appendix G, Dwg. OSU 600701 of the Final Data Report.m One long and one short tube of each SG are instrumented with shell-side (secondary-side) wall thermocouples and tube-side (primary or RCS-side) fluid thermocouples as shown in Appendix G, l Dwg. OSU 600301 of the Final Data Report.m a
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m:W23w-2a.no :tb ioo395 2.2-3 REVISION: 1
4 1 l l l (O) 3.0 TEST
SUMMARY
P 1he 'ollowing sections describe the test matrix and the data validation processes used for the low-pressure integral systems tests performed at OSU. The initial test validation process was based on i specific instruments that were required to function or selected backup instruments that were available. The OSU Final Data Report
- describes the instrumentation from the tests and assesses their performance and reliability.- This report utilizes the valid instrumentation to interpret the data.
J l l l l i l msp60m23n3w.noa:1bioo995 3-1 REVISION: 1
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- 1. 3.I' Test Validation l
As described in the OSU Final Data Report,* the OSU test facility data were reviewed and validated using a three-step process. The first step was performed at the OSU test facility. Immediately l following each test, the data were recorded on a compact disc as a read-only file, and documented in a Day-of Test Report. The Day-of-Test Report evaluated the test from a very basic standpoint, including ! operability of_ key instruments and deviations from specified initial conditions. The Day-of-Test Report also documented any facility modifications or onsite test observations. The Day-of-Test Report assessed whether the test needed to be rerun because of some significant problem observed during the performance of the test. The overall test acceptance criteria are shown in Table 3.1-1. The critical instruments were the minimum set ofinstruments required to perform a t/ansient, component-by-component mass and energy balance. The second step in the data validation process was performed by the test engineering personnel at the l Westinghouse Energy Center in Monroeville, Pennsylvania. This step was performed after receiving I the Day-of-Test Report and processing of the data. This data validation was documented in the Quick l Look Report (QLR). The QLR prosided a preliminary validation of all test data. The key purpose of I the QLR was to issue some pedigree of the data, without specifically evaluating the data for code j validation purposes (reviewed, but not yet validated, data) shortly after the test was performed. The third step was a detailed review of the transient progression, facility and component performance, and cross-test comparisons as reported in the Final Data Report.m j This report utilizes the insight derived from the instmmentation performance and the transient progression gained in the Final Data Report
- to determine and understand the thermal-hydraulic j behavior of the facility during each test. The objective of the analysis is to provide the explanation of the test facility response. l I
l l l \ m%sxmu-3...onab.noo995 3.1-1 REVISION: 1
TABLE 3.1-1 O OVERALL ACCEL'TANCE CRITERIA
+
Test initial conditions shall be achieved in a specified tolerance.
- Setpoints shall be achieved in an acceptable tolerance band.
+
Sufficient instrumentation shall be operational before the test (exceptions shall be approved by the Westinghouse test ecgineer),
+ Critical instruments not operating shall be identified to the Westinghouse test engineer before the j tests. These instruments must be operational before and during the test, or exceptions should be l approved.
- A zero check of LDPs, DPs, and FDPS shall be in acceptable tolerances.
The zero check was eliminated from tbc acceptance criteria for Category III tests. The earlier pre-test and post test checks of zero shift showed acceptable variation in the readings of these instruments. Performing these checks required that each inr ament be manually isolated and then returned to service. Based on the consistency of the readings from earlier tests and the large numler of manual operations, it was decided that the risk of an instrument remaining isolated after the check was greater than an instrument having a zero shift. O l l l l l l i l l l O mvan23n3w.no :it>.too995 3.1-2 REVISION: 1
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. 3.2 Test Matrix i
Before the test' matrix was initiated, a series of pre-operational tests were performed to provide an understanding of facility control and operating characteristics, to confirm design features essential to i scaling, and to check that the instruments and the data acquisition system (DAS) were performing as : expected. Tests were conducted while the facility was in cold conditions to measure system pressure . ! drops and volume of the components. Pressure drops in the test facility were adjusted to the desired scaled AP600 values by using orifice plates. Pre-operational testing was also performed in the hot condition to characterize system heat losses. Results of the pre-operational tests are presented in i Sectio,n 4 of the OSU Final Data Report.m
. System volume determination tests were performed for the accumulators, CMTs, pressurizer, IRWST, e
sumps, SG secondary sides, and reactor vessel to compare the actual volumes with the calculated
. volumes for the facility used in the safety analysis computer codes. i Line resistance determination tests were performed to measure line resistance for the accumulator lines,
. IRWST lines, and sump injection lines for a given flow rate. De pressure drop in CMT injection and ADS 1-3 lines was measured over a range of flows. Resistance of the RNS injection lines was measured to demonstrate that the pressure drops were within 10 percent of each other. De RCP head was measured for full flow and pump coastdown conditions. De line resistances also provided data h G for the computer models of the test facility. Three separate hot functional tests were performed. The objectives of the tests were to show proper operation of the equipment prior to the formal matrix test program and to provide data necessary to ; document temperature characteristics of the system. 'He first hot functional test measured the steady. I state heat loss, natural-circulation flow, and forced-flow characteristics. These tests verified the calculations used to help size the test facility. OSU-HS01 was performed to determine surface heat losses from the system at 100',200',300*, and 400'F; characterize passive residual heat removal (PRHR) under natural circulation and forced cooling; characterize the primary cooling system at 100, l 300,500, and 600 kW: and characterize the CMT natural-convection characteristics. De objective of l the second test (OSU HS02) was to verify the measuring capability of the break and ADS measurement system (BAMS) and the control of the ADS. The third hot functional test was an l ) inadvertent ADS-1 actuation to assess overall facility performance prior to executing the formal matrix ! test program. l 4 The formal test matrix summary given in Table 3.2-1 was designed to provide a wide range of test data on the performance of the passive emergency core cooling systems used in the AP600 The tests were intended to provide overlap with the full height, full pressure SPES-2 tests
- as well as the AP600 plant SSAR calculations. De test parameters investigated were:
= Cold-leg break sizes of 1/2,1,2,4 (bottom of cold leg), and 4 in. (top and bottom of cold leg) f -~
i mAnp600Q344-3w.nomlb-100995 3.2-1 REVISION: 1
. 13reak locations of cold leg with CMTs, cold leg without CMTs, top of pipe, bottom of pipe, balance line breaks, DVI line breaks, and no breaks with inadvertent ADS actuation
. Different single failures, one of four ADS-4 valves, ADS-1 valves, and ADS-3 valves
. Beyond design-basis tests with multiple failures l l
l
- Spurious S signal (no break) !
The break-size range and location, as well as the different assumed single-failure disruptions, provide a thorough and comprehensive set of integral systems data for code validation. o e O m:\ar6000344-3w. con:Ib-100995 3.2-2 REVISION: 1
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B TABLE 3.2-1 w OSU MATRIX TEST
SUMMARY
7 Break -
'l Test No.
Sire and PRHR CVS RNS ADS 4-1 - ADS 4-2 _ p. Location HX Pump Pump (HI 1) (HI 2) Comanents
?
.~2 SB01 2-in. Cl-3 On Off Off 50-percent flow area ' 100-percent flow
$ Failure of one of two lines in . [
. bottom of coM leg in AP600 area in AP600 ADS 4-1; reference cold-leg break (CMT side) case -
.l SB06 4-in. Cl-3 On Off Off 50-percent flow 100-percent flow Failme of one of two lines in bottom of coM leg area in AP600 area in AP600 ADS 4-1 R 3
(CMT sMe)
.i SB09 2-in. Cl-3 On Off Off 50-percent flow 100-percent flow Same as SB01 except different break i to CMT-1 areain AP600 Areain AP600 location; asymmetric behavior of I balance line CMTs SBIO ' DEG Cl 3 to CMT-1 On Off Off 50-percent flow 100-percent flow Limiting break on balance line; M balance line area in AP600 area in AP600 y asymmetric behavior of CMTs; failure of one of two lines in ADS 4-1 SB12 DEO DVI-I On Off Off 100-percent flow 100-percent flow Limiting break on DVI line; failure line break area in area in AP600 of one of two lines of ADS-1 and -
AP600 ADS-3 SB13 2-in DVI-I On Off Off 50-percent flow 100-percent flow Same as SB01 except different break line break areain AP600 areain AP600 location k SB14 Inadvertent ADS On Off Off 50-percent flow 100-percent flow No-break case with one failure of (no break) area in AP600 area in AP600 two lines in ADS 4-1 SBIS 2-in. HI 2 On Off Off 50-percent flow area 100-percent flow Same as SB01 except break location bottom of phe in AP600 area in AP600 ' W m .
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B E TABLE 3.2-1 (Continued) m OSU MATRIX TEST
SUMMARY
y Break g Test Size and PRHR CVS RNS ADS 4-1 ADS 4-2 E. No. Iecation IIX Pump Pump (Hiel) (IIL-2) Comments T s SBIS 2-in. Cl 3 On Off Off 50-percent flow 100-percent flow Repeat test of SB01; confirm behavior S bottom of cold leg area in AP600 area in AP600 of system and instrumentation (CMT side) SB19 2-in. Cl-3 On Off Off 50-percent flow 100-percent flow Same as SB01 except containment bottom of cold leg area in AP600 area in AP600 backpressure simulated (CMT side) SB21 4-in. top of and 4-in. On Off Off 50-percent flow "10-percent flow Same as SB01 except larger break bottom of CL-3 area in AP600 areain AP600 size; largest break size simulated in (CMT side) matrix tests SB23 1/2-in. Cl 3 On Off Off 50-percent flow 100-percent flow Same as SB01 except smaller break bottom of cold leg area in AP600 area in AP600 size [ l (CMT side) i 5 l Note: DEG - double-ended guillotine PRA - probabilistic risk assessment i i 1 l N E cl
.3 1
O O 9
- 4.0 - DATA REDUCTION METHODOLOGY V
The OSU test facility is a scaled model of the AP600 and its passive emergency core cooling systems that were used to perform selected tests to study the thermal-hydraulic behavior of those systems under postulated accident conditions. Data collected from the facility included fluid, metal, and ambient e temperatures, pressures, liquid levels, differential pressures, volumetric flow rates and power supplied
- to the core simulation. Rese data were used as inputs to calculations of instantaneous and integrated
- masses, mass flow rates, energy, and energy transfer into and out of facility components.
The AP600 incorporates and takes advantage of passive safety systems features, that is, flow is driven by differences in temperature (natural convection) and elevation heads (gravity). He Oregon State f j University (OSU) test facility is a scaled model of the systems and components important to the { accident mitigation capability of the AP600 design. Data collected from the OSU test program are j used to validate the safety analysis codes used to predict the performance of the passive safety system ^ for design-basis accidents. i' Computer software has been developed to reduce the experimental data collected from the OSU test { facility into engineering parameters against which code predictions may be directly compared. The l data analysis software performs calculations and generates plot files that contain these calculated q engineering parameters. De calculations performed by this software include mass and energy-balances for the system. i Sections 4.1 through 4.22 describe the underlying assumptions, conversions, corrections and compensations made to the data; the formulation of the mass and energy equations for each component included in the OSU test facility; and the formulation of the overall system mass and energy balance l equations. Also, the nomenclature used in the development of the mass and energy equations is } defined in Subsection 4.0.1. The equations described in this section were coded in FORTRAN to run 4 on workstations. Results obtained from this code were used as the basis of the analyses presented in ! Sections 5,6, and 7, 1 4.0.1 Nomenclature 1 Ris section contains the nomenclature utilized throughout Section 4. In certain subsections, specific
- conversion constants and subscripts are defined, where applicable, within the subsection.
- j. Variable Definitions
, p = Density, Ibift.' '
A = Area, in.2 l l cp = Specific heat, at constant pressure, Blu/(1bm *F) , c, = Specific heat, at constant volume, Btu /(lbm 'F) Az = Height of water, ft.
- Q D = Diameter, in.
m:\ap6004sec-4\2344w-40.aon:Ib 100395 41 REVISION: 1
e = Emissivity E = Energy, Btu g = Gravitational acceleration,32.2 ft/sec.2 g, = Conversion constant,32.2 (1bm ft/lbf-sec.2) h = Enthalpy, Btu /lbm 11 = IIcat transfer coefficient, Btu /hr - ft.2. p k = Mean thermal conductivity, Btu-infar - ft.2, .p L = Elevation, in. LDP = Level transducer reading, in. M = Mass, Ibm M = Mass flow rate, Ibm /sec. P = Pressure, psia Q = Rate of energy transfer, Btu /sec. R = Heat transfer resistance, (hr. 'F)/Bru S = Entropy, Bru/lbm 'F span = Water level span within a tank, in. t = Time, sec. T = Temperature, 'F U = Stored energy, Btu v = Specific volume, ft.'/lbm V = Volume, in.' vf = Void fraction W = Volumetric flow rate, (ft.8/ min. for steam, gpm for liquid) X = Flow quality x = Insulation thickness, in. Z = Compressibility factor Subscririt Definitions ACC = Accumulator ADS = Automatic depressurization system air = Air AMB = Ambient AVG = Average i BRK = Break separator tank CL = Cold leg CLBL = Cold-leg balance line CMT/DVIL = Interface of CMT and DVI-line CMT = Core makeup tank COMP = Compensated corr = Corrected value CVS = Simulated nonsafety Chemical and Volume Control System m :\ap600%s ec-4\1344 w-40. noo : I b- 100395 42 REVISION: 1
f~ DC = Downcomer Q) DEF DVIL
=
=
Deficit DVI line f- = Liquid phase of water fg = Vaporization FL = Fluid in the control volume flux = Flux flux _ cony = Convective flux flux _ rad = Radiative flux g = Steam HO 2
= Water (liquid and steam)
HL = Hot leg .. . initial = Initial inner _ metal = Inner surface of metal l IRWST INJ = IRWST injection LOCAL = Local LOW = Low threshold -
- mean_ metal = Average metal surface metal = Metal i mix = Mixture of liquid and steam i out = Out 1O outer _conv_ metal = Convection at outer metal surface outer _ metal = Outt.r surface of metal outer _ rad _ metal = Rac'.iation at outer metal surface OVERFLW = Overflow 4'
READING = Recorded by the data acquisition system ! i ref = Reference ! RNS = Simulated nonsafety Normal Residual Heat Removal System l RODS = Heated rods j t SAT = Saturation SG = Steam generator ) STM VElfl' = Steam vent I STM XHST = Steam exhaust transient = Transient zone = Zone i-4.0.2 Energy Equation Approximation Several assumptions were made in calculating the total energy associated with the working fluid in the test, these are:
- 1. The total energy of the fluid is approximated by the enthalpy of the fluid. Kinetic and ,
j potential energy are neglected. en:Wec-44344w-40,non:Ib-100395 43 REVISION: 1 i
- 2. He enthalpy of the fluid is approximated by the specific heat of the fluid.
- 3. The change in physical properties between two consecutive time-steps is negligible.
He equation for calculating the change in fluid energy may be represented as: dE dU d(c, M T) A(c, M T) _= = = 4.0.2-1 dt dt dt At where all terms are for fluid at local conditions. Assumption 3 provides for the specific heat at constant volume to be taken as constant over a time interval: dE A(M T) _dt = c' At 4.0.2-2 Assumptions 1 and 2 provide for the specific heat at constant volume to be replaced by the specific heat at constant pressure: I dE A(M T) _dt = c" At 4.0.2-3 For the liquid phase of water, the specific heat at constant volume is approximately equal to the O specific heat at constant pressure: c, = c, 4.0.2-4 Dus, equation 4.0.2-3 holds for the liquid phase of water. For vapor, the definition of enthalpy is noted to be: h=U+Py 4.0.2-5 where the product of the pressure and specific volume may be significant. Ilowever, both the mass of vapor (steam) and the change in vapor mass is small for the OSU tests. Rus, the contribution of the pressure / specific volume term to the total energy associated with a given component, including the reactor vessel simulation,is very small. Rus, the specific heat at constant pressure provides a reasonable approximation of the total energy associated with the steam in the OSU tests. 4.0.3 Ambient Conditions Ambient pressure and temperature for the test facility environment were monitored during each test. He ambient pressure data channel was used to convert data from units of psig to units of psia for m:\a:WXAnc-4\2344w-40.non:1b.100395 4-4 REVISION: 1 l
i
/~\ subsequent calculations, as opposed to using the constant 14.7. ' Ambient temperatures were used to D : evaluate heat loss from the test facility to the ambient. The ambient test facility environment was monitored by the following data channels:
- Channel ID Function
- PT-003 . ambient pressure TF-005 . ambient temperature, lower elevation of test facility; [ ]'6" elevation 7F-006 ambient temperature, middle elevation of test facility; [ ]'6# elevation TF-007 ambient temperature, top elesation of test facility; [ ]'6d elevation I
i: As ' . J
- m:\ap60owec-4\2344w-40 Loon:lt>100395 45 REVISION: 1 a
.p 4.1 LDP Compensation Function
\
Level transducers measure the variation in density of a vertical span relative to a reference. When the l reference and the measured span are not at the same temperature, the effect of density differences needs to be accounted for to obtain a "true" level. The method of applying a compensation to level transducer outputs to account for differences in fluid density between the measured region and the reference leg is described below. First, an average temperature for the column of water being monitored was calculated: a i i 4 I'I TAvo" "[ { L, where: T = Temperature, 'F L = Span along the column that a temperature measurement was applied, in. A and the subscripts: i = Index of measurement arrays n = Total number of fluid temperature measurements in a column of water ] AVG = Span-weighted average fluid temperature in the monitored water column ) The fluid density in the monitored water column was then calculated as: l PAvc = p (P. min (T Ava, TsAT)) 4.1-2 l l where: ) Tsar = Saturation temperature at local pressure l l l mnap600'sec-4\2 M4 w.4a non : l b l 00395 4,].1 REVISION: 1
i The corrected f'uld level of the monitored water column (thai is, the compensated LDP reading denoted as LDPeoup) was calculated as:
~
LDPcoup = LDP aranno 4.1-3 P4vo where: 62.303 = Reference density of water, Ibm /ft.8 LDP = Reading from a level transducer, in. and the subscripts: COMP = Density-compensated value READING = Data as recorded by the data acquisition system (DAS) All level transdacer data used in the calculation of mass inventories, energy of fluid volumes, and inferred flows were density-compensated prior to being used in calculations. O O m:\apuhec-4\2344 w-40.noa:1b- 100395 4,}.2 REVISION: I
gg 4.2 Selected Level Compensations I ' A number oflevel transducer (LDP) readings were density-compensated, the majority of which were calculated as part of the normal mass and energy calculations performed to support analysis of the data. However, a number of other compensated LDPs were calculated to support the OSU Final Data Report."' 'Ihe list of these compensated LDPs, together with Qe pressures and temperatures used to . compensate them, are listed in Table 4.21. The compensated levels are provided in the OSU Final Data Report.* Note that, since these LDPs are not directly used in the mass and energy calculations performed as part of the analysis, no minimum or maximum range information is applied to the compensated LDP. Also, where more than one temperature data channel is listed for a level transducer, a straight numeric average on all the data channels is used. This is in contrast to the calculation modules, where, typically, only the subset of submerged channels is used. f% L) l p V !
. m:Wec 4\2344w-40.non:lb-100395 4.2-1 REVISION: 1
TAllLE 4.21 PRESSURES AND TEhfPERATURES FOR COAIPENSATED LDPs Name Pressure Temperature (s) CLDP-104 FT-102 TF-166, TF-lM, TF-165, TF-167. TF-147, TF-148, TF-149. TF-150 CLDP 109 FT-108 TR-001-1. TR-001-2, TR-001-3, TR-303-1, TR-303-2, TR-303-3, l TR-313 1. TR-313-2, TR-313-3, TR-308-1 l CLDP-Il0 FT-108 TR-001-4, TR-001-5, TR-001-6, TR-303-4, TR-303-5, TR-303-6, TR 313-4, TR-313-5, TR 313-6, TR-308-2, TR-308-3, TR 318-1, TR-318-2 CLDP-ll2 I'F-107 TF-169 LDP-ll5") DP-i l4"' LDP-l l3") CLDP-ll3 I'T-107 TF-169 LDP-l l5"' DP-i l4") CLDP-il5 FT-107 TF-171 CLDP-!!6 I'T-lli TF-126, TF-127, TF-162, TF-163, TF-lM, TF-165, TF-155, TF-156, TF-130, TF-131, TF-147, TF-148 CLDP-127 I'T-107 TF-126, TF-127, TF-162, TF-163, TF-164, TF-165, TF-155, TF-156, TF-130, TF-131, TF-147, TF-148 CLDP-138 E'T-108 TR-001-1, TR-001-2, TR-001-3, TR-303-1, TR-303-2, TR-303-3, TR-313-1, TR-313-2, TR 313-3, TR-308-1, TR-001-4, 'IR-001-5, TR-001-6, TR 303-4, TR-303-5, TR-303-6, TR-313-4, TR 313 5, TR 313-6, TR-308-2. TR-308-3, TR-318-1, TR-318-2 CLDP-139 FT-107 TF 169 CLDP-140 l'T-ll! TF-126 TF-127, TF-162, TF-163, TF-lM, TF-165, TF-155. TF-156, TF-130, TF-131, TF-147, TF-148 CLDP-801 i'T-202 TF-803 Note: (1) These level values were used to adjust the measured pressure from f'T-107, O mna;wxNec-4u344w 40.non:tb-loo 395 4.2-2 REVISION: 1
i I p . 4.3 Accumulators. ~ Each of the two accumulators consists of a tank with a fluid flow path at the bottom connected to the direct vessel injection line. De accumulator fluid mass conservation equations are described in Subsection 4.3.1, and the fluid energy conservation equations are described in Subsection 4.3.2. ! l 4.3.1 Fluid Mass Conservation Equations I
- he general fluid (H O) 2 mass conservation equation, which relates the change in stored fluid mass with respect to time (the fluid-mass time derivative) to the mass flow rates in and out, reduces to the l
, following for each of the two accumulators, ACC,, where i = 1,2: I 1
4 dM g,ojce, , 4,3,3
"' "' #ce, 3
dt where the subscript: . 112 0 = water (liquid and steam) Due to the fact that no steam is present in the accumulators, the fluid mass conservation equation further simplifies to: , N 4.3-2 uce, , , g "' " CC-dt I where the subscript-f = liquid phase of water From this point on in the discussions of the mass conservation calculations, the expressions will be in the liquid-only form. ! The left-hand side of the water mass conservation equation is approximated from the value of the ) i water mass at two consecutive time points (denoted by subscripts n-1 and n)- l l l dMuce, , AMuce, _ M uce,,n - M"" "- ' - 4.3-3 dt At t, - t,,, j i l U i m:Weew344w-40..on:itwioo395 4.3 1 REVislON: 1 l i
The water mass calculations are based on measured water level. In general, the water mass may be ' expressed as follows: hi ucc,
- Puce, x Vucc, x C 4.3-4 where:
C = Conversion constant, in.' to ft.' Due to the cylindrical shape of the accumulator tanks, the water volume is calculated as follows: V uce, = L .Acc, t x A,ce, 4.3-5 Thus, the water mass expression becomes: hi ucc,
- P uce x L,,3cc,x A4ce, x C 4.3-6 O
Rather than using the LDP compensation method to calculate the corrected water level from the measured water level as is done for many other OSU test facility components, an equivalent approach was employed for the accumulators. Using the concept of head and recalling that measured water levels (LDPs) are applicable (calibrated) at a reference density, the following is true: AP [ psi) = p, x L, x (.8 ) x C = p,,, x LDP x ( E.) x C 4.3-7 8 E, so that,in the case of the accumulators: pace, x Lucc, = p,,3cc x LDP-XXX Acc, 4.3-8 where: i I LDP-XXX = Level instrument channel measurement, in. O i m:Wec.4u344w-40..on:Is too395 4.3-2 REVISION: 1 l
I I
/~^ Thus, the water mass expression becomes:
M,,3w, = p,,,,3ce, x LDP-XXX,e xAAce xC 4.3-9 In certain OSU tests, the measured water level for the accumulators was adversely impacted by air in
- the sense line. For those tests, the following correction was developed
I LDP ,3z =YAcc, x F,cc, x LDP-XXXAcc, 4.3 10
- Y4cci si the following constant based on initial conditions
~ '
span 3ce, (Picc, x v,,,3cc/Z,,cc,)= 4.3-11
^CC' (LDP-XXX3ce,),, _
(P gy, x y,,,y,/Z,,,,y3),,_ Face is given by the following: 1
- ( . .
+ (P,y,xv-auaIZairAus)=w 4.3 12 p^ C' (P4ce
- Vw.ACC,/ZarACC,)msers 4
j In the previous equation, the specific volume (ft.3/lbm) and compressibility factor of air are given by the following functions of pressure (psia), respectively: v, = 301.39 x P ""88 4.3 13 Z,,, = 1.0x10-8x P 2 + 3.0x10-5 x P + 1.0002 In the tests without air in the sense line, the factors YA cc and F Acc are both set to 1.0 to deactivate the LDP corTection. O f , ' m:\ap600wec-4\2344w-4 anon:Ib-100995 4.3-3 REVISION: 1
Thus, the final expression for the water mass is as follows: Mucc, = preucc, x LDPecc, x AAcc xC 4.3 14 l This completes the discussion of the calculations related to the left-hand side of the water mass conservation equation. On the right-hand side of the water mass conservation equation, the calculation of the outlet water mass flow rate is performed directly from the liquid volumetric flow rate (FMM) measurement as follows: ( uce, = max (0.0, FMM-XXXAcc) x p ucc, x C, 4.3-15 where: FMM-XXX 4cc, = Liquid flow rate measurement from flow meter FMM-XXX, gpm Ci = , Conversion constant, gpm to ft.'/sec. The integrated outlet water mass flow is also computed for use in the overall system mass balance. The water density is simply the reciprocal of the water-specific volume: 1 I p uce = v 4.3-16 , ucc, I The water-specific volume is calculated as a function of pressure and temperature from the ASME steam table function VCL: l 1 l v acc, = VCL(PAcc, Tucc) 4.3-17 l i l The water temperature is given by the fluid thermocouple measurement. Care is taken to insure that l the water temperature value employed is at or below the saturation temperature. Thus, for i = 1,2: l l Tucc, = min (TF-XXXAcc, T,,,3cc,) 4.3-18 I i I 9l l mhp60lAsec-4W44w-40 non:lb- 100395 4.3-4 REVISION: 1 I l
l (N where: TF-XXX 4ce, l
= Temperature reading from fluid thermocouple TF-XXX, 'F l
1he saturation temperature is calculated as a function of pressure from the ASME steam table ftmetion TSL as follows: l T,3ce, = TSL(Pice) 4.3-19 1 Accumulator pressure, PA cc o is set equal to the value from the corresponding PT-XXXACCi measurement after conversion from gauge (psig) to absolute (psia) pressure. This completes the discussion of the calculations related to the right-hand side of the water mass l conservation equation. Note that in the case of the accumulators,it is possible to calculate all quantities on both the left-hand side and right-hand side of the fluid mass conservation equation. This information can then be used to l check the fluid mass conservation equation. l Finally, a complete list of the applicable data for both accumulators appears in Table 4.3-1, including the level instrument channel ID (LDP-XXXACCi), the pressure measurement D (PT-XXX ACC IhO 8 fluid thermocouple measurement ID (TF-XXXA cc), the outlet liquid volumetric flow rate measurement ID (FMM-XXX ce), 4 the water reference density, the water level span, and the tank cylindrical area. ] 4.3.2 Fluid Energy Conservation Equations The general fluid (H 2O) energy conservation equation, which relates the change in stored fluid energy with respect to time (the fluid-energy time derivative) to the energy rates in and out (due to the connected flow paths) and the energy addition rate due to other external devices, reduces to the following for each of the two accumulator, ACC,, where i = 1,2:
- d[M x c, x (T-T,)]n 03cc dt ' "P^CC. (**'* "Pl^cc,
- O m:pec-4u344 40..on:ib ioo395 4.3-5 REVISION: 1 l
Due to the fact that no steam is present in the accumulators, the fluid energy conservation equation further simplifies to: d[M x c, x (T-T,,,)]racc. " ~ 4 3-21 dt '" **'*** From this point on in the discussions of the energy conservation calculations, the expressions will be in their reduced, water +nly form. The left-hand side of the water energy conservation equation is approximated as follows: d[M x c, x (T-T,,,)]ucc, _ A[M x c, x (T-T,,,)]ucc, 4.3-22 dt At and: A[M x c, x (T-T,,)]ucc, A[M x (T-T,,,)]uce,
-c x At P" At ,
i 4.3-23
= cPuce, x (TfACC, Tre ,) x M " "'" - M " " *'
g ,g
"" """~'
+ c, xMucc, x a n -1 ,
where the subscripts: n-1, n = Two consecutive time points The water-specific heat capacity at constant pressure is calculated as a function of pressure and temperature from the ASME steam table function CPL as follows: c, = CPL (PAcc, Tuce,) 4.3-24 O m:hr'6cossec-4uw4 anon:tb ioo995 4.3-6 REVISION: 1
For use in the overall system energy balance calculations, the water stored energy in each accumulator
. is given by the following:
Uucc, = .M acc, x c p,uce, x (Tucc, - T,,,) 4.3-25 Eis completes the discussion of the calculations related to the left-hand side of the water energy conservation equation. On the right-hand side of the water energy conservation equation, the calculation of the outlet water energy transport rate is given by the following, where i = 1,2: Q . u cc, " M . u ce,x hucc, 4.3-26 The water-specific enthalpy is calculated as a function of pressure and temperature from the ASME steam table function HCL as follows (note that the water-specific entropy is also an output of HCL): huce, = HCL (PAcc., Tucc,, Suce,) 4.3-27 A Q The heat transfer between the accumulator tank metal and the accumulator fluid inventory during accumulator injection is negligible, for the following reasons:
- The accumulators are pre-filled before initiation of the tests, so that the fluid and tank metal walls are in thermal equilibrium at the beginning of the transients.
- Although the overpressure gas decreases in temperature as it expands to fill the volume left by the injected water, the water temperature remains approximately uniform throughout the
- injection phase.
[ . He mass and specific heat of the overpressure gas is small compared to the mass and specific heat of the tank metal, and therefore draws an insignificant amount of energy from the tank metal walls. Rus, the accumulator metal-to-fluid energy addition rate is neglected, and: Q n,3cc = 0.0 4.3-28 $ / His completes the discussion of the calculations related to the right-hand side of the water energy conservation equation. j. , mAap600sec-4\2344w-4Claos:ltv100395 4.3-7 REVISION: 1
Note that, in the case of the accumulators, it is possible to calculate all quantities on both the left-hand side and right-hand side of the fluid energy conservation equation. This information can then be used to check the fluid energy conservation equation. O O. l mAap60thec 4\2344w-40.noo:lb-100395 4.3-8 REVISION: 1 i
I I (3 i TABLE 4.3-1 , INSTRUMENTATION EMPLOYED FOR ACCUMULATOR FLUID CALCULATIONS l Description ACC1 ACC-2 Level ID (in.) LDP40! LDP-402 Pressure ID (psig) Irr-401 l'T-402 l l Water temp ID ('F) TF401 TF-402 ; Outlet flow ID (gpm) FMM-401 FMM-402 Ref water p (lbm/ft.') 62.40 62.40 Water level span (in.) 36.75 37.00 Tank area (cylin) (in.2) 416.6208 417.1392 1 l rr l l 1 l l lG m%*nseo4u)44.-4a non:It> too395 4.3-9 REVISION: 1
I l L/~N- 4.4 Core Makeup Tanks and Cold Leg flalance Lines
)
Each of the two core makeup tanks (CMTs) consists of a tank with a fluid flow path at the top for the cold-leg balance line (CLBL) connection aind a fluid flow path at the bottom for the direct vessel injection line (DVI) connection. Each of the two cold-leg balance lines consists of a series of connected pipes with a fluid flow path at the bottom for the inlet connection from the cold leg and a fluid flow path at the top for the outlet connection to the CMT. He CMT fluid mass, fluid energy, and metal energy conservation equations are described in Subsections 4.4.1,4.4.2, and 4.4.3, respectively. The cold-leg balance line fluid mass, fluid energy, and metal energy conservation equations are described in Subsections 4.4.4,4.4.5, and 4.4.6, respectively. 4.4.1 Core Makeup Tank Fluid Mass Conservation Equations The general fluid (H O)2 mass conservation equation, which relates the change in stored fluid mass with respect to time (the mass time derivative) to the mass flow rates in and out, reduces to the following for each of the two CMTs, CMT,i where I = 1,2: l dM 4,4.j ap. cur, , g,"
' *"" N "#M _ g4 "' " N "P"'
dt O
\.) De left-hand side of the fluid mass conservation equation is approximated from the value of the fluid mass at two consecutive time points (denoted by subscripts n-1 and n) as follows: l l
i dMap. cur, , AMap. cur, , Map. cur,.a -M ap. cur,."- 4.4-2 dt At t, -t, ,, i I The H 2O fluid mass is the sum of the liquid and steam masses: l l l i M ap. cur, = M .r cur, + M .s cur, 4.4-3 l 1 where the subscripts: HO2
= Water (liquid and steam) f = Liquid phase of water g = Steam m
m:\np600wec-4V344w-40.non:lb.100395 4,4. ] REVISION: 1
The water and steam mass calculations are based on the measured tank water level. A list of the level (LDP-XXX) instrument channel ids for both CMTs appears in Table 4.4-1. To span the entire tank, either the main LDP, or a sum of the three alternate LDPs, is employed for the indicated water level, which is then corrected for temperature effects using the LDP compensation method. As is generally assumed for each OSU test facility component, zero quality water is modeled below the compensated water level, and 100 percent quality steam is modeled above the compensated water level. This assumption is appropriate because the fluid below the water level is predominantly water, the fluid above the water level is predominantly steam, and any amount of frothing is indeterminate with the available instrumentation. To utilize the measurements from the numerous tank fluid thermocouples available, each Ch1T is divided into a number of axial fluid property zones for the calculation of various fluid conditions. Table 4.4-1 contains a list of the fluid thermocouples employed (ten per tank), along with their elevations. Axial fluid property zone boundaries are taken at the vertical midpoint between consecutive fluid thermocouples, which yields ten zones per tank. In general, the tank liquid and steam masses are given by: N cm. M.r cur, = { M . rcur, 4.4-4 N, i Ms . cur, = { M . scur, Jet where: N aon.. cur, = Number of zones in the CMT, (10) j = Specific zone within the CMT For zones j below those containing the compensated water level (which contain all water), the zone j I water and steam masses are given by, where i = 1,2 and j = 1,...,levzone-1: l l l M r. cur,
- Pr. cur" x Vcwr" x C 4.4-5 ;
M s. cur, = 0 1 l el l m4whe-4u344w-40.noa:15too395 4,4 2 REVIslON: 1 l l
where: C = Conversion constant, in.8 to ft.8 levrone = Zone containing the compensated water level For the zone j containing the compensated water level, the zone j water and steam masses are given by, where i = 1,2 and j = levzone: hi r.cstro " P r.curr, x V .cstr, r xC hi s.cstr,
- Ps.cstr,
- Ys .cstry xC For zones j above those containing the compensated water level, which contain all steam, the zone j water and steam masses are given by, where i = 1,2 and j = levzone+1,...,N,, .cstri:
hi r.cstr" =0 4.4-7 hi s. cur,
- Ps.cstr,
- Veur, x C The tank fluid volumes are calculated as a function oflevel from the volume versus height tabular data listed in Table 4.4-2, via linear interpolation within the table. No extrapolation at either end is performed; the first and last table points define the applicable range (the first point is [h ,(= 0.),
V , (= 0.)], and the last point is [h,,, V,,) ). During initialization, the volume versus height data / function is employed to calculate the total volume of each zone j as follows: forj=l: 4.4-8 Vcstr = VO)cstr. . IJ w for J-2,...,N a .cstr: 1-1 V ewr. = VO) cur. . - E Vcur. 4 y n.i a m:\agWXAsec-4\2344w-4 anon:Ib 100395 4,4 3 REVISION: 1
where: V(1) cur = Volume as a function of elevation, in.3/in., and 1 is the elevation of the top of zonej The zone j top elevations are determined from the elevations of the fluid thermocouples (or the elevation of the top of the tank in the case of the top zone). Du*ing the transient calculations, the volume versus height data / function is employed to calculate the watar volume of the zone containing the compensated water level,J = levzone, as follows: if j=levzone=l: 4.4-9 V r. cur. . = V(1,) cur' u if j=levzone>1:
;-n Vr . cur.
- V0r) cur, - { Vcur 8J t -1 LA where:
VO,) cur, = Volume as a function of elevation,in.8/in., and I,is the elevation of the O compensated water level. The steam volume of the zone containing the compensated water level, j = levzone, is then the following: V
- s. curg
=V cur - V .t cur 4,4-10 g
The zone j water and steam densities are the reciprocal of the water- and steam-specific volumes, respectively: 1 P r. IJcur. .r " v . cwt
; 4.4-11 O g. Cur y
vs. curg O maap60wec-4u344w-4 anon:Ib-too395 4.4 4 REVISION: 1
l i i The zone j water specific volume is calculated as a function of pressure and temperature from the 1 ASME steam table function VCL as follows, where 1 = 1,2: vr . cur = VCL(PcMr,j 4.4 12 T.r cur) - The zone j steam specific volume is calculated as a function of pressure and temperature from the ASME steam table function HSS as follows, where i = 1,2 (note that the steam-specific enthalpy is the main output of HSS, and the steam-specific entropy is also an output of HSS): hs. cur, = HSS(Pcury T.s cur,+'S.curj V s. cur,) 4.4-13 s he zone j water and steam temperatures are given by the zone temperatures, which are the fluid thermccouple measurements TF-XXX (see Table 4.4-1). Extra precautions ensure that the water temperature value employLd is at or below the saturation temperature, and that the steam te.mperature value employed is at or above the saturation temperature. Thus: n I. rcur" = ndn(Tcur", T i. cur") 4.4-14 a.CMT, CMr' y sat. Cur y ) where: Tcur, = TF-XXX cur, 4.4-15 The zonc j saturation temperature is calculated as a function of pressure from the ASME steam table function TSL as follows: T . cur, = TSL(Pcur,) 4.4 16 The pressure of the top zone is set equal to the value from the FT-XXXcuri measurement (see Table 4.41) (after conversion from gauge [psig] to absolute [ psia] pressure). For the zones below, the pressure is adjusted for hydrostatic effects to account for the density differences that occur from zone to zone throughout the tank. O ' l m%sxwc-4us44.-40..on:ib ioo395 4,4 5 REVISION: 1 m
I l This completes the discussion of the calculations related to the left-hand side of the fluid mass conservation equation. l l On the right-hand side of the fluid mass conservation equation, the CMT inlet fluid mass flow rate is ; equal to the cold-leg balance line outlet fluid mass flow rate (which appears on the right-hand side of l the cold-leg balance line mass conservation equation): l m (from CLBL) H,0,Chfr, " out (to ChfT) H,0,CLBL, Although a liquid volumetric flow rate measurement (FMM) is available for this line, it cannot be relied on for accurate mass flow rate calculations because two-phase fluid flow can occur in the line. Thus, the total mass flow rate must be inferred from a combination of the CMT and cold-leg balance line mass conservation equations, once all other terms have been calculated in these equations. The integrated inlet mass flow is also computed for use in the overall system mass balance. , Regarding the CMT outlet fluid mass flow rate to the DVI line, the calculation may be performed directly from the liquid volumetric flow rate measurement (FMM-XXX), since when there is flow, the water flowing in the line is at either subcooled or saturated conditions (see Table 4.4-1 for the instrument channel ids). 'Ihus: out (to Dv1) H,0.Chft, " out (to Dv!L) f.ChfT, 4.4-18
= max (0.,FMM-XXXChrriova)
- P r.Chrrinva, x C, i
where: FMM-XXX = Liquid volumetric flow rate measurement from flow meter FMMM-XXX, gpm Ci = Conversion constant, gpm to ft 8/sec. The integrated outlet water mass flow is also computed for use in the overall system mass balance. The outiet line water density is the reciprocal of the outlet line water-specific volume: 1 P t.Chrriovt' 4.4-19 v t.Chrr/Dyn., 1 O ms e60w4u344. 40 on:nwoo395 4,4 6 REVISION: 1
.. . - _ . . . - - ~ - _ . . _. . - . .
g The outlet line water-specific volume is calculated as a function of pressure and temperature from the I ASME steam table function VCL as follows: V.cunovt, t = VCL(Pcunova./ T.cunova,) r 4.4-20 The outlet line water temperature is simply given by the outlet line fluid thermocouple measurement TF XXX (see Table 4.4-1). Care is taken to ensure that the water temperature value employed is at or
- below the saturation temperature. Thus:
T.curova., r = ndn(TF-XXXcunovA, T,,,cuyova) 4.4-21 The outlet line saturation temperature is calculated as a function of pressure from the ASME steam table function TSL as follows: T.n.curova, = TSL(Pcunova.,) 4.4-22 The pressure in the outlet line is equal to the pressure at the bottom of the tank. Q This completes the discussion of the calculations related to the right-hand side of the fluid mass l conservation equation. 1 l 4.4.2 Core Makeup Tank Fluid Energy Conservation Equations i 1 The general fluid (H O) 2 energy conservation equation, which relates the change in stored fluid energy l with respect to time (the energy time derivative) to the energy rates in and out (due to the connected flow paths) and the energy addition rate due to other external devices, reduces to the following for j each of the two CMTs, CMT,, where i = 1,2:
)
i d[M x c, x (T-T,,,)]n,o.chrr I dt
" e a cw up. cur, - % = ovy up. cur, !
4.4-23
+ O(membe n,0), Cur, I
l
.v t O) .
m:Wec 4\2344w-40.non.it>100395 4,4 7 REVISION: 1
I l l l l The left hand side of the energy conservation equation is approximated as follows: d[M x c, x (T-T,,,)]u,0.c6rr, - " ~. A[M x c, x (T-T,,,)]r. cur, dl E,.i At 4.4-24
+
A[M x c, x (T-T,,,)]Scur, b.i 3 at where: A[M x (T-T,,,)]r. cur, A[M x c, x (T-T,,,)]r. cur a". x = c^ At , At ,
' ~
4.4-25 M"~" - MW"
- c,'""wx(Tr* cur9 -T,,,) x t,-1,,g
+
Pian, xM twt x T"I# T#"I~" w t , -t, .i and: A[M x c, x (T-T,,,)],. cur, - c A[M x (T-T,,,))s. cur-x At P.n,ry gg 4.4-26 M .chrr,n - M .chrr,n-i s
=c P.are, xU8 Chff,- ret T)x g ,g n n-1 ,
g.Chrry ~ 3,Chrr yn' l P,orr, g,Chrr, , where the subscripts: n-1, n = Two consecutive time points O mfugiaAsec-4\2344 w-40. non : l ts l 00395 4,4 8 REVISION: 1
i l The zone j water-specific heat capacity at constant pressure is calculated as a function of pressure and
'v - temperature from the ASME steam table function CPL as follows:
c,,mr, = CPL (Pcuro , T.c6rr.) r 4.4-27 The zone j steam-specific heat capacity at constant pressure is calculated as a function of pressure and temperature from the ASME steam table function CPV as follows (note that the steam-specific volume is also an output of CPV): I c, = CPV(Pckrrj T s.curj V .s cur,) 4.4-28 For use in the overall system energy balance calculations, the fluid stored energy in each of the two l CMTs is given by the following: a U np. cur, = U r. cur, + U .s cur, 4.4-29 where: V("N - ~__. U r. cur, = { M .cstr, r
- C .r.r cur,
- Ur.cstr, - T,,,)
Jet and: s_ _. AM U.s cur. - { M . cur,s
- C rcur,
- 6. cur, 5 - T,,,)
!=1 This completes the discussion of the calculations related to the left-hand side of the fluid energy conservation equation.
On the right-hand side of the fluid energy conservation equation, the CMT inlet fluid energy transport rate is equal to the CLBL outlet fluid energy transport rate (which appears in the CLBL energy conservation equation): Om (erom ctsLa n,o.c6rr, " Si <i. cstra a,o.ctsk 4M O' v m:sapsxwc-4u344w-40.no :tb ioo395 4,4-9 REVIs!ON: 1 I 1
In a manner analogous to that performed for the corresponding term on the right-hand side of the fluid mass conservation equation, the above CMT inlet fluid energy transport rate term must be inferred i from a combination of the CMT and CLBL energy conservation equations, due to the two-phase flow which exists, once all other terms have been calculated in these equations. l l The CMT outlet fluid energy transport rate to the DVI line is given by the following: out Do DVL) H,0.Chrr, out Go DVL) f.ChfT, 4,4-33 out go DYL) f. Cart, f.Cbfr/DV1, ne outlet line water-specific enthalpy is calculated as a function of pressure and temperature from the ASME steam table function HCL as follows (note that the water-specific entropy is also an output of IICL): h.chrrevn., t = IICL(PchmDyn.,, T.currevn.,' f S .chmDvn.,) t 4.4-34 Re energy addition rate to the fluid due to heat transfer from the metal walls is calculated from the O CONTRA inverse heat conduction" solution, and is discussed in Subsection 4.4.3. His completes the discussion of the calcuhtlons related to the right-hand side of the fluid energy conservation equation. 4.4.3 Core Makeup Tank Metal Energy Conservation Equations he general metal energy conservation equation, which relates the change in stored metal energy with respect to time (the metal energy time derivative) to the energy rates in and out (due to heat transfer), reduces to the following for each of the two CMTs, CMT,, where i = 1,2: d[M x c, x (T-T,,)..i.chrr, , _ Q(==i. n m , qimmi..w.chrr,4,4-35 dt , , To utilize the measurements from the numerous wall thermocouples available, the metal of each tank is divided into 13 metal segments which correspond to the axial positions where wall thermocouples are located in the CMTs. Table 4.4-3 lists the wall thermocouple channel ids used for each metal segment of both CMTs. In addition, Table 4.4-4 lists the mass, inner surface area, and outer surface mw6oossec.4us44. 40 non: b-100395 4,4-10 REVislON: 1
. - ~ . .- .. . . . .. . . . _ - - ..
I f^ area for each metal segment, as well as the ambient temperature channel ID, all of which are required in the metal energy calculations. The left-hand side of the fluid mass conservation equation is approximated from the value of the fluid - mass at two consecutive time points as follows: d[M x c, x (T-T,,,)]n tcur, , A[M x c, x U-T,.tH..i. cur. 4.4-36 dt At where: A[M x c, x (T-T,,,)]..i. cur, " T..tcur . - T..tcur n-i -4,4-37 At .i
**'"'! '-'"* ? t -I a n where: .
?
i N..t cur, = Total number of metal segments in each CMT, (13) and the subscripts; m = Specific metal segment of the CMT n-1, n = Two consecutive time points In the equation above, the temperature used for each metal segment is an average of the temperatures obtained from the thermocouples listed in Table 4.4-3 for that metal segment. l Die metal segment specific heat capacity r,t constant pressure is calculated as a function of the metal j segment average temperature from the metal c, versus temperature tabular data listed in Table 4.4-5, ,
- via linear interpolation within the table. No extrapolation at either end is performed; the first and last )
f table points define the applicable range (the first point is [T,,, c,,,,], and the last point is [T,,, c,,,,]). Riis completes the discussion of the calculations related to the left-hand-side of the metal energy conservation equation. 4 ? 4 1 i 1 , m%ecouc.4u344.-4ano.:ts too395 4,4-11 REVISION: 1
On the right hand-side of the metal energy conservation equation, the total metal-to fluid heat transfer rate is equal to the sum of the individual metal segment metal-to-fluid heat transfer rates: N .., (metal = H,0).Chfr, metal H,OI.Chfr.
..i The integrated metal-to-fluid heat transfer is also computed for use in the overall system energy balance.
For the detailed analyses of the heat transfer between the fluid and the metal wall in each of the two CMTs, the CONTRA inverse heat conduction"" solution is employed to calculate the transient inner surface heat fluxes and temperatures, given the transient temperature data collected from thermocouples imbedded in the walls of the CMTs. The CONTRA inverse heat conduction calculational method models a wall as a one-dimensional cylinder in radial coordinates, and applies to each axial position where wall thermocouples are located in the CMTs (refer to Table 4.4-3). A stand-alone pre-processor code ("osucontra", the OSU CONTRA driver software) employs the CONTRA inverse heat conduction solution on the transient wall thermocouple temperature data for the modeled metal segments of each of the two CMTs and calculates the transient inner surface heat flux and temperature for each of these metal segments. Therefore, the energy addition rate to the fluid due to heat transfer from the CMT metal is given by the following for e'wh metal segment m: l l Oi .i Hpi. Cur,, " Onnim.w. Hpi.Chfr,, * ^=,r_mmi. Cur x C2 4.4-39 1 where: Qn, = Ileat flux, Btu /hr.-ft: C2 = Conversion constant, hours to seconds Also on the right hand-side of the inetal energy conservation equation, the total metal-to-ambient heat transfer rate is equal to the sum of the individual metal segment metal-to-ambient heat transfer rates: N,,,, lmetale ambiem],Chft, (metale ambent},ChfT,, m.i
'Ihe integrated metal-to-ambient heat transfer is also computed for use in the overall system energy balance. ;
O l 1 m:\ap600\sec-4\2344 w-40. non : l b- 100395 4,4 12 REVISION: 1
q ' The metal segment m metal-to-ambient heat transfer rate is given by the following: b Q tomaw amb no. Cur,, " 00 erm.iaw amb=1.CMr,,
- A oui .r...iai. Cur x C2 4.4-41 The metal segment m metal-to-ambient heat flux is expressed as the sum of the convective and radiative heat fluxes as follows:
Du lastaW ambame],Ch.",, Cus.conv (metakambm].CMT. 4.4-42
+
OuJad [metab ambnent),CMT. Note that for ambient heat transfer considerations, each CMT metal segment is treated as a vertical cylinder (CMTs are not insulated). The metal segment m outer surface convective heat flux is given by the following, where i = 1,2: Du_conv lactakambuent],CKr, outer _conv_ metal.CMr,, cuter metal,CMT,, ~ AMB where:
%/
H = Heat transfer coefficient, Btu /hr.-ft.2 op A natural convection heat transfer correlation"* of the following form is employed: Nu = C (Gr Pr)" 4 4-44 where, for vertical cylinders: C = 0.13, m = 1/3 ; for (Gr Pr) > 10' (turbulent range) 4.4 45 C = 0.59, m = 1/4 ; for (Gr Pr) > 10' (laminar range) X-n460msecauw4anoa:n>. loo 395 4,4-13 REVISION: 1 1
For the OSU test facility air at atmospheric pressure and ambient temperature (averaged over the tests), the turbulent range (Gr Pr > 10'; C = 0.13, m = 1/3) applies at the outer surfaces of the CMTs, and the above correlation simplifies to the following expression for the convective heat transfer coefficient for each metal segment m: li,,,,,,,,,,=tcur. = 0.19 x ( ITow.mmtCur,, - Tiu,I )v2 4.4-46 Employing the assumption of radiation heat transfer between infinite parallel planes in the limiting case of a convex object completely enclosed by a very large concave surface (typical assumption for the calculation of radiation energy loss from a hot object in a room), with an assumed emissivity of 0.84 for the uninsulated CMTs, the metal segment m outer surface radiative heat flux reduces to: 4 On.a.ot..a..mb-niiCur. = 1.4389 x 10 x [ (T== .tCur,, + C3 )' 4,4,47
- (T,,, + C )' ] 3 where:
C 3
= Conversion constant,459.6 ('F to 'R)
- Ihe coefficient 1.4389 x 104 is the product of the Stefan Boltzmann constant (1.713 x 10d Btu /(hr.-
ft.2 *R')) and the emissivity (0.84). Uds completes the discussion of the calculations related to the right-hand-side of the metal energy conservation equation. Note that it is possible to calculate all quantities on both the left-hand side and right-hand-side of the metal energy conservation equation, as shown above. Uds information can then be used to check the metal energy conservation equation. 4.4.4 Cold-Leg Balance Line Fluid Mass Conservation Equations The general fluid (110) 2 mass conservation equation, which relates the change in stored fluid mass with respect to time (the fluid inass time derivative) to the mass flow rates in and out, reduces to the following for each of the two cold-leg balance lines, CLBL, where i = 1,2: dM HACLB( 4,4 4g g . g{m (CL) HACLE( , 1mt (CMT) H,0.CLBL, O mangao ec 4u344w 4anoa: b-too395 4,4 14 REVIslON: 1
l The left-hand side of the fluid mass conservation equation is approximated from the value of the fluid
, mass at two consecutive time points (denoted by subscripts n 1 and n) as follows:
dM gp cat, _ AMgp. cat, M ap. cut. -M gp,ctng,,.i 4,4,49 di At t, - t,,, The 110 2 fluid mass is the sum of the water and steam masses: M ap,ct,q 4.4-50
= M r.ctat. + M s.cah The water and steam mass calculatioris are based upon the measured water level. A list of the level (LDP-XXX) instrument channel ids for both cold-leg balance lines appears in Table 4.4-6. These LDPs are corrected for temperature effects using the LDP compensation method. As is generally assumed for each OSU test facility component, zero quality water is modeled below the compensated water level, and 100 percent quality steam is modeled above the compensated water level. This assumption is appropriate because the fluid below the water level is predominantly water, the fluid above the water level is predominantly steam, and any amount of frothing is indeterminate with the 3
available instrumentation. O For the fluid property calculations, the measurements from the two fluid thermocouples available in each CLBL are utilized. Table 4.4-6 contains a list of these fluid thermocouples; one is located at the bottom of the line and one is located at the top of the line. If the line is completely filled with water (water level at top), an average of the two fluid thermocouple temperatures is employed for the water properties. If the line consists of all steam (water level at bottom), an average of the two fluid thermocouple temperatures is employed for the steam properties. Otherwise (water level in between top and bottom), the bottom fluid thermocouple temperature is employed for the water properties, and the top fluid thermocouple temperature is employed for the steam properties. In general, the liquid and steam masses are given by: 1 l M r.ctst,
- Pr.ctat. x Vme, x C 4.4-51 !
M,.ct,t, = p,cgq xV guq xC where: l l 1 C = Conversion constant, in.2 to ft.8
'b U I i
m:\np60wec-4u344 -40.non:1b-t00395 4,4 15 REVis10N: 1
The water volume is calculated as a function of compensated water level from the volume versus height tabular data listed in Table 4.4-7, via linear interpolation within the table. No extrapolation at l either end is performed; the first and last table points define the applicable range (the first point is [h , (= 0.), V , (= 0.)], and the last point is [h V.,1). Thus: Vr .ctst, " V(I)ctat, 4.4-52 where: V(1)caq = Volume as a function of elevation, in.'/in., and I is the compensated water level The steam volume is then given by: Vs .ctat, =V cgg - V .ctat, t 4.4-53 in the above, the cold leg ualance line total volume, Veat, h given by the last point (V .) in the aforementioned volume versus height table. The water and steam densities are the reciprocal of the water- and steam-specific volumes, respectively: 1 O f.CLB(
- y'"
4.4 54 1 Ps.ca4 " y .ctat, The water-specific volume is calculated as a function of pressure and temperature from the ASME steam table function VCL as follows: vr .Ctsl, = VCL(Pcat,, T,,ct,t,) 4.4-55 9 m:Wesu 4u344w-40 noa:tb-ioo395 4,4-16 REVISION: 1
g-
~
The steam-specific volume is calculated 'as a function o'f pressure and temperature from the ASME-1 steam table function HSS as follows (note that the steam-specific enthalpy is the main output of HSS, and the steam-specific entropy is also an output of HSS): h,, cat, = HSS(Pct,q, T,ctat,, S,, cat,. v,,ctat,) 4.4-56 As stated above, the water and steam temperatures are given by the fluid thermocouple measurements , 1F-XXX (see Table 4.4-6). Care is taken to ensure that the water temperature value employed is at or below the saturation temperature, and that the steam temperature value employed is at or above the saturation temperature. Thus: T.caq r = min (T,, cat,, Toi.cLat) . T,,ct,q = max (T s.ctat,, Tut.ctsq) The saturation temperature is calculated as a function of pressure from the ASME steam table function TSL as follows: T,ct,q = TSL(Pcaq) 4.4-58 in all of the above, the cold-leg balance line pressure, Pctat , is set equal to the value from the corresponding PT-XXXcm , measurement (see Table 4.4-6 for list) (after conversion from gauge [psig] to absolute [ psia] pressure). , This completes the discussion of the calculations related to the left-hand side of the fluid mass conservation equation. ; i 4 Although a liquid volumetric flow rate measurement (FMM) is available for each cold-leg balance
- line, the measured values cannot be relied on for accc ate mass flow rate calculations since two-phase l fluid flow can occur in the line. Thus, the total mass flow rates on the right-hand side of the cold-leg I
! balance line fluid mass conservation equation must be inferred from a combination of the CMT and cold-leg balance line fluid mass conservation equations, once all other terms have been calculated in these equations. The integrated cold-leg balance line outlet (to CMT) mass flow is also computed for , use in the overall system mass balance. This completes the discussion of the calculations related to the right-hand side of the fluid mass conservation equation. O 1 l
.. i m:\np60thec-4u344w-40.aos:1b 100395 4,4.]7 REVislON: 1 1 i
I
+ < - 1
I \ l 1 1 4.4.5 Cold-Leg Italance Line Fluid Energy Conservation Equations 1
'Ihe general fluid (110) 2 energy conservation equation, which relates the change in stored fluid energy with respect to time (the fluid-energy time derivative) to the energy rates in and out (due to the connected flow paths) and the energy addition rate due to other external devices. 4e.1uces to the following for each of the two cold-leg balance lines, CLBL,, where i = 1,2:
d{M x c, x (T-T,,,)],,o,ct,t ~
'" "'
- cy H,OM( out Oo C@ H,0.CLB( 4,4,$g dt I
+
( Imetalue H,0).CLBL, The left-hand side of the fluid energy conservation equation is approximated as follows: d[M x c, x (T-T,,,)]w,0.ctat, , A[M x c, x (T-T,,,)]r.ctsq dt At 4.4-60 A[M x c, x (T-T,,,)],ct3t, At where (subscripts n-1 and n denote two consecutive time points): A[M x c, x (T-T,,,)],,ct3e' = A[M x (T-T,,,)]r.ctat-c x At P=. At .
=
c, , x (T,,ct3t, T,,,) x
"* " MM
( T,,ct3t,,, - T .ctsq.n-i r
+ c, x M,,ctat, x t -I n n-t ,
O m:wuec-4um-4a os:1b.ico395 4.4-18 REVISION: 1
6i 1-
- and; 4.G i /-
+
- A[M x c, x (T-T,,)],, cat, A[M x (T-T,,,)],ct,t, At - c' a x . At. ..
M,,ct,q,, - MSctsq.n-1 - 4,4 62
" 0 ~
p% g.Cla( ref g a n-1 1 s.CtB(,n ~ g.CtB(.n-t
+ c, x Ms.caq
- tn -I n-1
.The water-specific heat capacity at constant pressure is calculated as a function of pressure and temperature from the ASME steam table function CPL as follows:
c, = CPL (Pctaq, T,, cat,) 4.4-63 D Q The steam-specific heat capacity at constant pressure is calculated as a function of pressure and ! temperature from the ASME steam table function CPV as follows (note that the steam-specific volume is also an output of CPV): c, = CPV(Pctsq, T,,ctat,, v s. cat,) 4.4-64 For use in the overall system energy balance calculations, the fluid stored energy in each of the two cold leg balance lines is given by the following: Uso,ctag = U,,ct,q + U.cuq s 4.4-65
;;~
n where: Ur .ca( = M,,caq x c,,,,caq x (T,,ct,q - T,,,) 4.4-66 l m:We 4u344w-40.no : b ioo395 - 4.4-19 REVISION: 1
i and: 4 Us .ctst, = Ms .cLat, x pc ,,ct,q x(T,,ctag - T,,) 4.4-67 1his completes the discussion of the calculations related to the left-hand side of the fluid energy conservation equation. On the right hand side of the cold-leg balance line fluid energy conservation equation, the energy addition rate to the fluid due to heat transfer from the cold-leg balance line metal is calculated from the cold-leg balance line metal energy conservation equation, which is discussed in Subsection 4.4.6. Once the metal-to-fluid heat *ransfer rate is known, the fluid energy transport rates on the right-hand side of the cold-leg balance line fluid energy conservation equation are inferred from a combination of the CMT and cold-leg balance line fluid energy conservation equations, in a manner analogous to that performed for the mass flow rates on the right-hand side of the fluid mass conservation equation, once all other terms have been calculated in these equations. This completes the discussion of the calculations related to the right hand side of the fluid energy conservation equation. 4.4.6 Cold Leg Balance Line Metal Energy Conservation Equations
'Ihe general metal energy conservation equation, which relates the change in stored metal energy with respect to time (the metal-energy time derivative) to the energy rates in and out (due to heat transfer),
reduces to the following for each of the two cold-leg balance lines, CLBL,i where i = 1,2: d[M x c, x (T-T,,)]=='CL84 , 4.4-68 dt gI"*"" '*b _ g'"*" "*""* b l Due to the fact that no metal thermocouple instrumentation exists in the cold-leg balance lines, fluid thermocouples are employed to obtain pseudo-metal temperatures. Each cold-leg balance line is divided into two metal segments. Table 4.4-8 lists the following data required in the metal energy calculations for each metal segment: fluid thermocouple channel ID for the pseudo-metal temperature, metal mass, outer surface area, and mean surface area. In addition, Table 4.4-9 lists the following data required in the metal energy calculations (which are not on a segment basis): ambient temperature channel ID, insulation tidekness, and insulation mean thermal conductivity. O m:wxnec-4u344.-4 anon: moo 395 4.4-20 REVislON: 1
.q The left-hand side of the metal energy conservation equation is approximated as follows:
Q d[M x c, x (T-T,,,)],,tcaq _ A[M x c, x PT,,,)]==LCut. 4.4-69 dt At where (subscripts n-1 and n denote two consecutive time points):
"~~
A[M x c, x (T-T,,,)]meistCut, , T"*'C'8 4=" ~ I==tcL84.a- 4.4-70 f[ MmetalCLBL, npm,q, x g _g g where: Nm.tetet = Number of metal segments in each cold-leg balance line, (2) The metal segment specific heat capacity at constant pressure is calculated as a function of the metal segment temperature from the metal c, versus temperature tabular data listed in Table 4.4-10, via O linear interpolation within the table. No extrapolation at either end is performed; the first and last table points define the applicable range (the first point is [T ,,, c,,,,,], and the last point is [T,., c,. ]). This completes the discussion of the calculations related to the left-hand side of the metal energy conservation equation. On the right-hand side of the metal energy conservation equation, the total metal-to-ambient heat I transfer rate is equal to the sum of the individual metal segment metal-to-ambient heat transfer rates: N. lmetak ambent),CLB( * (metak ambwnt),CLBL, m.i f The integrated metal-to-ambient heat transfer is also computed for use in the overall system energy balance. 4 m: Wee 4u344.-4 anon:ib-too395 4.4 21 REVISION: 1
The metal segment m metal-to-ambient heat transfer rate is given by the following: i (Tami.ctsy - T,,,,,,,) 04ma .mb,-tctat , 4.4-72 C* (Router.metalCLB6 + RmsulemLCLB6) where: C2 = Conversion constant, hours to seconds The metal segment m outer surface heat transfer resistance is given by the following:
.0 R = 4.4-73
- +
A,,,,r_m.i i.ctat, x (H,,,,,,,,,,, ,,,,,,ct,y H,,,,,_,,a_,,,,,, cut,,) Note that for ambient heat transfer considerations, each cold-lC balance line metal segment is treated as a horir.ontal cylinder (also recall that the cold-leg balance lines are insulated). A natural convection heat transfer correlation"* of the following form is employed: 4.4-74 Nu = C (Gr Pr). where, for horizonal cylind:rs: C = 0.13, m = 1/3; for (Gr Pr) > 10'(turbulent range) 4.4-75 C = 0.53, m = 1/4; for (Gr Pr) < 10' (laminar range) For the OSU test facility air at atmospheric prersure and ambient temperature (averaged over the tests), the laminar range (Gr Pr < 10'; C = 0.53, m = 1/4) applies at the outer surfaces of the cold leg balance lines, and the above correlation simpli0es to the following expression for the convective heat transfer coefficient for each metal segment m: ( TIM H ouar.=>v mmtetsy = 0.28 x 4.4-76 x (lT,,,caq, - T,3l) "
\ }
O m%sxNec-4m44. 40.no :ib.ioo395 4,4 22 REVISION: 1
- . . . . . - - . . . -- ~ ~ - . - . - . . . - -- . . , . .
7 .- r $: Using an outer diameter of 3.25 in. 'for the cold-leg balance lines, the above correlation further V ' simplifies to the following expression:
'H ,,,y,,,ct,g = 0.21 x (lT,,cuy - T,,l) /d 4.4 77 r
I
, i Employing thh assumption of radiation heat transfer between infinite parallel planes in the limiting - >
case of a convex object completely enclosed by a very large concave surface (typical assumption for the calculation ~of radiation energy loss from a hot object in a room), with an assumed emissivity of
- 'O.8 for the insulated cold-leg balance lines, the metal segment m outer surface heat transfer coefficient due to radiation reduces to the following
- 1 1 e 4
Hm,_ con,_m tcuq, = 1.3704 x 10 x [ Tit,.tetat, _ + (T,,w + C 3)2] l
! x [T,t,,ct,y + (T,,, + C 3)] ,
l' where: C3 = Conversion constant,459.6 ('F to 'R) )
'Ihe coefficient 1.3704 x 10 4is the product of the Stefan-Boltzmann constant (1.713 x 104 Btu /(hr.-ft.2 E - *R')) and the emissivity (0.8).
'the insulation temperature for metal segment m is approximated by the following (assumes that the i' temperature drop in the insulation is 90 percent of the overall temperature drop):
L I T,,g,,tcaq, = 0.1 x (T,,,.i,Cuq, - T,,%) + T,,, + C3 4.4-79 j i The metal segment m insulation heat transfer resistance is given by the following:
=
malamlCLBL, e R,,g,,tct, t,, Amean_ mesal.CLB(, x kmalamLCLB4 a where: Ax =- Insulation thickness,in. g ; k = ' Mean thermal conductivity of insulation, Btu-in/hr.-ft.2,op l
. m:pe-4us44 -40.noa: b-too395 4,4 23 REVISION: 1 l
I Finally, having calculated both the change in stored metal energy (from the left-hand side) and the metal-to-ambient heat transfer rate (from the right-hand side), the metal energy conservation equation is rearranged to solve for the metal-to-fluid heat transfer rate. Recall that the latter is used in the fluid energy conservation equation calculations which were described earlier. The integrated metal-to-fluid heat transfer is also computed for use in the overall system energy balance. This completes the discussion of the calculations related to the right hand side of the metal energy conservation equation. O l 1 mhrWXNec-4W44w-40.non:lb-100495 4,4 24 REVISION: 1 l
s ,r 3 , t, ! TAllLE 4.41 INSTRUMENTATION EMPLOYED FOR CMT FLUID CALCULATIONS Description CMT1 CMT2 Levels - main LDP-507 LDP-502
- alternate LDP-501 LDP-504 LDP-503 LDP-506 LDP-505 LDP-508 Top pressure I'F-501 EYF-502 Temperatures /elev (in.) TF-501 0.300 TF-504 0.300 (Bottom to top) TF-507 21.170 TF-510 21.170 TF-509 37.191 TF-512 37.191 TF-513 40.891 TF-516 40.891 TF-515 43.711 TF-518 43.711 TF-519 46.531 TF-522 46.531 TF-523 49.351 TF-526 49.351 TF-527 52.171 TF-530 52.171 r r
( ) TF-547 54.541 TF-548 54.541 TF-529 56.911 TF-532 56.911 DVI line - flow out bottom FMM-501 FMM-504
- temperature TF-549 TF-550 f%
f i NJ m :\agNXAsec-4\2344 w-40. non : I b- 100395 4,4 25 REVISION: I
l l l l T'J1LE 4.4-2 l VOLUME VERSUS IIEIGIIT TAllLFS FOR CMT FLUID VOLUME CALCULATIONS l CMT1 CMT2 lleight Volume IIeight Volume (in.) (in.') (in.) (in.') 0.00 0.0 0.00 0.0 3.25 122.0 0.50 69.0 6.25 393.0 3.50 294.0 9.25 1103.0 6.50 1080.0 12.25 2009.0 10.50 1956.0 15.25 3006.0 13.50 2942.0 17.00 3599.0 15.00 3455.0 31.50 8888.0 30.00 8957.0 36.75 10786.0 35.00 10763.0 44.25 13484.0 42.50 13451.0 48.75 15105.0 46.625 14961.0 51.75 16130.0 50.00 16050.0 54.75 17036.0 53.00 17061.0 57.75 177N.0 56.00 17706.0 58.90 17945.0 59.00 17911.0 0; mhpoomsee-4u344 40..on:ib.ioo395 4,4 26 REVISION: 1 1
TAllLE 4.4-3 CMT METAL WALL TliERMOCOUPLE INSTRUMENTATION Metal Segment CMT1 CMT-2 1 TW-503, TW-501 TW-5N, TW-502 2 TW-507 TW-505 TW-508,1W-506 3 TW 511, TW 509 TW-512, TW-510 4 TW 515, TW-513 TW-516, TW-514 5 TW-521 TW-519, TW-517 TW-522, TW-520, TW-518 6 TW 525, TW 523 TW-526 TW-524 7 TW-529, TW-527 TW-530, TW-528 8 TW-533, TW-531 TW-534, TW-532 9 TW-535, TW-537 TW-538, TW-536 10 TW-541, TW-539 TW-542, TW-540 11 TW-545, TW-543 TW 546, TW-544 12 TW-551, TW-549, TW 547 TW-552, TW-550, TW-548 13 TW-555, TW-553 TW-556, TW-554 s Note: Thermocouples for each metal segment are listed in the following order: inside surface, centerline (if applicable; only if three thermocouples are listed), outside surface. I \ mAapeakec-4\2344w-40.non:Ib too395 4,4 27 REVISION: 1
l l l l l l TAllt.E 4.4-4 DATA FOR Ch1T 51ETAL ENERGY CALCULATIONS CAIT 1 or ChfT 2 Afetal Metal Segment Metal blass (Ibm) Inner Surface Outer Surface Area (ft ) Area (It') 1 347.4 5.50 6.28 2 91.3 1.70 1.89 3 91.3 1.70 1.89 4 187.I 3.48 3.87 5 77.9 1.45 1.61 6 77.9 1.45 1 61 7 155.7 2.90 3.23 8 72.1 1.34 1.49 9 72.1 1.34 1.49 10 162.9 2.80 3.15 11 123.0 1.83 2.09 12 86.4 1.03 1.57 13 86.4 1.03 1.57 Ambient Temperature ID [*F) TF-006 l l l 1 mAquxNec-4uu4 40.non:ib-too395 4,4 28 REVISION: 1
I 4 TABLE 4.4-5
'v/ SPECIFIC llEAT CAPACITY VERSUS TEMPERATURE TABLE FOR CMT METAL ENERGY CALCULATIONS M etal C , Metal Temperature
[Htu/0bm *F)] ['F] 0.1085 70. 0.1109 100. 0.1175 200. 0.1223 300. 0.1256 400. 0.1279 500. 0.1297 600. I l TABI.E 4.4-6 INSTRUMENTATION EMPLOYED FOR COLD-LEG
,q HALANCE LINE FLUID CALCULATIONS (d
Description CLBL1 CLBL-2 Level LDP-509 LDP-510 Temperatures: bottom TF-533 IIPS201-3 top TF-531 TF-546 Pressure l'T-501 IrT-502 A U m:\ap60wec-4u344w-40.non:th loo 395 4,4 29 REVISION: 1
1 l TAllLE 4.4 7 VOLUME VERSUS IIEIGIIT TAllLES FOR COLD-LEG BALANCE LINE FLUID VOLUME CALCULATIONS CLBI-I CLitL 2 Ileight Volume IIeight Volume (in.) (in.') (in.) (in.') 0.000 0.000 0.000 0.000 11.000 16.453 9.000 13.461 44.440 49.398 38.440 42.465 45.560 66.147 39.560 55.765 47.375 85.687 41.500 75.316 50.815 108.477 44.000 90.298 51.935 158.539 48.191 124.537 75 250 181.449 49.309 147.629 78.693 218.059 74.500 172.387 80.000 257.284 77.000 196.405 78.949 209.478 80.000 225.(M6 TABLE 4.4-8 DATA FOR COLD-LEG llALANCE LINE METAL ENERGY CALCULATIONS (PER SEGMENT) CLilL 1 Metal Segments Pseudo-Metal Outer Surface Mean Surface Temperature ID Metal Mass Area Area Metal Segment ('F) (Ibm) (ft.') (ft.8) 1 TF-533 42.9 9.03 5.82 2 TF-531 42.9 9.03 5.82 CLBL-2 Metal Segments Pseudo-Metal Outer Surface Mean Surface Temperature ID Metal Mass Area Area Metal Segment (*F) (Ibm) (ft.') (ft.') 1 HPS2013 31.9 6.00 5.16 2 TF-546 31.9 6.00 5.16 O m:Wec-4\2344w-40.non:lb-100395 4,4-30 REVISION: 1
I 1 l 1 l r
, i TABLE 4.4-9 4 DATA FOR COLD LEG BALANCE LINE METAL ENERGY CALCULATIONS Description CLBL1 CLBL-2 l
Ambient temperature ID TF-006 TF-006 (*F) ) l Insulation thickness 1.0 1.0 1 (in.) Insulation mean thermal 0.31 0.31 conductivity [(Btu-in.)/(br.-ft.2 'F)]
- I TABLE 4.4-10 SPECIFIC llEAT CAPACITY VERSUS TEMPERATURE TABLE FOR COLD-LEG BALANCE LINE AfETAL ENERGY CALCULATIONS Metalc, Metal Temperature
[ Btu /(lbm 'F)] (*F) l ,p 0.1085 70 O.1109 100 0.I175 200 0.1223 300 0.1256 400 1 , 0.1279 500 0.1297 600 i i i
,{
%)
mhransec-4u344w-40.aon:Istoo395 4,4-31 REVISION: 1 1 l l
. 4.5 In Containment Refueling Water Storage Tank (I.RWST)
. For the' AP600, the IRWST is a reservoir of coolant held inside containment at ambient containment 1 pressure. Once the RCS has been depressurized by actuation of all four stages of the . ADS, the IRWST inventory flows to the core by gravity through piping and valves connecting the IRWST to the DVI lines. The IRWST injection phase continues until the IRWST inventory is' depleted and ' -
recirculation of the containment sump inventory through the RCS is initiated. The OSU test facility ' simulated the IRWST with a tank and associated piping. In addition to piping . > and valves connecting the IRWST to the DVI lines, the IRWST simulation accounted for the following AP600 features:
- A sparger to distribute flow from ADS 1-3 to tie IRWST
- A submersed heat exchanger to simulate the operation of PRHR during simulated events
. An overflow line between the IRWST and the ' containment sump tank to simulate the interaction of the IRWST directly with the sump De operation of each of these features is accounted for in the mass and energy balance performed for A the IRWST simulation.
4.5.I General Mass and Energy Balance Formulation The liquid mass inventory in the IRWST is calculated as: Mrawsr " Pr xV rawsr 4.5-I De vapor mass inventory of the IRWST is calculated as: Meawsr " P, x Vaawsr 4.5-2 where the subscripts: f = liquid phase of water g = steam Flow meter FVM-701 is located in the exhaust line from the IRWST. If the measured flow from FVM 70I is less than an input threshold value, then no steam property calculations are performed. Thus, no steam inventory is modeled in the tank until there is a measured exhaust flow from FVM-701. mAap600sec-4\2344w-41 mon:lb.100395 ~ 4,$.1 REVISION: 1
.-c
The general mass balance equation for the IRWST is expressed as: l 1 dM i
"" Aos - sm vrxr -b OVRFLW ~ IRWSTINs " ~
d l where the subscripts: ADS = Mass flow rate, both vapor and liquid, from ADS 1-3 separator STM VENT = Vapor mass flow rate exhausted from IRWST through FVM-701 OVRFLW = Liquid mass flow rate through IRWST overflow line measured by FMM-703 IRWST INJ = Total mass flow rate injected from IRWST through both injection lines Minitrum flow criteria: If the mass flow rates injected from the IRWST are less than a threshold value, then the mass flow rates are set to zero A similar adjustment is made to Mans' Ms m VENT' EUd OVE W' If M,,y is, < Mmwn txi low, then M,,y g, = 0. If Mios < M,os tow, then M,os = 0. IfMom,<Moym, tow, then Moyg,1, = 0. Mmwn mi low " M ADS LOW, f OVELW LOW = 0.5 gpm M,o, tow, , = 0.5 cfm Similarly, the general energy balance on the IRWST is written as: [ M' c dT' = Qrana xos hios +
- KiNi sm vsxt h, - N1 0vartw h,-Ni,wng,h,,yg,
- Q,a - Qm 4.5-4 Qrana is caiculated using Equation 4.5-4.
O m:\ap600\sec-4u M4 w-45.aon: Ib- 100395 4.5-2 REVis!ON: 1
S O -( ). where: Orasia = Heat addition rate from PRHR, Bru/sec. Q,a . = Energy loss to metal QAwa = Energy loss to ambient environment and the subscripts: J = Control volume within the IRWST associated with axial thermocouple locations The IRWST simulation may experience four modes of operation: Case 1: IRWST inventory is subcooled; ADS has not been actuated; there is no overflow of IRWST inventory. Case 2: At least some portion of the IRWST inventory is saturated (due to actuation of PRHR); ADS has not been actuated; there may be overflow from the IRWST. ; Case 3: At least some portion of the IRWST inventory is saturated; ADS has been actuated; both PRHR and ADS energy is added to the IRWST. .O ! O Case 4: After ADS actuation, IRWST inventory is injected into the primary system. The simplification of the general mass and energy balance equations for these four cases are as follows. 4.5.2 Case 1 For this case, there is no liquid or vapor flow out of the IRWST, implying the following:
- No injection flow from the IRWST l No liquid swell (due to heating) that provides for liquid flow through the overflow lines No steam generation and, therefore, the bulk liquid temperature of the IRWST inventory is lower than its saturation temperature (T, < TSAT)
According to the first two items, measured flow from the IRWST is given as: 6t ryy,,g = Nipyy,,,=Ni ggy ,g = si pvu ,g = Nim = Niguy, = 0 4.5-5 O m:Wec-4u344w-45.noa:ib. loo 395 4.5-3 REVISION: 1
where the subscripts: FMM-701 FMM-702 FMM-703
= Flow meters in lines containing both liquid and vapor flow from the IRWST FVM-701 FVM-601 FMM-601 With no mass entering or leaving the IRWST, the time rate of change in mass of the IRWST is defined as:
dMS " = 0 4.5-6 dt Under these conditions, the energy balance on the IRWST is expressed as: { M; c, dTi 4.5-7 dt
.Q""
The fluid thermocouples used in calculating the mass inventory of the IRWST and their respective elevations referenced to the top of the ceramic filler in the IRWST are given in Table 4.5-1. The difference form of the preceding equation is: { M c,j ATj = Q,, At 4.5-8 Liquid mass inventory of the IRWST is then calculated as follows: Step 1: Calculate a temperature-compensated water level in the IRWST using outputs from level transducer LDP-701 and the thermocouples identified in Table 4.5-1 as input to the subroutine developed to perform this calculation. If the temperature compensated level is less than the threshold value of 90.5 in., then M oyxa,w = 0. Step 2: Calculate a pressure and density gradient for the water inventory in the IRWST using outputs from total pressure transducer FT-701, the compensated level measurement, and the thermocouples identified in Table 4.5-1. This is accomplished as follows: O m:pecaumwas.noo: b too395 4.5-4 REVISION: 1
O) g First, the IRWST is divided into regions. Located in the middle of each region is one of the thermocouples identified in Table 4.5-1. The boundary between adjacent regions is defined as the midpoint between adjacent thermocouples. The bottom boundary of the region at the bottom of the IRWST is the top surface of the ceramic filler. The top boundary of the region at the top of the IRWST is the top of the IRWST. 4 Next, water colurnn in the IRWST is partitioned into the regions defined above. In the event that the water level resides in a region, but is below the thermocouple in that region, then the height of water in that region is applied to the adjacent region. Then, the pressure recorded by instrument PT-701 is applied to the top surface of the IRWST inventory and assumed to be representative of the pressure in that region. The liquid density in the first region is calculated as: p , = p(P, , T 3) 4.5-9 where Pi is the pressure measurement of instrument FT-701 and Ti is the first thermocouple below the surface of the compensated water level. The pressure at the top of the next lowest region is then calculated as: O V P, = P,,, + .g_ jp , Az,.i 4.5-10 E. I where the subscript: l j = Control volume within the IRWST associated with axial thermocouple locations 1 The calculations defined in Equations 4.5-9 and 4.5-10 are repeated for each of the ; regions in the IRWST that are below the compensated level indicated by LDP-701, Step 3: Liquid mass in the IRWST is then calculated as the sum of the masses of each region: M mwsr *[Pj(z) V 4.5-11 ) where: V(z) = Volume of the IRWST as a function of height from the top surface of the ceramic fill located in the bottom of the tank () m:wah4um45.non:ib.ioo395 4.5-5 REVISION: 1
i l 4.53 Case 2 In this case, the IRWST liquid inventory is warmed sufficiently due to PRHR so that the liquid volume has swelled above the overflow line. The temperature of the IRWST inventory at or near the ' top of the IRWST is expected to be at or near saturation. Thus, there may be both liquid and vapor flow out of the IRWST. This situation may be described as:
- There is no injection flow from the IRWST
- Liquid flow may occur through the overflow line, measured by FMM-703, due to volume swelling resulting from heating
* 'Ihere may be steam generation, measured by FVM 701 From these three conditions, the mass balance on the IRWST given by Equation 4.5-3 is simplified as:
dM**" = - Nim.,o, - Niruu-703 4.5-12 dt Noting that the left-hand side of the previous equation may be approximated as a difference: dM,,y M,,y,, - M,,y, ,,, O dt t, - t, ., where the subscript: i = Index of data and time arrays The mass depletion of the IRWST may now be written as: M,,y, , = M,,y, ,,, - Nigyy,,oi, , At - Nigyy ,03, , At 4.5-14 where: At = t, - t,.i e m:peaumas.no.:ib ioo395 4.5-6 REVISION: 1
a l l 1 i
'I 7-
-( )- Density and enthalpy of vapor and liquid flow are evaluated using the ASME steam table routines and the outputs from the following instruments: . l Pressure Temperature ! Vapor (steam) PT-701 TF-722 Liquid (overflow) IT-701 TF-723 The units of measurement for the vapor and liquid flow meters are ft.8/ min. and gpm, respectively. The vapor mass flow is calculated as: S rvu-7oi " P <rr-7oi. TF-722> W(,yy.,,,, x C, GM l Similarly, the liquid mass flow rate is calculated as: Meuw-tos
- Per-7oi, TF-m> W,y,,o3, x C 2 4.5-16 where:
b
.g_ C = Conversion constant, min. to sec.
C2 = Conversion constant, gpm to ft.8/sec. The rate of energy transfer associated with the IRWST inventory is simplified from Equation 4.5-4 and expressed as: Cp dT i=Qp ,-My .70, h, - Mguy.,o3 h, 4.5-17 The difference form of the previous equation is written as: { M, c, AT j= QemAt - M rvu-7oi h, At - M gyy,,o3 h,At 4.5-18 4.5.4 Case 3 1 In this case, the operation of the ADS increases the mass in the IRWST. Also, operation of PRHR and the ADS may heat up the IRWST inventory. The combination of added mass and inventory heatup may cause the IRWST liquid volume to swell to above the overflow line. The temperature of the IRWST inventory is expected to be at or near saturation, and not all ADS flow may be condensed. 'O / I Thus, there may be both liquid and vapor flow out of the IRWST, implying the following: V m:\ap60(Avec-4\2344w 45.non:lb 100395 4.5-7 REVISION: 1
- There is no injection flow from the IRWST to the DVI line
- ADS flow, possibly both steam and water, is being exhausted into the IRWST
- Liquid flow may occur through the overflow line, measured by FMM-703, due to volume swelling resulting from heating and injection of water from ADS 1-3
- There may be steam generation, measured by FVM-701 From these four conditions, the mass balance on the IRWST is given as:
dM 4M As ~b evu-7oi -M ruw-7o3 dt The ADS flow may consist of both vapor and liquid components: 4.5-20 MADS " rVM 41 + rMM W1 where the subscripts: FVM-601 = Vapor flow meter downstream of ADS 1-3 separator FMM-601 = Liquid flow meter downstream of ADS 1-3 separator Expressing the left-hand side of Equation 4.5-20 as a difference and expanding, the liquid mass inventory of the IRWST is calculated as: M m ,37,,= M ,,37, ., +Mgyu i, , At + Mruwei. , At - Mn .y 7,i, , At - M ruu-7oa. i At 4.5-21 Similarly, the energy balance for this case may be written as: {McdT. p .
~
dt " * ' ' '**'*'* ""~ The enthalpy of the vapor and liquid components of the ADS flow, measured by FVM-601 and FMM-601, respectively, is evaluated using the ASME steam table routines and data from the following instruments: Pressure Temperature Vapor phase PT-605 TF-617 Liquid phase PT-605 TF-616 I m%wxNec-4u)44.-45..on:tb. loo 395 4.5-8 REVISION: 1 ) I 1
..p. '( ,
The difference form of the energy equation is written as:
. [ M; cp ATj = spyy , h, At + s pyy o, h, At - s pyy .,o, h, At - s pyy ,o3h,At 4.5-23 4.5.5 Case 4 i
This case occurs late in the transient: e PRHR is no longer active and ADS 1-3, while open, has low flow
= While there may be some steam venting of the IRWST volume, there is most likely no liquid overflow
- At this time, the IRWST may be injecting into the DVI lines The mass balance on the IRWST is given by Equation 4.5-3 where:
Mayn x = bewu-7oi + bruu-702 4.5-24 The mass balance for this case is expressed as: d Mawsr. i " Mmwsr..-i+brvuei. At + s pyy oi, At - s pyy .,oi, At -s pyy .,o3, At 4.5-25
~Sruu-7oi. At -sruu 7n At Similarly, the energy balance is written as:
{Mcdr p
" #' 8 ""~' ' " ' ' ' 8 ** ~7 ' '
.t
- s pyy .,o, h, - sruu-702 h, OM Expressing the preceding equation in its difference form:
i { M c,Mj =$rvu-6oi h, + s pyy oi h, - s pyy .,o, h, 3 i 4.5-27
~
FMM-703 f ~ FMM-701 f ~ FuM-702 f m:pc-4u344 45. on:ib.ioo395 4.5-9 REVISION: 1
4.5.6 Direct Vessel Injection Line Flow Reversal ( Liquid flow reversal can occur in the DVI 1. The total injection flow is calculated as: I Mmwn w " Mmwsr m. i + M awsr m.2 4.5-28 Mmwn w. , - p W,yy.,,n xC2 4.5-29 Mawsr w.2 = p Wmuu-?o23 xC 2 4.5-30 where: p = water density of the lowest-most region calculated in Equation 4.5-9 C2 = Conversion constant, gpm to ft.'/sec. Flow reversals can occur for injection line number 1. If the measured flow is less than a threshold q value, Wyyy ,,g < W,,37 , go, 4.5-31 l then the reverse DVI flow is calculated as: i MawsT w. a " M awsT w. mi -M ACC,4 - SUMPi ~ CMr,1
- l where:
MmwsT w. m, = p W auu-2os xC 2 4.5-33 l l where the density of injection liquid is calculated at pressure and temperature conditions using PT-109 and the minimum temperature from TF-ll3 and TF-II5. sg , is calculated using Equation 4.3-15. M suur. I is calculated using Equation 4.9-9. Mcur. , is calculated using Equation 4.418. I m:wwcc-4unsw.45mo:ib ioo395 4.5-10 REVISION: 1 1
,d If Sawrr w,'mer=4 < 0.0 and if $ guy.,oy < M,,y , w, then b mwrr w. : "S mwn w. r.vor=d 4 5-M otherwise, s,,y3,; = 0.0 4.5-35 Note: Flow reversal does not occur for injection line number 2. i iib mww-7en 2b mwst w t.ow then i M mwsr m. =pWyyy , ,x C 2 4.5-36 i
; otherwise \ Mmwst m. 2 = 0.0 4.5-37 4.5.7 Energy Loss due to Ambient Heat Transfer Rate An equivalent heat transfer coefficient, accounting for heat resistance due to any insulation applied to the outside surface of the IRWST volumes and natural convection from the outside surface to the ambient, is used to evaluate the heat loss to ambient from the metal surfaces of the IRWST. De same metal surface temperatures used to evaluate metal heat storage of the IRWST is used to calculate heat j loss to the ambient. Surface areas included in this calculation account for IRWST and associated I piping.
De metal-to-ambient heat transfer calculation methodology for the IRWST is similar to methodology used for other insulated components of the system. Each IRWST metal segment is treated as an insulated vertical cylinder. De turbulent range form of the natural convection heat transfer correlation from Reference 16 applies, based upon the conditions of the OSU test facility air at atmospheric ; pressure and ambient temperature (averaged over the tests). In the radiation heat transfer correlation, an emissivity of 0.8 is assumed. j De detailed calculations and equations are identical to those for the metal-to-ambient heat transfer for the cold-leg balance lines in Subsection 4.4.6, the only exception is that in the final expression for the
\g natural convection heat transfer coefficient, the coefficient multiplier 0.21 is replaced with 0.09 and the
=%**8" 42M4*-41aon:lt 100395 4.5 REVISION: 1
{
4 I l exponent 1/4 is replaced with 1/3 (these differences are due to both the vertical orientation and the turbulent range of applicability for the IRWST). The total energy loss to ambient rate is equal to the sum of the individual metal segment heat transfer rates: L 4.5-38 Oambient " b Q. mal T,,t, - T*"6""' 9* , 4.5-39 (R,,,,t,, +R .t) 1.0
= 4.5-40 R'"T "*'" A metatm x (H conv.metata +Hrad_metalm)
H,,,,,,,, = 0.09 x ( l(T,,,, - T.,,,) l)" 4.5-41 H,,,,, t, = 1.713x104 x e x [T' ,.t.
+ (T,,,,,, + C)2] x 4.5-42
[T,,,,, + (T,,,., ,, + C)] The insulation temperature for metal segment, m, is approximated by: T,_,,,, = (0.1 x (T,, - T ,,) + T.,,,) + C 4.5-43 This assumes that the temperature drop in the insulation is 90 percent overall temperature drop: R m .imt" = Ax'""' "*' 4.5-44 Amean_metalm xk inml_ metal where: C = Conversion constant, 'F to *R Ax i_,u = Insulation thickness A..t,, = Segment outer surface area mAap600\sec 4\2M4w-45.non:lb-100395 4.5-12 REVISION: 1 l
p A. . .. = Mean segment area e = Emissivity of insulation layer (0.8) k_i = Mean thermal conductivity ofinsulation T 3,,,,
= . Ambient temperature N.,,r = Number of surface metal segments R.,r_,u,, = Metal segment outer surface heat transfer resistance 11
. .u = Metal segment outer surface heat transfer coefficient due to convection l i,,(,,,,t,, = Metal segment outer surface heat transfer coefficient due to radiation 4 d 1.713 x 10 = Stefan Boltzmann constant BTU /hr. ft' *R 4.5.8 Energy Loss to Metal L
Qmetal
= cp metal xM metalm x T'"* '" - T" """ 4*5-45 m.i t, - t, ., _
where: c,, ,,,,, = Ileat capacity of the metal M,u,, = Mass of the metal segment () m T,,,,,,3 t, =
= Metal temperature at time n Time at time-step n Nu o = Number of metal segments 4.5.9 Fluid Stored Energy For use in the overall system energy balance calculations, the fluid stored energy in the IRWST is !
given by the following: U ng.mwn = Ur .mwn + U,,,,y 4.5-46 where: L. Ur.mwn = {
)=l M,,,,y x c ,,,, , ,y x (T,,,,y - T,,,) 4.5-47
- where
N, mwn = Number of zones in the IRWST (12) maap60(Asec 4Q344w-45.noo:Ib 100395 4.5-13 REVISION: 1
U, mwn is not considered here. Note that there is a steam inventory in the IRWST during ADS 1-3 actuation. Ilowever, the stored energy due to steam is negligible compared to the stored energy due to the large volurne of liquid in the IRWST. O J l O EkMwo4\2344w-45.non:lb-100'95 4.5-14 REVISION: 1
^
i i l [ i [ TABLE 4.51 IRWST MASS AND ENERGY CALCULATIONS IDINI'IFICATION OF FLUID THERMOCOUPLES AND ELEVATION Elevation'" Index Thermocouple ID (in.) 1 TF-701 1.00 ! 2 TF 702 8.99 3 TF-703 16.97 4 TF-7N 26.85 5 TF-705 36.73 6 TF-706 46.61 l 7 TF-707 56.49 8 TF-708 66.36 9 TF-709 76.24 10 TF-710 87.36 11 TF-711 98.47 12 TF-712 109.59 Note: (1) Elevations referen d from top surface of plate in bottom of IRWST and run vertically upwards. TABLE 4.5-2 VOLUME VERSUS HEIGHT TABLE FOR IRWST FLUID VOLUME CALCULATIONS IIelght On.) Volume (in.') 0.0 0.0 10.5 73,391.0 42.0 297,302.0 56.625 403,788.0 74.5 531,187.0 90.0 640,577.0 93.75 666,472.0 115.0 749,140.0 0% V m:\ap60(Asec-4\2344 w.45.non : l b.100395 4,$.15 REVISION: 1
i TABLE 4.5 3 Ol l DATA FOR IRWST METAL ENERGY CALCULATIONS (PER SEGMENT) Metal Metal Temperature ID Metal Mass l Segment ('F) (Ibm) 1 TFM-701 2148 2 TFM-702 1353 3 TFM-703 2394 4 TF-719 146 TABLE 4.5 4 DATA FOR IRWST METAL ENERGY CALCULATIONS Description Data Ambient Temperature ID ( F) TF-006 Insulation Thickness (in.) 1.5 Insulation mean thennal conductivity 0.141 [(Btu-in.)Ahr.-ft.2 F)] TABLE 4.5-5 DATA FOR IRWST ENERGY LOSS DUE TO AMBIENT IIEAT TRANSFER Metal Temperature Surface ID ('F) Outer Surface Area (ft.2) Mean Surface Area (ft.') 1 TFM-701 107.72 53.I1 2 TFM-702 89.46 44.10 3 TFM-703 75.32 86.68 l l mhr60msee-4u344 -45.non:thtoo395 4.5-16 REVISION: 1
]
l
O 4.6 Automatic Depressurization System 13 Separator Q. The ADS provides a means of depressurizing the RCS in a controlled, staged manner through the use of four pairs of valves, with each valve pair sequenced to open at different primary system pressures.
'Ihe first three pairs of valves, called ADS stages 1,2, and 3 (ADS 1-3), are located in parallel piping paths running from the top of the pressurizer to the IRWST. 'This portion of the ADS operates independently of the fourth and final stage of the ADS.
The ADS 13 flow path was simulated in the OSU test facility by three valves (one valve each having a scaled stage 1, stage 2, and stage 3 flow area), a steam / water separator tank, a vortex (vapor) flow meter, a magnetic (liquid) flow meter, and associated piping. During testing, flow through the ADS 1-3 flow path was measured by separating the vapor and liquid components of the flow, measuring the flow rate of the component flows, recombining the flows, then directing the total metered flow to a sparger located in the IRWST. Flow through ADS 1-3 may be calculated as: ADS 1-3 SEP M ADS i-3 "SADS 1-3, f + ADS 1-3, g , 4,6.] where the subscripts: f = Liquid phase of water g = Steam
)
ADS 13 = Stages 1 through 3 of the ADS l SEP = Steam water separator tank for ADS 1-3 Energy is transported out of the primary system by both the vapor and liquid flows. Also, the stored energy of the ADS 1-3 separator inventory may change due to changes in the amount of liquid and I vapor in the separator, changes in temperature and pressure of the liquid or vapor inventory in the seprator, or a change in the temperature of the separator tank metal mass. Accounting for these terms, the energy rate equation for ADS 1-3 flow may be expressed as: AS SP QADS1-3 " OADS 1-3. f + OADS 1-3. g *C P
+OADS l-3, METAL, + OADS 1-3. AMB 4.6-2 ,
l L, mA,600cN11.noa:1b. ton 395 4.6-1 REVISION: 1 1
where the subscripts: METAL = Metal mass of the ADS 1-3 separator tank and associated piping AMD = Energy loss to ambient environment 4.6.1 Automatic Depressurization System 13 Separator Liquid Inventory he ADS 1-3 separator is, in its simplest form, a tank. Liquid inventory in the ADS 1-3 separator is monitored by a level transducer. De functional steps and associated system of equations for operating on the output from the ADS 1-3 separator level transducers to calculate liquid mass in the ADS 1-3 steam / water separator tank follows: Step 1: Compensate the readings from the ADS 1-3 separator level transducer listed in Table 4.6-1 to account for temperature differences between fluid in the separator tank and fluid in the reference leg of the instrument line. The local pressure and fluid temperature instruments to be used to accomplish the compensation are also identified in Table 4.6-1. Step 2: The local pressures and temperatures measured using the instruments identifled in Table 4.6-1 are used as inputs to the ASME steam tables to calculate the density of the liquid and vapor in the ADS 1-3 separator: p,os 3,7, = p, (PT-605, TF-616) O p,os i.s , = p, (PT-605, TF-617) 4.6-3 where: PT-605 = Data channel ID for local pressure measurement in the ADS 1-3 separator tank TF-616 = Data channel ID for local liquid temperature measurement, *F
'IF-617 = Data channel ID for local vapor temperature measurement, 'F As identified previously, both liquid and vapor flow meters have an associated local fluid temperature used to evaluate the thermodynamic properties of the liquid and vapor phases in the ADS 1-3 separator.
Step 3: Using the compensated liquid level and the ADS 1-3 separator volume as a function of height (see Table 4.6-2), determine the volume of liquid in the ADS 1-3 separator as: e V, ,,3 sep,, = V(1),os 3 3 se, x LDP-610cour 4.6-4 maapemawii. oa:tt,. loo 395 4.6 2 REVISION: 1
4 1 (D l g where: V(1) = Volume of the ADS 13 separator as a function of elevation, in.8/in. LDP-610m = Compensated fluid levels data from level transducer LDP-610, in. l l Step 4: Liquid mass inventory in the ADS 1-3 separator is now calculated as: M,os i.3 sy, , = p,os i 3, , x V,03 , 3 3y, , 4.6-5 Vapor mass in the separator is then calculated as: M,os i.3 3y, , = p,os i 3, , x (V Aos i-3 se. ror -VAos :-3 se. ) 4.6-6 where the subscript: TOT = Total volume of ADS 1-3 separator, ft.' Step 5: The rate of change in mass inventory of the break separator tank may be approximated by A e differencing two consecutive calculated values ofliquid and vapor mass: I dM Aos -3 se
, AM Aos i 3s e , M ,os i 3 s y,,,, -M Aos -3 se. t. 4-i dt At t, - t, .,
M,03 i 3 sy, ,, , - MAos :-3 se. s. 5-3
. 4.6-7 t, - t,,,
where the subscript: 1 = Index of data and time arrays e m:WwI t.non:tb ioo395 4.6-3 REVISION: 1
4.6.2 Steam Flow Rates The density of steam in the ADS 1-3 exhaust line, evaluated using the pressure and temperature inputs from the data channels identified in Table 4.6-3, is used to calculate the steam mass flow rate through the ADS 1-3 valves as: NI ADS 1-1 g *C 1 k ADS l-1 g W FVM el 4'6-8 where: W= = Volumetric flow rate of steam, ft.'/ min Ci = Conversion constant, minutes to seconds and the subscript: FVM-601 = Instrument channel ID for steam vapor flow meter in line between ADS 1-3 separator and sparger 4.6.3 Liquid Flow Rates Liquid from ADS 1-3 is ducted from the separator through a magnetic flow meter into a header where O it is mixed with steam flow before passing to the sparger in the IRWST. Mass flow from the ADS 1-3 separator to the sparger is calculated as: NIADS 1-3, f ' C2 b ADS 1-3, f W rwuei 4.6-9 where: W = Volumetric flow rate of liquid, gpm C2 = Conversion constant, gpm to ft.'/sec. and the subscript: FMM-601 = Instmment channel ID for liquid flow meter in line between ADS 1-3 separator and sparger The density of the liquid passing through the flow meter is evaluated using data from the pressure and temperature instruments identified in Table 4.6-3. mAap60tA21144w ll.non:lb-100395 4.6-4 REVISION: 1 l l
d t O) 4.6.4 Total Flow Rate De total liquid and vapor flows through ADS 1-3 are then calculated as: 9
# ADS 1-3, f 4,6 ]O Q ADS l-3. f .Q FMM-601 g
- s,o, i.3, , = s gyu , + i -3. 4.6-11 e
i . The total flow through the ADS is zero when the three ADS 1-3 valves are closed. Thus, the above )
- total liquid and vapor flows are used only when at least one valve is open.
)
The total flow through the ADS (for t 2 valve open time) is then calculated as: i b ADS l-3, TOTAL
- ADS 1-3, f + ADS 1-3, 3 4*6-l2 l I t O}
, For t < valve open time, Mios 3 3 = 0. The total mass flow rate is integrating step-wise over time to calculate the total mass inventory passed by the ADS 1-3: At) 4.6-13 MADS :-3 " [ (SADS l-3 TUT 4 The flow quality of the ADS 1-3 flow is also calculated as:
- x. b^Ds 1-3 s 4.6-14 ADS 1-3, TOTAL J
{ 4 i
.t f%..
m4ar60cn2344w-11.non:ib.ioo395 4.6-5 REVISION: 1
4.6.5 Energy llalance Energy flow through ADS 1-3 consists of the following:
- Rate of change in stored energy of the ADS 1-3 separator liquid and steam inventory
- Energy transport rate from the ADS 1-3 separator by exiting steam flow
- Energy transport rate from the ADS 1-3 separator by exiting liquid flow
. Rate of change in stored energy of the ADS 1-3 metal components
- Rate of energy loss from the ADS 1-3 components to the environment The expressions for evaluating these five energy transfer or transport terms are developed in the following sections.
The combination of these five terms results in energy rate equation 4.6-2. Consistent with the influence of valve position on total flow equation 4.6-12, the energy flow rate through the ADS is zero when the three ADS l-3 valves are closed. Thus: Q,os i.3 = 0, for the t < valve open time. 4.6.5.1 Rate of Change in Stored Energy of the Automatic Depressurization System 13 Separator Liquid and Steam Inventory The rate of change of energy in the liquid and steam in the ADS 1-3 separator may be expressed as: d(higo3i.,,,T) QADS 1-3 *C P g
= c,, , (T, - T,,,) d(hi,os 3 3 sy, ,)p + c , , Af^ * 8" ' d(T,)
dt 4.6-15 d(hi,os33sy,,) d(T,)
+ ce , , (T, - T,,,) '8 ^"8 "'
dt dt where: T,.r = 32*F Expressing the previous equation as a difference and solving for consecutive data, the rate of change of energy of the fluid in the ADS 1-3 separator is calculated: m:\a[WXA2344w ll.non:lt>100395 4.6-6 REVISION: 1
O A(M,33 g3 sep T) Cp at ADS F3 SEP. f f
= c . f (T P f - Tref) g ,c.f P q ADS 1-3 SEP. f gg 4,6-16
^ S "' 88' 8
+ c . s (T P g - Tref) gg
+c.g P M ADS 1-3 SEP, g g The specific heats of the liquid and vapor phases, ce ,, and ce,, respectively, are evaluated using the output of the data channels identified in Table 4.6-2.
4.6.5.2 Energy Transport Rate from the Automatic Depressurization System 1-3 Separator by Exiting Steam Flow l l The enthalpy of the steam exhaust from the ADS 1-3 separator is determined using the output from the pressure and temperature sensors associated with FVM-601 (Table 4.6-2): h.y n i, = h, (PT-605, TF-617) 4.6-17 l FVM-601 is the data channel ID for the ADS 1-3 vapor flow meter, and PT-605 and TF-617 denote the vapor pressure and temperatures, respectively, associated with that flow meter. The rate of energy transport of the ADS 1-3 flow due to the steam component, then, is expressed as: O ADS 1-3 3
- h n.y 4.6-18 ADS 1-3.3 ,i, 4.6.5.3 Energy Transport Rate from the Automatic Depressurization System 1-3 Separator by Exiting Liquid Flow 4
The enthalpy of ADS liquid flow is determined using the output from the pressure and temperature sensors associated with FMM-601 (Table 4.6-2): hguy,1, , = h, (PT-605, TF-616) 4.6-19 (V't
.4*n23h11.mA100395 4.6-7 REVISION: 1
FMM-601 is the data channel ID of the ADS 13 liquid flow meter, and PT-605 and TF-616 denote the data channel ids for the liquid pressure and temperatures, respectively, associated with that flow meter. l The rate of energy transport due to ADS 1-3 liquid flow, then,is expressed as: j i 4*6 20 Q AD S 1-3. f ADS 1-3. f
- hWMMI, f 4.6.5.4 Rate of Change in Stored Energy of the Automatic Depressurization System 1-3 Metal Components The ADS is heat-traced downstream from the ADS valves 'Iherefore, the change in stored energy of the ADS piping due to energy loss from the fluid is negligible.
I I 4.6.5.5 Rate of Energy Loss from the Automatic Depressurization System 13 Components to the Environment Piping between the ADS 13 valves and the separator, and between the separator and the IRWST were provided with heat-tracing. This had the effect of off-setting any energy loss to the ambient environment. 'lhus, for the purpose of evaluating the rate of energy loss from the fluid to the ambient environment (through piping): Qxy , e 0.0 4.6-21 1 I t 9 m:sgenochIi.noa:1b ioo395 4.6-8 REVISION: 1 l
1 l l p l t \ TABLE 4.61 I INSTRUMENTATION TO BE USED FOR ADS 13 LEVEI.S I I INSTRUMENT CORRECTION I Lesel Pressure Fluid Location Function Transducer Transducer Temperature ADS 1-3 Density LDP-610 FT-605 TF-616 Separator compensation of levels data j TABLE 4.6-2 VOLUME VERSUS IIEIGIIT FOR ADS 13 VOLUME CALCULATIONS Ileight (in.) Volume (in.8)
- 0. 656.
86.5 2858. p 98.72 4436. 121.72 11013. I 140.62 16047. l , 1 I 146.87 17725. i l l l TABLE 4.6-3 ADS 13 SEPARATOR STEAM AND LIQUID PRESSURE AND TEMPERATURE INSTRUMENT CIIANNELS Flow Meter Flow Meter Temperature Channel Description Channel ID Pressure Channel ID ID ADS 13 separator FVM-601 Irr-605 TF-617 steam flow ADS 13 separator FMM-601 Irr-605 7F-616 liquid flow A ( I m:wvxt344w. t i.no.:t b-too395 4.6-9 REVISION: I l
- (
V 4.7 Automatic Depressurization System 4 Separators The ADS provides a means of depressurizing the RCS in a controlled, staged manner through the use of four pairs of valves, with each valve pair sequenced so that they will open at different primary system pressures. Mass, flow, and energy calculations associated with the first three pairs of valves, called ADS 13, are described in Section 4.6. A redundant fourth pair of valves, are located on each of the two hot legs and exhaust directly to the containment atmosphere. The ADS-4 valves are used to complete depressurization of the RCS to near containment pressure. The ADS-4 flow paths were simulated in the OSU test facility by a valve, a steam / water separator tank, a vortex (vapor) flow meter, a magnetic (liquid) flow meter, and associated piping from each of the two hot legs. During testing, flow through the two ADS-4 flow paths was measured by separating the vapor and liquid components of the flow, measuring the flow rate of the component flows, recombining the flows, then directing the total metered flow to a simulation of the containment sump. Flow through each of the ADS-4 flow paths may be calculated as: , dMADS 4-X SEP M ios 4.x "SADS 4-X. f + ADS 4-X, g
, 4,7.] !3 U where the subscripts:
ADS-4 = Founh-stage ADS X = Hot leg to which the flow path is connected where: 1 X = 1, for HL-1 X = 2, for HL-2 f = Liquid component of ADS-4 flow SEP = Steam water separator tank for ADS-4 g = Vapor (steam) component of ADS-4 flow The total ADS-4 flow rate is then calculated as: M Aos 4 = M ios 4.i
+M ads 4-2 4.7-2 Energy is transported out of the primary system by the ADS-4 vapor and liquid flows. Also, the stored energy of the ADS-4 separator liquid inventory may change due to a change in the amount of liquid in the separator, a change in temperature of the liquid inventory in the separator, or a change in 1, 3
() the temperature of the separator tank metal mass. Accounting for these terms, the energy equation for ADS-4 flow may be expressed as: massoasne-4u)44.-47..on: i b- t oo395 4.7 1 REVISION: 1
l l d ( M,3, ,_x ,sp T, ) O 0 4os 4-x " 0 40 s 4-x. t + 0 os 4 4-x. : + C' di 4.7-3 I
+ Q ios 4-x. MerAt. + Q40 s 4-x. Aus l
where the subscripts: 4 l l METAL = Metal mass of the ADS-4 separator tank and associated piping AMB = Energy loss to ambient environment De total energy associated with the ADS-4 (accounting for both separators) is calculated by summing the terms in the previous equation for each of the two separators. 4.7.1 Automatic Depressurization System 4 Separator Liquid Inventory The ADS-4 sepEator is, in its simplest form, a tank. Liquid inventory in the ADS-4 separators is monitored by level transducers. He functional steps and associated system of equations for operating on the output from the ADS-4 separator level transducers to calculate inventory mass in the separator tanks are: Step 1: A differential head is calculated to account for the orifice in the liquid line. De O inputs to this calculation are: pipe diameter, beta ratio of orifice to pipe restriction (p=1.0 = no orifice), pressure, temperature, and liquid flow. The output differential head is added to the measured LDP value. He readings are compensated (adjusted for orifice effect) from the ADS-4 separator level transducers listed in Table 4.7-1 to account for temperature differences between fluid in the separator tank and fluid in the reference leg of the instrument line. He local pressure and fluid temperature instruments used to accomplish the compensation are also identified in Table 4.7-1. Step 2: De local pressures and temperatures from the instruments identified in Table 4.7-1 are used to calculate the density of the liquid and vapor in the ADS-4 separator: Pr. Aos 4-x
=
p, (PT-YYY, FZZL) 4.7-4 P s. ADS 4-x kg (PT-YYY, TF-ZZV) O imwec-4u344w-47.non:ib.ioo395 4,7 2 REVISION: 1
A Q e/here: I'r-YYY = Channel ID for local pressure measurement in ADS-4 separator tank TF-ZZL, = Channel ID for local temperature measurement of liquid exiting ADS-4 separator TF-ZZV = Channel ID for local temperature measurement of vapor exiting ADS-4 separator As noted previously, both liquid and vapor flow meters have an associated local fluid temperature used to evaluate the thermodynamic properties of the liquid and vapor phases in the ADS-4 separator. 4 Step 3: Using the compensated liquid level and the ADS-4 separator volume as a function of height, determine the volume of liquid in thc ADS-4 separator as: i V,93 ,,x 3,p, , = V(1)3g3,,x 3g, x LDP-XXX 4.7-5 l , t I i where: l 'g V(1) = Volume of ADS-4 separator as a function of elevation, ft.3/ft. (Tables 4.7-4 ( and 4.7 5)
- LDP-XXXmp = Compensated fluid level data from level transducer for ADS-4 separator, ft.
XXX = 611 for ADS 4-1 separator 4 XXX = 612 for ADS 4-2 separator ! Step 4: Liquid mass inventory in the ADS-4 separator is now calculated as: 4 l h1ADS 4-X SEP ' 4'7~6 l b f, AD$ 4-X V ,f ADS 4-X SEP 4 The vapor mass inventory of an ADS-4 separator is then calculated as: hI ADS 4-s SEP, g
- P, x ( VADS 4-x ssp. nnA1, ~ VADS 4-X $9, f ) 4'7~7 4
where the subscript: TOTAL = Total volume associated with an ADS-4 separator, ft.' ]
~l V
1 m:\ap600%sec.4\2344w-47.non:lt> 100395 4,7 3 REVislON: 1 1
Step 5: The rate of change in mass inventory of an ADS separator tank may be calculated by differencing two consecutive calculated values of the liquid and vapor masses: dhi,33,,x ,, , AM,o3 4.x 3y dt At hi ADS 4-X SP. f. i -M Aos 4.x S&. r. '-'
, 4.7-8 t, - t,,,
. Mins ,_x sy, , ,, - MAos 4 x Sm. s. i.i
(, - t,,, where the subscript: 1 = Index of data and time arrays 4.7.2 Steam Flow Rates The density of steam in the exhaust line from the ADS-4 separator is evaluated using the ASME steam O table routines with pressure and temperature inputs from the data channels identified in Table 4.7-2. Using this density, the steam portion of flow through the ADS-4 valves is calculated as: NI ADS 4-X. 3
" C kg WN -XXX 4*7~9 where:
W = Volumetric flow rate of steam, ft.8/ min. C = Conversion constant, minutes to seconds and the subscript: FVM-XXX = Instrument channel ID for steam vapor flow meter XXX = 603 for ADS 4-1 separator XXX = 602 for ADS 4-2 separator 9; m:\apMX4ec-4\2344w 47.nos:1b-100395 4,7 4 REVISION: 1 l
A
. () 4.73 Liquid Flow Rates Liquid from ADS-4 is directed from the separator through a magnetic flow meter into the primary sump tank. Mass flow from the ADS-4 separator to the primary sump is calculated as:
AADS 4-X. f 'C 1 k ADS 4-X. f W FMM-xxx 4.7-10 where: W = Volumetric flow rate ofliquid, gpm Ci = Conversion constant, gpm to ft.8/sec. and the subscript: FMM XXX =- Instmment channel ID for liquid flow meter XXX = 603 for ADS 4-1 separator XXX = 602 for ADS 4-2 separator
'The density of the liquid passing through the flow meter is determined using data from the pressure i 7g and temperature instruments identified in Table 4.7-2.
Minimum Flow Criteria: l Until the measured level from LDP-502 or LDP-507 is less than an input threshold value, all mass i flow rates are set to zero. 4.7.4 Total Flow Rate Total liquid and vapor flow through ADS-4 are then calculated as:
^D8d-X ' 4.7-11 Mins ,,x, , = Myyy.xxx +
^DS d-X 8 4.7-12 N1,os,,x,, = K1,yy,xxx +
O J
'mpe4u)44w-47.nos:tb-too395 4.7 5 REVISION: 1
Total flow through one ADS-4 flow path is then calculated as: b ADS 4-X
- AD5 4-X. f + ADS 4 X. g 4.7-13 he total ADS-4 flow rate is calculated by summing the total flow rate for each of the two flow paths.
The tatal mass passed by ADS-4 is calculated by integrating step-wise over time: ( dt 4*7-l4 MADS 4. TOTAL " ADS 4-1. TUTAL
- ADS 4-2. TUTAL )
he flow que.lity of the each of the ADS-4 flow paths is also calculated as:
= ADS 4-X. s 4.7 15 X ,g3 ,_x ADS 4-X. TUrAl.
Calculation of Local Flow Qualities: Flow qualities are calculated at several locations to provide results for comparison against code O predictions for modeled system boundaries. His is accomplished by calculating the flow quality at a separator for the ADS flows and then calculating the quality at the point of interest in the test facility by assuming the expansion process associated with the exhaust line is isentropic (constant entropy). He entropy associated with the flow to the separator is: S3 , = (1 - X 3,) S,,3, + X 3,S,,3, 4.7 16 he entropy at the location of interest is calculated as a function of the local temperature and pressure at that location. Ss .Loca = S (P T,) 4.7-17 Sr, toca = S (P, T,) 4.7-18 9 mnpoomsee-4u 344 w.47.non : 1 b- i oo395 4,7 6 REVISION: 1
i i fh h he entropy of vaporization is calculated as: Sr , mt = S,, toeg - Sr. mL 4.7-19 De local flow quality at the location of interest is calculated as:
= S"~S "'
8 8-X' C^' S rs 4.7 20 The calculation for Xcoca is performed at four locations: ADS 1-3 separator, ADS 4-1 separator, ADS 4-2 separator, and the break separator. De instruments used are identified in Table 4.7-3. l 1 4.7.5 Energy Balance ' Energy flow through the ADS-4 separator consists of the following:
- Rate of change in stored energy of the ADS-4 separator fluid inventory ,
. Energy transport rate from the ADS-4 separator by exiting steam flow e' Energy transport rate from the ADS-4 separator by exiting liquid flow g = Rate of change in stored energy of the ADS-4 metal components ]
- Rate of energy loss from the ADS-4 components to the environment )
The expressions for evaluating these five energy transfer or transport terms are developed in the i following sections. I l i m%wone-4u144 47..on:id-loo 395 4.7 7 REVISION: 1
4,7.5.1 Rate of Change in Stored Energy of the Automatic Depressurization System-4 Separator Fluid Inventory l The rate of change of energy associated with the fluid (both steam and liquid) inventory of a ADS-4 separator may be expressed as: d ( htAos 4-x ss, T, ) O ADS 4-X SEP U.t P g
= cP. t (I t~ IREF} g(
4.7-21 d(T,)
+ cp, , M^ ' ' '* ' '"* '
dt d ( hi d ( T, )
+ c,, , (T, - Tag,) gos ,.x 3ep, , ) + c,, gos, *
hi '8"' ' dt Expressing the previous equation as a difference: c,, ^D8 4 -* '" ' *""' e = cp, , (T, - T,,,) dt At
+ c,,
p hi go, ,.x 3ep, , AT' 4.M Ah! * ' AT8
+ c, , (T, - Tasp) + c,,, higos ,,x 3gy, ,
( O m:pc.4us44.-47.ntw:ib.ioo395 4.7-8 REVISION: 1
73 V Expanding the terms on the right-hand side of the equation: A S 4.x SEr, f cp, , (T, - Tagp) At MADS 4-X sEP, f, i - M^ 8 d -* 8 5"' '
=c p (Tc , - T,) 4.7-23 t, - t,,3 c p,, MAos 4-x sse, f
' AT'
= c,,Mm p 4-x sse.f,i T ' - T ' '
At _ cp , , (T, - T,,p) **
= c, , (T,, , - Tasp) s
[*"*5"
-l 4.7- M /^%
(v) AT' T'-T S SN c,,, Maos 4.x ser, s g " Cp s M,03 ,_x ,gg, ,, , i ,, 6-1 where: f
- i = Index of data and time arrays The specific heat of the liquid, cp,,, is evaluated using the pressures and temperatures measured using the instruments identified in Table 4.7-2. The total change in energy associated with the change in
- inventory of the two ADS-4 separators is then calculated as
Q Aos 4 QAos 4.i + Q ns A 4-2 4.7-25 i l t
% 1-
. t i m: Wee-au344w-47.aon:lb 100395 4,7 9 REVISION: 1 I l 1
}
4.7.5.2 Energy Transport Rate from the Automatic Depressurization System-4 Separator by Exiting Steam Flow The enthalpy of the steam exhaust from the ADS-4 separator is evaluated using the output from the pressure and temperature sensors associated with vapor flow meters (Table 4.7-2): h,,pyy.xxx = h,(PT-XXX, TF-XXX) 4.7-26 FVM-XXX is the data channel ID for the ADS-4 vapor flow meter, and PT-XXX and TF-XXX denote the vapor pressure and temperatures, respectively, associated with that flow meter. The rate of energy transport of the ADS-4 flow due to the steam component, then,is expressed as: O ADS 4.x.g " NI ADS 4-X. g 4.7-27 h,,n,y_xxx 1he energy transport associated with steam flow from the two ADS-4 separators is then calculated as: O ADS 4. g " O ADS 4-1. g +O ADS 4-2. g
- O 4.7.5.3 Energy Transport Rate from the Automatic Depressurization System-4 Separator by Exiting Liquid Flow The enthalpy of the ADS liquid flow is determined using the output from the pressure and temperature sensors associated with the liquid flow meters as given in Table 4.7-2:
l h r . ruu-xxx
= h, (PT-XXX, TF-XXX) 4.7-29 FMM-XXX is the data channel ID of the ADS-4 liquid flow meter ofinterest, and PT-XXX and TF-XXX denote the data channel ids for the liquid pressure and temperatures, respectively, associated with that flow meter. The rate of energy transport due to liquid flow from the ADS-4 separator, then, is expressed as:
O ADS 4-x. f
- h,,pyy.xxx 4.7-30 ADS 4-x. f e
mAap600ssec-4u344.-47.non:lb 100395 4.7-10 REVISION: 1
O l h he total energy transport associated with liquid flow from the two ADS-4 separators is then calculated as: O ADS 4. f ' 4.7-31 OADS 4-1. f + O ADS 4-2. f 4.7.5.4 Rate of Change in Stored Energy of the Automatic Depressi.rization System-4 Metal Components l J
'The ADS system is heat-traced and, therefore, the change in stored energy of the ADS metal l components due to energy loss from the fluid is negligible.
4.7.5.5 Rate of Energy Loss from the Automatic Depressurization System-4 Components to the Environment he rate of energy loss from the ADS-4 separator and its associatM piping to me environment is taken to be zero since the separator and steam exhaust line are heat-traced. Thu , the separator and steam lines are treated as an adiabatic boundary. p.s 4.7.5.6 Fluid Stored Energy For use in the overall system energy balance calculations, the fluid stored energy in each of the two ADS-4 separators is given by the following: U ,o,, = U,,,o3, + U,, 333 , 4.7-32 where: U,, 333 , = M,, ,o3 , x c,, ,, ,os , x U,, 333 , - T,,) 4.7 33 and: U, gos , = M, 333, x cp,,, 333, x U,, 333, - T,,,) 4.7-34 , O v maapmsee-4u)44 -47.non:tb-100395 4.7-11 REVISION: 1-
i l l TABLE 4.71 9 INSTRUMFNTATION TO BE USED FOR ADS 4 SEPARATOR LEVELS INSTRUMENT CORRECTION Level Pressure Fluid Location Function Transducer Transducer Temperature ADS 41 Density Separator compensation of LDP-611 l'T-611 TF-619 levels data ADS 4-2 Density Separator compensation of LDP-612 PT-610 TF-618 levels data TABLE 4.7-2 ADS 4 SEPARATOR STEAM AND LIQUID PRESSURE AND TEMPERATURE INSTRUMENT CIIANNELS Flow Meter Flow Meter Pressure Temperature Description Channel ID Channel ID Channel ID ADS Separator 4-1 steam flow FVM-603 FT-611 TF-623 ADS Separator 4-1 liquid flow FMM-603 FT-611 TF-619 ADS Separator 4-2 steam flow FVM-602 FT-610 TF-622 ADS Separator 4-2 liquid flow FMM-602 FT-610 TF-618 9 mwwasec-4u144w-47.non:tb-t oo395 4.7-12 REVISION: 1
\ / TAllLE 4.7 3 INSTRUMENTS TO DE USED IN CALCULATION OF LOCAL FLOW QUALITIES Separator Local Primary Loop Location Location Pressure Temperature Pressure Temperature ADS 13 Separator =o Pressuruer l'r-605 TF-616 (liquid) I'T-604 TF-602 TF-617 (vapor) (liquid and vapor) ADS 4-1 Separator => IIL-1 FT-611 TF-619 (liquid) Irr-205 TF-609 TF-623 (vapor) (liquid and vapor) ADS 4-2 Separator => HL-2 l'r-610 TF-618 (liquid) Irr-202 TF-610 TF-622 (vapor) (liquid and vapor) Break Separator => Break Irr-905 TF-912 (liquid) Define as a Namelist in input; Location TF-613 (vapor) Dependant upon Break Location V j l 1 f3 V mNewec 4u)44w-47.non;1b-too395 4,7-13 REVISION: } l
TABLE 4.7 4 O VOLUME VERSUS IIEIGIIT FOR ADS 41 FLUID VOLUME CALCULATIONS Ileight (in.) Volume (in.2) 0 0 2.0 54.9 27.75 234.4 40.25 1792.0 62.25 8076.0 89.75 16,171.0 100.75 17,232.0 TABLE 4.7 5 VOLUME VERSUS IIEIGHT FOR ADS 4 2 FLUID VOLUME CALCULATIONS Height On.) Volume (in.') 0 0 27.75 96.0 41.25 1672.0 62.5 7843.0 90.75 16,183.0 101.25 17,016.0 O m:\np60 wee-4\2344w.47.no :16.too395 4.7-14 REVISION: 1
i 4.8 Break Separator
'Ihe purpose of the break separator is to separate break flow into liquid and vapor components and t i ;
measure the flow rates of the single-phase flow components. Once measured, vapor flow is exhaured to the ambient environment, and liquid is directed to the primary sump simulation of the test facility. The break separator consists of a tank (separator); a vortex (vapor) flow meter; a magnetic (liquid) i flow meter; and associated valves, piping, and instrumentation.
'Ihe total break flow may be calculated as:
dM 8RK82
$1,as,g = Manxuo + baaxsm + 4.8-1 where the subscripts:
BREAK = Total break flow BRK LIQ = Liquid flow from break separator BRK SEP = Break separator tank BRK STM = Steam exhaust from break separator l O To address possible reverse flow from the sump to the break separator, a condition of reverse flow is l t determined by evaluating the level from LDP-901. If LDP-9012: 75.275 in., then-l dM S. tsam ssp " ~
- f. ADS 4-1 SEP dt 1
l 4.8-2
+
- M .Aos4-2 r sEP ~OVREW SUMP INJ where:
Nf oyan, = Liquid mass flow rate through IRWST overflow line measured by FMM-703 l This new value for M,,,,x 3g, is then used in Equation 4.9-1 of Section 4.9, h calculate the general ! mass balance for the sumps. j fm b I maap600Q344w-48.noo:Ib too395 4.8 1 REVISION: 1
Energy is transported out of the primary system by both the vapor and liquid components of break flow. Additionally, the following occurrences cause the stored energy level of the break separator to vary: a change in the fluid inventory held within the separator; a change in temperature of the fluid inventory held within the separator; a change in the temperature of the metal mass of the separator tank or a loss of energy to ambient. Rus, the energy balance for the break separator, accounting for the change in stored energy of both the liquid inventory of the separator and the separator tank, may be expressed as: O BREAK B R K LlQ + BRK$1N + CP +Q BRK SEP METAL # BRK SEP AMB where the subscripts: METAL = Metal mass of break separator tank and associated piping AMil = Energy loss to ambient environment 4.8.1 lireak Separator Liquid Inventory he break separator is, in its simplest form, a tank. Liquid inventory in the break separator is monitored by a level transducer. For all tests, with the exception of doubleended DVI line breaks, an orifice was in place within the span of the level transducer. Herefore, a correction was made to the level indication to adjust for the pressure drop through the orifice. The functional steps and associated system of equations for operating on the output from the break separator level transducers to calculate liquid mass in the sump tanks follows: Step 1: Calculate the pressure drop through the orifice and add this value to the resdings from the breat separator level transducer listed in Table 4.81 to account for the pressure drop through the orifice. The instruments used to measure level, local pressure, ar.J fluid temperature are identified in Table 4.8-1. He flow meters used to calculate the pressure drop are identified in Table 4.8-2. Step 2: Compensate the adjusted level indication for the break separator level to account for temperature differences between fluid in the separator tank and fluid in the rLference leg of the instrument line. The local pressure and fluid temperatures used to accomplish the compensation are also identifled in Table 4.8-1. O mwwa.o :th-toows 4,g.2 REVISION: 1
Step 3: The local pressures and temperatures identified in Table 4.8-2 are used to calculate the density of the fluid in the break separator: ' Pt. sax ser
- Pr (PT-905, TF-912)
~
4.8-4 p,,,,,3ep = p,(PT-905, TF-9i 3) .i where: FT-905 = Local pressure in break separator tank TF-912 = Local liquid temperature of fluid in break separator tank TF-913 = Local vapor temperature of fluid at break separator exhaust and the subscripts: f = Liquid phase of water g = Vapor phase of water There is only one liquid temperature measurement instrument for the break separator, located in the drain line to the primary sump just downstream of flow meter FMM-905 'Ihis temperature measurement is used to evaluate the thermodynamic properties ofliquid in the break separators. Step 4: Using the compensated liquid level and the break separator volume as a function of height, determine the volume of liquid in the break separator as: V,, ,,g 3g, = M,,,,,, x DMMcoup 4.8-5 where:
- V = Volume, ft.'
V(1) = Tank (steam / water separator) volume as a function of elevation, ft.8/ft. i (See Table 4.8-3) LDP-905 coup = Compensated fluid levels data from level transducer LDP-905, ft. i Step 5: The liquid mass inventory in the break separator is now calculated as: M,, ,,g sep = p,, 3,x ,,, x V,, ,,x 3g, 4.8-6 b( mA p6000w4s. on:;e :00395 4.8-3 REVISION: 1
The vapor mass inventory of the break separator is then calculated as: 1 Ms. sax ser " P, x (Veux see, wr4t. - V,, ,,, 3,) 4. M I where the subscript: TOTAL = Total volume associated with the break separator, ft.' Step 6: The rate of change in mass inventory of the break separator tank may be calculated by differencing two consecutive calculated values of the liquid and vapor masses: dMann ser , Manx see , M,, ,,, 3,p, , - M,, , gg 3,p, ,,, M ,, ,,g 3,p,, - M ,, ,,g 3,p,,.
. 8-8 dt At t, - t,,, t,-t,i where the subscript:
i = Index of data and time arrays 4.8.2 Steam Flow Rates l The density of steam in the exhaust line from the break separator is evaluated using the pressure and l temperature values recorded from the data channels identified in Table 4.8-2. Using this density, the steam portion of the break flow vented to ambient is calculated as: Ni m.m = C x p, x W m_m 4.8-9 where: W = Volumetric flow rate of steam, ft.'/ min. C = Conversion constant, minutes to seconds and the subscript: FVM-XXX = Instrument channel ID where: XXX = 905 (6-in. line) 906 (8-in. common exhaust line) 9. mMp60lA2344w-48.non:Ib-100395 4,g.4 REVISION: 1 ; I l
l . Total break steam flow is calculated as: d Ms. SRx SEP 4.8-10 s,, ,,, = Mrvum5 + Mrvum6 + 4.83 Liquid Flow Rates . A single line directs liquid flow from the break separator into the primary sump tank.- The mass flow from the break separator to the primary sump tank is calculated as: Mruums " Ci x p, x W ruums 4*8'Il where: W = Volumetric flow rate of liquid, gpm Ci = Conversion constant, gpm to ft.8/sec. and the subscript: O FMM-905 = Instrument channel ID for liquid flow meter in line between break separator and primary sump The density of the liquid passing through the flow meters is determined using data from the pressure and temperature instruments identified in Table 4.8 2. Total liquid break flow rate passed to the sump simu!ation is calculated as: s,, ,,g = M ruum5 + d Mr8RK sEP 4.8-12 4.8.4 Total Flow Rate l I The total break flow rate is then calculated as: 4 Marx. TUTAL " f.BRK + g,BRX
- The flow quality of the break flow is also et . ilater' as:
X ,,, = k* *"" 4.8-14 N sRx. TOTAL 1 m:Wm-48.noa:tb-too395_ 4.g.5 REVISION: 1
4.8.5 Energy Balance The energy flow through the break separator consists of the following:
- Rate of change in stored energy of the break separator fluid inventory
. Energy transport rate from the break separator by steam exhaust flow
- Energy transport rate from the break separator by liquid flow to the sump
- Rate of change in stored energy of the metal of the break separator tank
- Rate of energy loss to the environment
'Ihe expressions evaluating these energy transfer or transport terms are developed in the following subsections: 4.8.5.1 Rate of Change in Stored Energy of ;he Break Separator Fluid Inventory The rate of change of energy in the break separator fluid may be expressed as: d(M, AT,) d(T,) Qgggsop=c p = cp , , F, - Tagg) d(M, ,gg 3,p) p+ c , , M,, ,"* 8'" dt where: AT, = T,- Tag, Taor = Reference temperature,32*F T, = Measured liquid temperature, 'F Expressing the previous equation as a difference: 4.8-16 d( M,,,,gsep T, ) " AM .t saK ser Mr r C.r P, f f ~ REF +C,iP f. BRK SEP g dt At Expanding the terms on the right-hand side of the equation: c,, , (T, - Tarr)
" Cr. , (T, - Tagg)
*** ~ "*H at t, - t,,,
4.8-17 cp,, M ,, ,,g sep AT'
= c p,, M ,, ,gg3sp,,
T' ' - T' " At t, - t, ,, m.W3m-48.noo: b ioo395 4,g.6 REVISION: 1
1 1
.( where:
l 1- = Index of data and time array l l The specific heat of the ligid, c ,,,e is evaluated using the pressuies and temperatures identified in Table 4.8 i. The rate of change of energy in the steam in the break separator may be expressed as:
- g. BRx sEp T, ) 8' 8" 8" : 4.8-18 c.$
P gg
= c,3P (T g- TgEF) Tg g +c,gP M . BRK g SEP g Expressing the previous equation as a difference:
- g. BRK SEP g g. BRK SEP g g 4,8.} 9 C.g P
gg
= c,3 P (T - TRE) T3 gg P. g g.BRK W gt Expanding the two terms on the right-hand side of the equation:
- s. K sE" g. BRK SEP, i g. BRK SEP, i-l
.( c,, , (T, - Tag,) T, = -Ty e i-i AT T.-T 5H U.g g. BRK SEP = ce,, M ,, ,,g ,g,,, 4.8-M P
g i , :-i The specific heat of the steam, c e,, , is evaluated using the pressures and temperatures measured by the instruments identified in Table 4.8-1. d i 4.8.5.2 Energy Transport Rate from the Break Separator by Steam Exhaust Flow The enthalpy of the steam exhaust from the break separator is calculated using the output from the pressure and temperature sensors associated with FVM-905 and FVM-906 (Table 4.8-2): h,, m xxx = h,(PT-YYY, TF-ZZZ) 4.8-21 m:W344w-48.non:lb-100395 4,8 7 REVISION: 1
i FVM-XXX denotes a specific flow meter, and I'T-YYY and TF-2ZZ denote the vapor pressure and temperatures associated with that flow meter. De rate of energy transport from the break separator due to exhaust steam, then, is expressed as: 4.8-22 Opvu.m = Ni ryu.m x h,, evu_m where: Qrvu m = Rate of energy transport due to steam flow through flow meter FVM-XXX The total energy transport rate, then, is calculated as: Osax. :
- Orvues + Orvu-w, 4.8-23 4.8.5.3 Energy Transport Rate from the lireak Separator by Liquid Flow to the Sump The enthalpy of liquid flow into the break separator is calculated using the output from the pressure and temperature sensors associated with FMM-905 (Table 4.8-2):
hr . ruums = h, (Irr-905, TF-912) 4.8-24 FMM-905 denotes the flow meter to which the enthalpy is applicable, and e >05 and TF-912 denote the liquid pressure and temperatures associated with that flow meter. The rate of energy transpon from the break separator due to liquid overflow into the sump, then, is expressed as: Osax,t " Nisnx.,x h,,puu y, 4.8-25 4.8.5.4 Rate of Change in Stored Energy of the Metal of the Ilreak Separator Tank De break separator tank and associated piping is heat-traced. Therefore the change of stored energy of the metal components due to energy loss from the fluid is negligible. 4.8.5.5 Rate of Energy Loss to the Environment The rate of energy loss from the break separator and its associated piping to the environment is zero because the separator and the steam exhaust lines are heat-traced. Thus, the separator and steam lines are treated as an adiabatic boundary. Ileat loss to the environment from the liquid drain line running from the break separator to the sump is neglected since the run of pipe is small. O mMp60003h-48.non%l00395 4.8-8 REVtSION: 1
O. \ U TABLE 4.8-1 1 INSTRUMENTATION TO BE USED FOR BREAK SEPARATOR MASS AND ENERGY BALANCE 1 s Level Pressure ' Location Function Transducer Transducer Fluid Temperature Break separator Density compensation LDP-905 fvr-905 TF-912 of levels data TABLE 4.8-2 BREAK SEPARATOR STEAAf EXHAUST AND LIQUID PRESSURE AND TEMPERATURE INSTRUMENT CIIANNELS Flow Meter Description Flow Meter Channel ID Pressure Channel ID Temperature Channel ID Break separator FVM-905 frr-905 TF-915 steam exhaust flow FVM-906 0.5 x (PT-902 + frr-905) TF-918 Break separator FMM-905 PT-905 TF-912 gi t liquid flow
%}
I i l l l l l l l l l l l
\
v
) 1 l
l mAap60m2344w48.non:ltr100395 4,g.9 REVISION: 1 1
TABLE 4.8-3 VOLUME VERSUS IIEIGIIT FOR BREAK SEPARATOR FLUID VOLUME CALCULATIONS IIelght Volume (in.) (in.') 0.0 607.4 39.5 1243.5 55.5 6648.1 81.0 26447.0 82.25 28204.6 88.5 32338.0 106.25 46324.7 136.25 69545.3 145.75 76835.0 151.75 81884.1 181.75 104933.0 196.25 107575.2 O m:W344w-48.non:ib-loo 395 4.8-10 REVISION: 1
>I l
-)
.j i I 4.9 Sumps
. . . l
. In the AP600, the sump collects all liquid released from the primary system and serves as the source {
. of post-accident LTC water inventory. In the OSU test facility, the sump is modeled by two tanks: a primary sump tank and a smaller secondary sump tank. These tanks have associated with them piping, vapor and liquid flow meters, and other pressure and temperature measurement instnamentation.
w Accounting for all possible flow paths associated with the sump, the general mass balance on that component may be expressed as: dM BRK SEP ADS 4-1 SEP ADS 4-2 SEP IRWST dt 4,9,1 >
-b sTu xusT -MSuwPim where the subscripts are defined as:
SUMP = Both primary and secondary sump tanks ! BRK SEP = Liquid flow from break separator through FMM-905 ) V ADS 4-1 SEP = Liquid flow from the ADS 4-1 steam / water separator through FMM-603 ADS 4-2 SEP = Liquid flow from the ADS 4-2 steam / water separator through FMM-602 IRWST = Liquid overflow from IRWST into sump through FMM-703 j STM XHST ' = Steam exhaust from sump through FVM-903 l SUMP INJ = Liquid flow from sump to DVI line through FMM-901 and FMM-902 l Energy is transported into and out of the sump by both the vapor and liquid components of flow associated with the sump. In addition, the stored energy associated with the sump may change due to ,
- a change in the liquid inventory held within the sump, a change in temperature of the liquid inventory I held within the sump, or a change in the temperature of the metal mass of the sump tanks, or heat loss l to the ambient. Thus, the energy balance for the sump, accounting for the change in stored energy of both the liquid inventory of the sump and the sump tank (s) may be expressed as:
4 i d( M T) '
- P BRK SEP ADS 4-1 SEP ADS 4-2 SEP IRWST
- Osru xust - OsuMP rsi - GUMP METAL
- QsUMP AMB I where the subscripts are defined as:
4 METAL = Metal mass of sump tanks and associated piping (. AMB = Energy loss to ambient environment I m:\.ptewu4u344.-49.nos: n>. t oo395 4,9 1 REVISION: 1
4.9.1 Sump Liquid Inventory For convenience, liquid mass in the primary and secondary sump tanks are calculated separately, then summed to yield the total liquid mass in the sump. Liquid inventory in the sump tanks are measured by a level transducer. Load cells installed on the two sump tanks are not used for inventory calculations. After reviewing the primary- and secondary-sump fluid temperatures for all tests (using thermocouples in Table 4.9-1, and including TF-907 which is located at the top of the primary-sump tank), it was determined that the sump temperatures were always less than the saturation temperature evaluated at the total pressure. It was concluded that 1) the water region clearly remained subcooled in the sump tanks, and 2) it was not possible for a steam region to be supported above the water level in the sump tanks (as occurs in some internal system components, such as the CMTs). He region above the water level is a cool mixture of air and some steam, and an accurate calculation of its composition is not possible since the partial pressure of the steam in the mixture is unknown. Berefore, steam inventory is not modeled in the sump tanks and the region above the water level is not included in sump inventory. 4.9.1.1 Use of Level Measurement for Mass Calculation
%e functional steps and associated system of equations for operating on the output of the sump level transducers to calculate liquid mass in the sump tanks follows:
Step 1: Compensate all readings from the primary and secondary sump level transducers to account for temperature differences between fluid in the tanks and fluid in the reference legs of the instrument lines. He two channels of level data to be compensated are identified in Table 4.9-1. The instruments used to measure local pressure and fluid temperatures to be used to accomplish the compensation are also identified in Table 4.9-1. Step 2: De local pressures and temperatures as recorded from the instrument channels identified in Table 4.9-1 are used to calculate the density of the liquid in the sump tanks: P r 3 " P r. j(P, T) 4.9-3 where the subscripts are defined as: f = Liquid phase of water j = Either sump tank maap600\aec4Q344w-49.noo:Ib-100395 4.9 2 REVISION: 1
1 (O - The instrument ids and the elevation of the instruments for the local fluid temperature measurements to be used in the calculation of fluid densities for both the primary and secondary sumps are given in Table 4.91. Step 3: Using the compensated liquid level and the sump volume as a function of height, determine the volume ofliquid in the sump:
=
V,, , V(1)suur, x LDP-XXXcoup 4,9 4 where: s V = Volume, in.8 V(1) = Volume as a function of elevation,in.8/in. LDP-XXXcoup = Compensated fluid levels data from the level transducer (in.) identified as XXX = 901 (primary sump LDP transducer) XXX = 902 (secondary sump LDP transducer) Step 4: The liquid mass inventory in the primary and secondary sump is calculated as: y l
=
M,,syyp; p ,, , x V,, , x C 4.9-5 ) Q where: C = Conversion constant, in.8 to ft.8 I The total liquid mass inventory in the sump is calculated as: ; i M r. suur = M,, syy, p + M .r suur s 4.9-6 i l where the subscripts: P = Primary sump S = Secondary sump Step 5: The rate of change in mass inventory of the sump tanks may be calculated by differencing two consecutive calculated values of the liquid masses: dMsuur , AM3yy, , M r. suur. i - M .r suur, i-i , dt At t, - t,., 4,9 7 O mwamu4uwa9.noo: b.mn95 4,9 3 REVISION: 1
i where the subscript: i = Index of data and time arrays 4.9.2 Sump Steam Exhaust Flow Although steam inventory is not modeled in the sump tanks (Subsection 4.9.1), steam exhaust flow can be calculated using a vapor flow meter on either of the two exhaust lines (3/4-in. schedule 40 piping containing FVM-903, and the common 8-in. header containing FVM-906). The meters measure flow in standard cubic feet per minute, scfm. Positive flow was indicated by these two vapor flow meters at various times during the tests. Ilowever, after checking the readings of the exhaust line fluid thermocouples (Table 4.9-2), the temperaturcs were usually less than their corresponding saturation temperatures (evaluated at the total ' pressures), although some were above their saturation temperatures. As discussed in Subsection 4.9.1, when vapor-space temperatures are less than saturation, an air / steam mixture may exist and there is no way of accurately measuring how much steam is present. Therefore, steam exhaust flow modeling was limited to positive vapor flow meter readings in conjunction with fluid thermocouple readings indicating temperatures at or above the saturation temperature. 4.9.2.1 Steam Flow Rates The density of steam in the 3/4-in. and 8-in. common header exhaust lines is evaluated using ttx: ASME steam table routines with pressure and temperature inputs from the data channels identified in Table 4.9-2. Using these densities, the steam mass flow is calculated as: M m -xxx = C, x p,, m.xxx xWm -xxx 4.9-8 where: W = Volumetric flow rate of steam, ft.8/ min. Ci = Conversion constant, minutes to seconds and the subscript: FVM-XXX = Instrument channel ID where: XXX = 903 (3/4-in exhaust line) XXX = 906 (8-in. common exhaust line) O m Aap600\sec4\2344 w-49.non: l tw l 00395 4.9 4 REVISION: 1
r A g The total steam mass flow rate exhausted by the primary sump and the break separator is calculated as: Msm xusr " bevu-m + M rvu-m 4 9-9 The total steam mass flow is calculated by integrating the mass flow rate step-wise over time: Msm xust "1 ($sm xust x At) 4.9-10 4.9.3 Sump Injection Two lines provide for liquid to flow from the primary sump into the DVI lines and the reactor
- pressure vessel simulation. Mass flow from the sump through either of the two sump injection lines is calculated as:
Mpyu.xxx = C2 x ppyy.xxx x W ,yy.xxx 4,9 11 where: W = Volumetric flow rate of liquid, gpm C2 = Conversion constant gpm to ft.'/sec. and the subscript: FMM-XXX = Instrument channel ID where: XXX = 901 (piping run to DVI-1) XXX = 902 (piping run to DVI-2) The density of the liquid passing through the flow meters is determined using data from the pressure and temperature instruments identified in Table 4.9-2 as input to the ASME steam tables. The total mass flow rate to the DVI lines from the sump is calculated as: M seurim "bruumi + S mum 2 4.9-12 The total mass injected by each injection line into the DVI line is calculated by integrating the product of the measured mass flow rates and the time interval over which the measurement is taken: l My y.xxx ={(Mm u-xxxa x At ) 3 4.9 13 5 m:Wec44344w-49.non:Ib 100395 4,9 5 REVISION: 1
The total liquid mass injected from the sump is calculated as: hi suur m,
= M yyy ,,, +M euu-902 4 9-14 4.9.4 Total Flow Rate Out of the Sump Total mass flow rate out of the sump is then calculated as:
Msuur " b scur m) +M sru xusr 4.9-15 Total mass flow out of the sump is calculated by integrating the mass flow rate step-wise over time: Msuur.Torrt " E Msuup x At 4.9-16 4.9.5 Energy Italance Energy into the smnp from the break separator, the two ADS-4 separators, and the IRWST overflow lines are calculated in and obtained from their respective modules. Thus, the calculation of an energy balance on the sump requires that the following parameters be evaluated:
- Rate of change in stored energy of the sump liquid inventory
- Energy removal rate from the sump by steam exhaust flow
- Energy removal rate from the sump by injection flow supplied to the DVI lines
- Rate of change in stored energy of the metal of the sump tanks
- Rate of energy loss to the environment The expressions for evaluating these five energy transfer or transport terms are developed in the following sections.
4.9.5.1 Rate of Change in Stered Energy of the Sump I,iquid inventory The rate of change of energy in the liquid in the sump may be expressed as: d( hi.suwe r T, ) d( M .r sump r
+C.r dt P' ' "8' dt r t.suur dt 4.9-17 Expressing the previous equation as a difference:
d( M.r suur T, ) " M .suur r Mr C.' r dt P' r ~ REF g
*C.r P r. SUMP g 4,g,} g mwwwec4u344w-49.no :ib.too395 496 REVIs!ON: 1
d i l i Expanding the two terms on the right-hand side of the equation: 8" "E ' ~ "E" cp, , (T, - Taoy) = c,, p U,,, - Tagg) At t, - t, , 4.9-19 l l AT T -T I c,,hi.suur p t = c ,p, M . rsuue, '
-t 49a 4
where: 4 d i = Index of data and time arrays
- l
! Each of the terms on the right-hand side of the previous equation is to be evaluated for both the 1 primary and secondary sump tank inventories. The specific heat of the liquid, c r.r , is evaluated using :
- the pressures and temperatures identified in Table 4.9-1. Note that, for the primary sump, the specific
{ i heat of the liquid may be evaluated for more than one zone or region, depending on the water level in the sump. 4.9.5.2 Energy Removal Rate from the Sump by Steam Exhaust Flow (
- d The enthalpy of the steam exhaust from the sump is calculated using the output from the pressure and temperature sensors associated with FVM-903 and FVM-906 (Table 4.9-2)
I
=
hs . rvu-m h,(PT-YYY, TF-ZZZ) 4.9-21 ; 1 FVM-XXX denotes a specific flow meter, and FT-YYY and TF-ZZZ denote the vapor pressure and temperatures associated with that flow meter. The rate of energy transport from the sump due to exhaust steam, then, is expressed as: 1 i Orvu-m " NI rvu-m x h s. rvu-m 4.9-22 ' V . The total energy transport rate, then, is ca.culated as: Osm xasT Omos + Orvu-906 4.9-23 l ., O a Q/ r mAap60tNec4\2344w 49.non:th-100395 REVISION: 1 4.9-7
l l l l 4.9.5.3 Energy Removal Rate from the Sump by Injection Flow Supplied to the Direct Vessel Injection Lines The enthalpy of the injected liquid from the sump is determined using the output from the pressure and temperature sensors associated with FMM-901 and FMM-902 (Table 4.9-2) a xi the ASME steam table routines. Expressed rnathematically: ht puu_xxx
=
h, (PT-YYY, TF-ZZZ) 4.9-24 FMM XXX denotes a specific flow meter, and FT-YYY and TF-ZZZ denote the liquid pressure and temperatures associated with that flow meter. The rate of energy transport from the sump due to injected liquid, then, is expressed as: Qruu-xxx
= M g m .xxx x h.r ruu-xxx 4.9-25 The total energy transport rate, then, is calculated as:
Osuur isi Oruu-4oi + Oruu-m 4.9-26 4.9.5.4 Rate of Change in Stored Energy of the Metal of the Sump Tanks The change in stored energy of the sump tanks and associated piping may be expressed as: d( TSUMP METAL SUMP METAL CP . MmL M suur ustrAL 4,g,27 dt where the subscript: METAL = Metal of the primary and secondary tanks and all associated piping For each sump tank, the calculations parallel those for the cold-leg balance lines (see Subsection 4.4.6). Each sump tank is divided into a number of metal segments. Fluid thermocouples are employed to obtain pseudo-metal temperatures when necessary. Tables 4.9-3,4.9-4, and 4.9-5 list the data required for the calculations. O m Aap600\sec4\2344 w-49.non: I b- 100395 4.9-8 REVISION: 1
]
1 J4.9.5.5 Irate of Energy Loss to the Environment De rate of energy loss from the sump and its associated piping to the environment may be expressed as:
'I
.QSUMP AM8 = H x A x AT 4.9-28 )
I i where: A = Effective external surface area of the sump, ft.2 H= Overall effective heat transfer coefficient, Btu /(sec.-ft.2 *F) AT = 1 Difference between ambient air temperature and bulk metal temperature, 'F An equivalent heat transfer coefficient, accounting for heat resistance due to any insulation applied to (the outside surface of the sump volumes and natural convection from the outside surface to the
~ ambient, is used to evaluate the heat loss to ambient from the metal surfaces of the sump tanks. The same metal surface temperatures used to evaluate metal heat storage of the sumps is used to calculate heat loss to the ambient. Surface areas included in this calculation account for sump tanks and associated piping.
O. The metal-to-ambient heat transfer calculation methodology for the sumps is similar to methodology l
.V used for other insulated components of the system. Each sump tank metal segment is treated as an insulated vertical cylinder. The turbulent range form of the natural convection heat transfer correlation from Reference 16 applies, based upon the conditions of the OSU test facility air at atmospheric l pressure and ambient temperature (averaged over the tests). In the radiation heat transfer correlation, I an emissivity of 0.8 is assumed. l I
The detailed calculations and equations are identical to those for the metal-to-ambient heat transfer for l the cold-leg balance lines in Subsection 4.4.6, the only exception is that in the final expression for the natural convection heat transfer coefficient, the coefficient multiplier 0.21 is replaced with 0.09 and the exponent 1/4 is replaced with 1/3 (these differences are due to both the vertical orientation and the , turbulent range of applicability for the sumps). A 1 ). mhp600cc4u344w 49.mos:lb-100395 4,9 9 REVISION: 1 l
TAllLE 4.91 INSTRUMFNTATION TO IIE USED FOR SUMP MASS AND ENERGY IIALANCE Lesel Pressure Function Location Transducer Transducer Fluid Temperature TF-901 Density Pnmary sump LDP-901 FT-901 SC-903 compensation TF-905 of level data Secondary sump LDP-902 PT-901 SC-902 TAllLE 4.9-2 SUMP STEAM EXilAUST AND INJECTION PRESSURE AND TEMPERATURE INSTRUMENT CIIANNELS Flow Meter Flow Meter Description Channel ID Pressure Channel ID Temperature Channel ID Sump steam FVM-903 IT-901 TF-906 exhaust flow FVM-906 TF-918 0.5 (I'F-901 + I'T-905) Sump injection W M 901 N 901 TF-909 0* FMM-902 17T-901 TF-904 I O mhp60thec4uM4w-49.non:Ib 100395 4,9. l() REVISION: I
i
\
l 1
'd TABLE 4.9 3 DATA FOR SUMPS METAL ENERGY CALCULATIONS (PER SEGMENT)
Primary Sump Metal Segments Pseudo-Metal Outer Surface Mean Surface Temperature ID Metal Mass Area Area Metal Segment ('F) (Ibm) (ft.8) (ft.8) 1 TF-901 1907.7 84.65 82.25 2 TFM-901 1677.8 115.70 114.I8 3 TF-907 2132.4 84.65 82.25 Secondary Sump Metal Segments Pseudo-Metal Outer Surface Mean Surface Temperature ID Metal Mass Area Area Metal Segment ('F) Obm) (ft.') (ft.8) 1 SC-902 1096.6 43.43 41.71 2 TFM-902 2889.7 152.93 149.15 O L) l TABLE 4.9-4 DATA FOR SUMPS METAL ENERGY CALCULATIONS s Description Primary Sump Secondary Sump Ambient temperature ID TF-005 TF-005 (*F) Insulation thickness 1.5 1.5 On.) Insulation mean thermal 0.141 0.141 conductivity ((Btu-in.)/(hr. ft.' 'F)] l (D c) m:\ap60 msec 4W44w 49.non:lb-100395 4.9-11 REVISION: 1
TABLE 4.9 5 O\ SPECIFIC IIEAT CAPACITY VERSUS TEMPERATURE l TABLE FOR SUMPS METAL ENERGY CALCULATIONS l Metal c, Metal Temperature [Iltu/(lbm- F)] (*F) 0.1085 70. 0.1109 100.
)
0.1175 200. 0.1223 300. 0.1256 400. 0.1279 500. 0.1297 600. O 1 l l O l 1 m:\asuun 4u)44. 49.non:ib tco395 4,9 12 REVISION: 1
.[ \
V 4.10 Passive Residual Heat Removal The passive residual heat removal heat exchanger (PRHR HX) consists of a number of tubes with a fluid flow path at the top for the inlet connection from hot leg-2 (HL-2) and a fluid flow path at the bottom for the outlet connection to steam generator-2 (SG-2) channel head. The PRHR HX fluid mass, fluid energy, and tube metal energy conservation equations are described in Subsections 4.10.1, 4.10.2, and 4.10.3, respectively. 4.10.1 Fluid Mass Conservation Equation The general fluid (H O) 2 mass conservation equation, which relates the change in stored fluid mass with respect to time (the fluid mass time derivative) to the mass flow rates in and out, reduces to the following: dM gp,paga
" ~
dt " ' ' ' ""'"> """" "'('** "#'"""" 4.10-1 The left-hand side of the fluid mass conservation equation is approximated from the value of the fluid n mass at two consecutive time points (denoted by subscripts n-1 and n): dM ,o, g pay, M ,o, g pan, M g,o,,,ga,, - M ,0 g p,g,,, i dt At t, - t,,, The H 2O fluid mass is simply the sum of the water and steam masses: H3 0,PRHR f,PRHR g.PRHR 4 The water and steam mass calculations are based on the measured water level. The level (LDP) instrument channel ID for PRHR is listed in Table 4.10-1. This LDP is corrected for temperature effects using the LDP compensation method. As is generally assumed for each OSU test facility component, zero quality water is modeled below the compensated water level, and 100 percent quality steam is modeled above the compensated water level. This assumption is appropriate because the fluid below the water level is predominantly water, the fluid above the water level is predominantly steam, and any amount of frothing is indeterminate with the available instrumentation.
. To use the available measurements from the fluid thermocouples, the PRHR HX is divided into five axial fluid property zones for the calculation of various fluid conditions. Table 4.10-1 contains a list U
mecc4u344-4to.non:ib-too395 4.10-1 REVISION: 1
of the fluid thermocouples employed for each axial fluid property zone, along with the zone top elevations. Note that, since LDP-802 (which spans just the tubes) is being employed for the water level of the entire PRllR llX (including the inlet and outlet), instead of the usual method of using the axial fluid property zone temperatures / elevations in the LDP compensation calculations, it is n cessary to use an alternate set of temperature channels / elevations for the LDP corrections. Table 4.11-1 contains a list of the fluid thermocouples employed for each LDP compensation region, along with their elevations. The LDP compensation region boundaries are taken at the vertical midpoint between consecutive fluid thermocouples. In general, the water and steam masses are given by: N f,PRH R f.PRH R, Jet 4.10-4 N_ g.PRNR g.PRHR, en where: N,, paH, = Number of zones within the PRHR, (5) For zones j below those containing the compensated water level (which contain all water), the zone j water and steam masses are given by the following. where j = 1,...,levzone-1: hI r.eRHa' " Pr.rRHR, x Vpag,' x C 4.10-5 hi g.rRHR, " 0. where: 1 levrone = The zone containing the compensated water level l C = Conversion constant, in.2 to ft.8 l O m:\ap60&acc4\2344-410.non : l b- 100395 4.10-2 REVISION: 1
I l l
~
For the zone j containing'~the compensated water' level, the zone j water and steam masses are given by - the following, where j = levzone: f.PRHR[ b f,PRHR, f,PRHR, a g.PRHB, ", bg.PRHR, g.PRHR,f 4g. 4 For zones j above those containing the compensated water level (which contain all steam) the~ zone j water and steam masses are given by the following, where j = levzone+1, .,Nmpana: 4 f,PRHR, g.PRHR, " bg PRHR, PRHR, i i The fluid volumes are calculated :u a function of level from the volume-versus-height tabular data listed in Table 4.10-2, via linear interpolation within the table. No extrapolation at either c'id is performed; the first and last table points define the applicable range. The first point is [h,,, (= 0.), V., (= 0.)], and the last point is [h,,, V,.).'
- During initialization, the volume-versus-height data / function is emp,'oyed to calculate the total volume of each zone j as follows
for j = 1: Vpaga, = V(1) pag,, for j = 2,...,N_ pag,: 4.10-8 j-1 V pag ,, = W) pan,, - { V pag,, k =1 4 where: 4 V(1)p,g, = Volume as a function cf elevation,in.8/in., and 1 is the elevation of the top I of zone; i l , : O: , i
. m:\ap600\sec4\2344410.aos:th100305 4.10.3 REVISION: 1
l l ~ f During the transient calculations, the volume-versus-height data / function is employed to calculate the water volume of the zone containing the compensated water level, j = levzone, as follows: if j=levzone=l: V,,pagg, = W,) pag, 4.10-9 if j=levzone>l: s-n V,, pag, = W ,)p,g, -{Vpaga, i. where: V(1,) = Volume as a function of elevation,in.3/in., and 1r is the elevation of the compensated water level 'The steam volume of the zone containing the compensated water level, j = levzone, is then the following: 4.10-10 O V,.pyg g,=V pag , - V,,,,ga, The zone j water and steam densities are the reciprocal of the water- and steam-specific volumes, respectively: I br.PRHR' V.PRnR r ' 4.10-11 hg.PRHR g.PRHR, The zone j water-specific volume is calculated as a function of pressure and temperature from the ASME steam table function VCL as follows: v,,pagy, = VCL(Praga,T,,pagg) 4.1412 O m:\ap60 msec 4\2344-410 non:Ib 100395 4.10-4 REVISION: I
'(
The zone j steam-specific volume is calculated as a function of pressure and temperature from the ASME steam table function HS9 as follows (note that the steam-specific enthalpy is the main output - of HSS, and the steam-specific enucoy is also an output of HSS):
'li,,paga, = MWpaga, T,,paga,, S,,,,ga,, v,, pag,) ,
13
' The zone j water and steam temperatures are simply given by the zone temperatures, which are'the fluid thermocouple measurements TF-XXX (Table 4.10-1). Care is taken to ensure that the water temperature value employed is at or below the saturation temperature and that the steam temperature value employed is at or above the saturation temperature:
f.PRH R, PRE R,' sat PRHR l's.PRHR, = m W pan ,,, T,,,,,,g ,)
}4 i
v l where: i
.l Tpan% = E mpgn,> 4.10-15 i
'0- . O I 1he saturation temperature is calculated as a function of pressure from the ASME steam table function
- TSL as follows:
- T,,, pan, = TSL(Ppag,)
4.10-16 4
- The pressure is set equal to the value from the l'I' XXXpag, measurement (Table 4.10-1) after conversion from gauge (psig) to absolute (psia) pressure.
This completes the discussion of the calculations related to the left-hand side of the fluid mass conservation equation. f d b mAap600\sec4\2344-410.non: I b- 100395 4.10-5 REVISION: 1
4 On the right-hand side of the fluid mass conservation equation, the calculation of the outlet fluid mass flow rate (to the SG 2 outlet) is performed directly from the liquid volumetric flow rate measurement (FMM-XXXp,3,_,,o,,), due to the ist that the water flowing in the line remains subcooled (Table 4.10-1 for the instrument channel ID). Thus: out (to SC) H,0.PRHR out (to SG) f,PRHR 4.10-17
= max (0.,FMM-XXXpagt,,,,) x p ,,,,g,, x C, where:
Ci = Conversion constant, gpm to ft.3/sec. The integrated outlet water mass flow is also computed for use in the overall system mass balance. The water density is given by that from axial fluid property zone 1 (at the bottom). Although a hauld volumetric flow rate measurement (FMM) is available for the inlet line, it cannot be relied on for accurate mass flow rate calculations due to two-phase flow. Thus, having calculated both the change in stored fluid mass (from the left-hand side) and the outlet fluid mass flow rate (from the right-hand side), the fluid mass conservation equation is rearranged to solve for (infer) the inlet fluid mass flow rate (from the IIL-2). His completes the discussion of the calculations related to the right-hand side of the fluid mass conservation equation. 4.10.2 Fluid Energy Conservation Equation ne general fluid (110)2 energy conseivation equation, which relates the change in stored energy with respect to time (the fluid-energy time derivative) to the energy rates in and out (due to the connected flow paths) and the energy addition rate due to other external devices, reduces to the following: d[M x c, x (T-T,,,)]3,of,y, dt a m m as n, fRun une sapn,ofanx 4.10 18
+
kmetat=o H,0).PRHR O m4p60t%ec4u3444to.non: b-ioo395 4,10 6 REVis!ON: 1
.i 1
l 1 bsj The left-hand side of the energy conservation equation is approximated:
.i I
d[M x c, x (T-T,,,)]g,o, pag, . N-~ 'A[M x c, x (T-T,,,)]c,,ga, di .[ At 33 4,10 19-N-= A[M x c, x (T-T,,)],, pan, I
. g At 33 where:
A[M x c, x (T-T,,,)],,,,,,' @ x U-T,,,)],,pasa, c'a' x
=
1 At . M - I
. . i M
, ,,paga,,, - Mr .PRHa, .:
4OM c,,,, x (T,,,,,,,-T,,,) x j n _ n-1 1
""">'"~""">'*~'
Pme, M f.PRHR, x
" , g"'I 3
'~ and: A[M x c, x (T-T,,,)],,p,3,,' @ x PT,,,)],,paga,
= c x At P***=>
At .
=
x (T,,paga -T,,,) x 8 # "> ' c,, a n-1 ,
~
g,PRHR,,a 3.PRH R,,m-1 Pm, 3PRHR, , i l l n
- . . m
- \ap600\sec4\2344-410.non:lb-100395 4.10-7 REVISION: 1
-- .=. --. . - - , - ,
\
l The zone j water-specific heat capacity at constant pressure is calculated as a function of pressure and temperature from the ASMB steam table function CPL as follows: c,, = CPL (Prag,, T,paga,) The zone j steam-specific heat capacity at constant pressure is calculated as a function of pressure and temperature from the ASME steam table function CPV as follows (note that the steam-specific volume is also an output of CPV): c,, = CPV(Ppaga, T,,,,3,,, y,, pan,), For use in the overall system energy balance calculations, the fluid stored energy in the PRIIR is given by the following: U g,o, pan, = U,,,,3, + U 8P "H" 4 3-24 where: N, ,,, U,,,,g , = { M,,,,g,, xc p x (T,,,,3,, - T,,,) 15 4.9-25 and: N, U , pan, = { M ,pgg, x c,, ,' x (T,p,3, - T,,,) J'I 4.9-26 This completes the discussion of the calculations related to the left-hand side of the fluid energy conservation equation. O m:pec4u344-4to.noa:tb-too395 4,10 8 REVISION: 1
i t ,[ P gg
;'\j ' On the right-hand side of the fluid energy conservation equation, the calculation of the outlet fluid energy transport rate (to the SG-2' outlet)is glven by the following:
i
- (to 50) Hp,PRHR t Do $0) f.PRHR
!.4.10-27 l
out Go 80) f.PRHR f.PRHR_ovoes
'Ihe outlet line water-specific enthalpy is calculated as a function of pressure and temperature from the ASME steam table function HCL as'follows (note that the water-specific entropy is also an output of llCL):
Ii r.PRHR_ouii.t " E@reuR, T,,,,y,,, S,, pag,,,,,,,)
- 4.Wh .
'Ihe water temperature is given by that from axial fluid property zone 1 (at the bottom).
The energy addition rate to the fluid due to heat transfer from the metal is given by: i
.f t ' *-- imetal* H,0).PRHR * (tute m.tal= H,0).PRHR + [niet ime metal
- Hpl.PRHR 4.10-29
. + O avo.t t un. m.iai.H,otPRHR l 4 j Due to the small size of the inlet and outlet lines, their contributions in the previous equation are l Ignored, i.e.: l Qi ,i,, ,,,,,,u,oi, pas , = 0.0 4.10-30 G iovan im. maai.Hpi.PRHR = 0.0 1 i- ~ l The contribution from the tube metal-to-fluid heat transfer rate is calculated from the tube metal j energy conservation equation, which is discussed in Subsection 4.10.3. .j Finally, having calculated the change in stored fluid energy (from the left-hand side) and the outlet . fluid energy transport rate and metal-to fluid heat transfer rate (from the right-hand side), the fluid
'energyfonservation equation is rearranged to solve for (infer) the inlet fluid energy transport rate
..(from the HL-2).
4 I m:pc4u344-41a o.:ib.ioo395 4.10 9 REVISION: 1 l . , v
, 1
1 This completes the discussion of the calculations related to the right-hand side of the fluid energy j conservation equation. j i 4.10.3 Tube Metal Energy Conservation Equation j The general metal energy conservation equation, which relates the change in stored metal energy with respect to time (the metal-energy time derivative) to the energy rates in and out (due to heat transfer), reduces to the following for the PRifR HX tube metal (note that here, the " ambient" is the IRWST fluid in which the tubes reside): d[Mxc,x(T-T,,,)],, ,,,,tpan,
" ~ O(rubs mMe Hp].PRHR ~ ruhe metalm!RWST1,PRHR dt 4.10-31 Due to the heation of the thermocouple instmmentation, the tube metal is divided into fou metal segments. Table 4.10-3 lists the tube metal thermocouple channe! ID and tube metal mass data for each tube metal segment, which are required in the tube metal energy calculations.
He left-hand side of the tube metal energy conservation equation is approximated: d[M xc,x(T-T,,,)),,, ,,,,, pan, , A[Mxc,x(T-T,,,)],,,, ,,,,,, pay, dt At 4.10-32 where: A[M x c, x (T-T,,,)],,,, ,,,,,, pan, '~ g tube metal,PRHR, x CP,,be mei,i,PRHR, tube metalPRHR,,n ~ rube metal,PRHR,n-1 t, - t"- 4.10-33 where: N ,3,m ,,,,ynna = Number of PRHR tube metal segments, (4) n, n-1 = Two consecutive time points he integrated tube metal stored energy is also computed for use in the overall system energy balance. He tube metal segment specific heat capacity at constant pressure is calculated as a function of the maap600*c4u344-4to.non:IMoo395 4.10-10 REVISION: 1
i i e k tube metal segment temperature from the metal c, versus temperature tabular data listed in Table 4.10-4, via linear interpolation within the table. No extrapolation at either end is performed; the first and last table points define the applicable range. The first point is (T,,, c,,,,), and the last point is (T,,, c,,,,). This completes the discussion of the calculations related to the left hand side of the tube metal energy conservation equation. On the right hand side of the tube metal energy conservation equation, the tube metal-to-IRWST heat transfer rate is inferred from the IRWST fluid energy conservation equation. Subsection 6.1.2 shows l that the inferred PRHR heat transfer is consistent with the external tube wall-to-IRWST water heat transfer rate. Finally, having calculated both the change in stored tube metal energy (from the left-hand side) and
. the tube metal-to-IRWST heat transfer rate (from the right-hand side), the tube metal energy conservation equation is rearranged to solve for (infer) the tube metal-to fluid heat transfer rate.
This completes the discussion of the calculations related to the right-hand side of the tube metal energy conservation equation. i 4 4 9 3 l t i l i i 4 1 lly . ' ntWwxrwee4u344-4to.non:1b-100395 4.10-11 REVISION: 1
I TABLE 4.101 O INSTRUMENTATION EMPLOYED FOR PRIIR FLUID CALCULATIONS Description item Level (in.) LDP-802 ! Pressure (psig) I'T-107 Oudet Line Liquid Flow Rate (gpm) FMM-8N Zone Temperature [ F]frop Elev (in.) TF-8N 4375 [ Bottom (oudet) to Top (inlet)] TF-805 7.000 TF-809 50.000 TF-811 52.875 TF-803 57.000 LDP Comp Region Temp ('F)/Elev (in.) TF-805 0.000 Bottom (oudet) to Top (inlet) TF-809 29.625 TF-811 57.000 TABLE 4.10-2 VOLUME VERSUS IIEIGIIT TABLES FOR PRIIR FLUID VOLUME CALCULATIONS Ileight Volume * (in.) (in.8) 0.00 0.0 1.75 0.0 7.00 871.3 50.00 1076.7 55.75 1513.4 57.00 1513.4 ' A deadband was included for the volume at the top (inlet) and bottom (oudet) to accommodate level instrument ; fluctuations that were observed in the tests. ' mAquasecau344-4to.non:Ib-too395 4.10 12 REVISION: 1 ,
.1
TABLE 4.10-3 DATA FOR PHIIR TUBE METAL ENERGY CALCULATIONS (PER SEGhfENT) Temperature ID hiass Metal Segment ('F) (Ibm) 1 (Bottom horizontal) TW-801 37.0 l 2 (Bot 1/2 vertical) TW-803 40.0 3 (Top 1/2 verucal) TW-806 40.0 4 (Top horizontal) TW-808 37.0 TABLE 4.10-4 SPECIFIC HEAT CAPACITY VERSUS TEMPERATURE TABLE FOR PRHR TUHE METAL ENERGY CALCULATIONS Metal c, Metal Temperature [Htu/(lhm *F)] ('F) 1 0.1085 70.0 ) [ 0.1109 100.0 0.1175 200.0 0.1223 300.0 0.1256 400.0 0.1279 500.0 0.1297 600.0 l l l
)
U 1 m.upemsecau344 4to..on:Is100395 4.10-13 REVISION: 1
-. .- - ~ . . . ... - . . .
=
4.11 ' ' Reactor Pressure Vessel The reactor pressure vessel (RPV) model includes the lower plenum, core, upper plenum, and upper head. He vessel analysis includes liquid level and mass, core power, steam production due to core power, and flow quality at the core exit. The core vessel is shown in Figure 4.11 1. A total of eight regions are modeled. He actual spans and elevations of the lower plenum, core, upper plenum, and upper head are used. Elevations are relative to the inside bottom of the vessel. Volumes for supporting structures are also included so as to support the use of actual elevations and simplify coding algorithms. Figure 4.11-1 also shows the six axial core heater power steps. Total heater power was varied during the tests; power history data is provided for each test. A top-skewed power distribution was used for all tests. Fluid thermocouple elevations in the core are shown in Figure 4.11-1. The fifteen locations include all elevations outside of the core region. Where multiple thermocouples are provided at a given elevation, the average was used in all analyses. Two issues were identified with the core fluid thermocouples-l. e The core fluid thermocouples are located in instrumented rods, with one rod at the core center j /\ and the remainder at the core perimeter. Fluid tempera'ure histories for core center and ! perimeter thermocouples were found to differ. This was attributed to smaller power-to flow l ratios for the perimeter heater rods than for the center rod due to larger flow areas. In order to ! best represent the average core temperature associated with the average flow area, the center-rod temperatures were used exclusively.
- The core fluid thermocouple histories were noisy. The noise was investigated and found to be unrelated to core thermal hydraulic phenomena. A time-history smoothing method was ,
selected to minimize the noise. This smoothing was applied to all core fluid thermocouples. The smoothing algorithm is provided in Section 4.20. I LDPs available to determine liquid levels are shown in Figure 4.11-2. Level instrumentation issues common to all tests were identified and are summarized below. Values for LDPs spanning the supporting structures tend to be influenced by flow through the supporting structures. The flow effect tends to bias the LDPs toward higher readings and may cause j an LDP to read full during an entire test. Since the bias is a function of flow quality and flow rate, the effect varies during the tests. As a result, DP-111 and DP-ll4 were not useful; they read either full scale or clearly overstated the liquid levels. These LDPs were not used in any analyses. p LDP 112 generally read full scale. Figure 4.11-3 provides an uncorrected plot of LDP-ll2 and the ( adjacent LDP 113 data. LDP-ll3 is directly above LDP-112. During the period around maap600nc4u344-411.no : b-too395 4.11-1 REVISION: 1
1 ) [ ] seconds, LDP-113 clearly indicated that the level had dropped into the span of LDP-112, yet LDP ll2 failed to respond. LDP ll2 was not used in any analyses. The pressure taps used in the vessel analysis are also shown in Figure 4.11-2. Since the pressure taps (IT-107 and IT-108) are located at the top and bottom of the vessel, the pressures should differ due to the effects of water column density. Figure 4.11-4 shows the pressure gradient (Pn.m - Pn.m) for Matrix Tests SB01 and SB18. The gradients are similar for about the first [ ]'6' seconds, after which the gradient for Matrix Test SB01 deviates to a value inconsistent with the height of the water column. This was found to be due to an error in 17-108 for Matrix Test SB01. To minimize the impact of pressure irregularities, RPV analyses are based on one of the pressure taps, which is selected on a test-by-test basis. PT-107 was used for Matrix Test SB01 and PT-108 was used for Matrix Test SB18. The pressure tap irregularities were feund to be on the order of two psi for Matrix Test SB01. The potential impact of a two psi error is related to the importance of the associated change in water and steam properties. For most analyses this effect is relatively unimportant. As discussed in Subsection 4.11.2.1, core steam production calculations may be sensitive to a two psi error. 4.11.1 Core Vessel Model The core vessel model is shown in Figure 4.11-5. Eight fluid regions and six core heated regions are modeled. The selected level instrumentation is shown. Fifteen temperature zones associated with the fifteen thermocouple elevations are also shown. The spans of the selected LDPs differ from the fluid regions and temperature zones and are .shown in Table 4.11-1. LDP levels are relative to their lower pressure taps. To be consistent with other elevations, region levels are defined relative to the inside bottom of the vessel. Region liquid levels are related to LDP readings as follows:
. When the LDP reading is empty or less then empty, the region level is set to either the bottom of the region or the bottom of the LDP span, whichever is greater. In cases where the LDP lower tap is above the bottom of the region, as is the case in the upper plenum, this logic results in a minimum calculated liquid level and a void fraction constrained to a value less than 1.
- When the LDP reading is between empty and full, the density-corrected level is used. The density correction accounts for temperature differences between the fluid in the vessel and the fluid in the reference leg of the instrument line. If this results in a level above the top LDP tap, the next method is used.
m Aap60lAsec4\2344-41 1.non: 1 b- 100395 4,j].2 REVISION: 1
t 0 V = When the LDP reading is full or more than full, the region level is set either to the top of the region or the top of the LDP span, whichever is less. De region is assumed to be full of liquid; the void fraction is set equal to 0. In cases where the LDP upper tap is below the top of the region, step changes in region void fraction and mass will occur when the LDP readings vary between full and less than full. He effects of this level and mass methodology are shown in Table 4.11-2. The fifteen temperature zones bound the thermocouple elevations. The midpoints between the thermocouple elevations define the intermediate zone boundaries. The top and bottom zone boundaries are set equal to the model top and bottom. Fluid and steam properties are assumed constant in each zone. Liquid densities are determined on a region basis using the average liquid temperature for the region. All other zone properties are based on the local pressure and temperature. De local pressure is based on the selected pressure tap, the liquid levels, and the density effects of water, f Region liquid masses and volume-based void fractions are detennined using lookup tables that relate liquid levels to region volumes. These tables account for axial variations in the region cross-sectional areas. Liquid and vapor masses for the temperature zones are then determined by mapping the region results to the zones. 4.11.2 Core Power and Flow Model ne core is subdivided into six axial heater regions that are 6 in, high. De axial-heater-region power distributions were the same for all tests and were constant during the tests. De core total power is calculated by averaging the redundant power instrumentation. The core total power and axial power distribution are defined below. t l Core total power is defined as: 1 {; PWR 7 =1 PWR, 4.11-1 2 ri where: PWR T = Total heater rod power, kW PWR, = Heater rod power data, kW Re heater rod instrument list is provided in Table 4.11-3. The power distribution is provided in Table 4.11-4. ma p60Nce44-4ti.no :1 Moo 395 4.11-3 REVISION: 1
Two methods are provided for calculating core steam production: the Tsat method and the DVI line flow method. Both methods view the core as a steady-state control volume. The rate of core mass j changes and energy changes are not considered. His limits the validity of both methods during the l first several hundred seconds of the transient. During the remainder of the tests, this approximation is I expected to have minimal impact. Two important results of this methodology are:
. The liquid mass flow rate from the lower plenum to core is the same as the mixture mass flow rate from the core to upper plenum.
. 100 percent of the power is applied to heating the core flow.
4.11.2.1 Core Steam Production - Tsat Method Core steam production and saturation line are calculated from core power, core fluid temperature, and local saturation temperature using the following steps: Step 1: Determine the saturation line by finding the intersection between curves of local saturation temperature and fluid temperature. 3 Step 2: Determine the power above the saturation line. Power above the saturation line is assumed to generate steam; power below this elevation is assumed to heat water. Step 3: Determine steam production from power above the saturation line and the enthalpy of vaporization. Fluid temperatures are defined at the thermocouple elevations. The fluid temperature curve is constructed by linearly interpolating between the elevations. The local pressures are also defined at the thermocouple elevations. He saturation temperature curve is defined by determining the saturation temperature at the local pressure at the thermocouple elevations and linearly interpolating between the elevations. Starting at the bottom and searching in the direction ofincreasing elevation, the saturation line is found by finding the intersection between the two curves. Thus, the saturation line is the elevation where core heating achieves saturated liquid. [ The temperature profiles for the entire vessel are used. As a result, the intersection may occur outside of the bounds of the core. If an intersection is not found, or the intersection is not in the core, then: If the core fluid temperatures are above the saturation temperature, the intersection elevation is set equal to the bottom of the core. if the core fluid temperatures are below the saturation temperature, the intersection elevation is set equal to the top of the core. m aap60(bec4\2344-41 1.nce : 1 b- 100395 4,i14 REVislON: 1
/7 V - Power above and below the saturation line is defined by:
P W R ,= FTP, dz 4.11 2 PWRg = PWR.r - PWR, where: z = Elevation, in. 36 = Total height of heater region, in. PWR, = Power above saturadon line, kW 5 PWR, = Power below ' saturation line, kW FTP, = Fraction of total power for region bounding elevation z per Table 4.11-4 Several methods for calculating steam and flow were investigated. The modeling of the core for steam and flow ca!culations is shown in Figure 4.11-6. The calculations are based on properties at die three locations shown. Properties at the saturation line are determined from the saturation temperature ( calculated by interpolation of the saturation temperature curve. Thermocouple elevations outside of i the core and close to the core top and bottom elevadons were selected as proxies; the corresponding thermocouple temperatures and local pressures were used to calculate associated properties. The selected method determines steam production directly from power above the saturation line and the enthalpy of vaporization. Steam production from the core is calculated as. 1 i l M' = PWR* x C 4.11 3 h,,,, l where: C = Conversion constant, kW to Btu /sec. 4 Ki, = Steam production, Ibm /sec. l hg., = Enthalpy of vaporizadon at saturation line, Blu/lbm The viability of this method for determining steam production is dependent on the accuracy of the saturation line calculation. Since the saturation line is the intersection of the fluid and saturation temperature curves, the accuracy of this elevation is a function of the slope of the two temperature curves. The saturation temperature curve is a function of local pressure and varies only slightly over m:\ap60L%ec4\2344-411.non:1b-100395 4,}15 REVISION: 1
the span of the core. The fluid temperature gradient varies widely during the transient as shown for Matrix Test SI301 in Figure 4.11-7. The gradient is very small during steady state operation; and is small during the first several hundred seconds of the transient, and during long term cooling. Since the applicability of this method is a function of the fluid temperature gradient, the range of applicability may vary from test to test. 4.11.2.2 Core Steam Production - DVI Line Flow Method Core steam production and saturation line are calculated from core power, core inlet enthalpy, and the DVI line flow rate using the following steps: Step 1: The core inlet enthalpy is based on the fluid temperature and local pressure in the lower plenum. Starting with the bottom core heated region, the enthalpy rise for each of the six heated regions is determined from the DVI line flow rate, less break flow from the cold leg, enthalpy at the bottom of the region, and the core power in the region. Step 2: The saturation line is identitled by locating the first region where the enthalpy exceeds that of saturated liquid and interpolating the enthalpy rise to determine the saturation line elevation in that region. The power above the saturation line is then calculated. Step 3: Core outlet quality is determined using the steam tables and the enthalpy at the top of the core. Steam production is then determined using the quality and DVI line flow rate. The saturation line is calculated as: C (h t, ,, - h,3) N1 = FTP, dz, for si > 0 4 11-4 bottom of core sat line = bottom of core, for $1 = 0 where: Ni = Liquid flow rate into core = DVI line flow - cold-leg break flow for hot-leg breaks and tests without breaks, the cold-leg break flow is zero h,, ,, = Liquid enthalpy at saturation line m:\ap600wec4us44-4 t i .non: i s ioo395 4,11 6 REVISION: 1
f h,, i = Liquid enthalpy at core inlet Core outlet flow quality can be calculated from the liquid flow rate and a core energy balance: PWR7 h = h,,, 2
+ xC 4.11-5 K1 X = f (h,, P7c) 2 4.11-6 The core outlet flow quality shown above applies to the flow exiting the core and not mixture conditions at the top of the core. Thus, core steam production and flow void fraction can be calculated directly without consideration for slip coefficients. The core steam production and flow void fraction are calculated as:
61, = X 2N1 4.11-7 X,p ct2 = 4.11-8 X2Pt.2 + (1 -X2 )Pp where: h2 = Mixture enthalpy at core outlet pp pr.2 = Vapor and liquid density at core outlet, respectively X2 = Flow quality at core outlet cc2
= Flow void fraction at core PcT
= Pressure at top of the core f (h, P) = Steam table function for quality he viability of this method for determining steam production is dependent on the assumption that DVI line flow minus cold-leg break flow is representative of the flow into the core from the lower plenum. This assumption clearly does not apply during steady state operation. During the first several hundred seconds of the transient, this flow is not representative due to draining of various components, a which includes the steam generators, hot legs, cold legs, and pressurizer. After approximately
[ ]** seconds, this is a good representation of the core inlet flow. The validity of the flow approximation for the DVI line flow method is further investigated in Subsection 6.2.1. This method is relatively insensitive to the core temperature gradient. The effects of pressure errors are primarily limited to the change in mixture enthalpy at the top of the core. As a result, this method O is well behaved during LTC and is preferred over the Tsat Method for this phase of the transient. Q m:up60msu4u344 411.non:ib.100395 4,)).7 REVISION: 1
4.11.3 Energy Italance The total liquid and steam energy is calculated by summing the results for the temperature zones as follows: U, = {
- i. m . non M,,,c,,o(T,,,-32'F) 4.11 9 U, = {
=mp.m. sone.
Mgc,,g(T,-32*F) 4.11-10 l l where the subscripts: f = Liquid phase of water g = Steam i = Temperature zone number The rate of energy change for the total reactor metal mass (includes downcomer) is based on 33 metal O i regions associated with metal thermocouples. Approximate metal masses were estimated for each l metal region. The rate of energy change of the metal is calculated as: 0.., = E M..y c,,o AT**" metal regions bI where the subscript: J = Metal region number 1 t
'Ihe rate of energy loss from the reactor vessel (including downcomer) to the environment is calculated using the sump methodology of Subsection 4.9.5.5. The reactor vessel and sump are similar in that they are insulated vertical tanks. The reactor vessel and sump modeling differences are limited to the constant multiplier in the equation for the free convection heat transfer coefficient. Subsection 4.9.5.5 defines the applicability of the detailed energy loss calculations for the CMT cold-leg balance line to the sump. The applicability relationship also applies to the RPV.
O l m Aap60thsec4\2344-41 1.non: 1 b- 100395 4,jj.8 REVISION: 1
1 4 s O V For reactor vessel-to ambient energy loss formulas, see Subsection 4.9.5.5 and related sections. The , textor vessel surface was subdivided into four areas, with a different constant multiplier for each area.
- These multipliers are compared to the sump multiplier in Table 4.11-5. The total heat loss is
- determined by summing the heat loss for the individual areas.
i t t l I } 4 i i k i i 1 i i k 4 l 1 i l 9 i e L 4 i 4 r m:Wec4u344-411.aon:lt>100395 4.11 9 REVISION: 1 4
+b - . - - - t- r- r - ~ --r---m.
I i TAllLE 4.11 1 O CORE VESSEL MODEL GEOMETRY i l Level llottom of Ilottom Top of l Region Instrument Region of LDP Region Top of LDP Lower plenum LDP-106 0 2 6.41 10.22 LDP 107 Lower core plate LDP-118 6.41 2 9.91 10.22 Core LDP-138 9.91 10.22 50.86 50.17 Core plenum LDP-139 50.86 50.17 52.75 74.13 Upper core plate LDP-139 52.75 50.17 53.5 74.13 Upper plenum LDP-113 53.5 58.71 74.42 77.83 Upper support plate LDP-115 74.42 77.83 77.42 97.58 Upper head LDP-115 77.42 77.83 97.58 97.58 TAllLE 4.112 MASS AIETIIODOLOGY EITECTS Region Minimum Level Effects Maximum Level Effects j Core Minimum liquid mass > 0 Small step changes in mass and void fraction Maximum void fraction < 1 when level varies to/from top of LDP span Upper plenum Minimum liquid mass >> 0 None Maximum void fraction << 1 Upper head Minimum liquid mass > 0 None Maximum void fraction < 1 TAllLE 4.11-3 IIEATER ROD LNSTRUMENTATION Pi Instrument Name P, KW-101 P2 KW-102 P3 KW-103 P, KW-104 9 maagoAsecau344-411.n<m:1b loo 395 4.11-10 REVISION: 1
l Or\ _ l TABLE 4.114 POWER DISTRIBUTION Axial Heater Region Fraction of Total Power in Region 6, top heater region 0.19367 5 0.24293 4 0.22487 3 0.1798 2 0.1172 1, bottom heater region 0.04153 TABLE 4.115 CONSTANT MULTIPLIERS FOR THE FREE CONVECTION HEAT TRANSFER COEFFICIENT l , Area Coefficient Basis j Sump 0.09 l Reactor Vessel: Vertical Surface 0.09 Simplined equation for free convection to air from a vertical surface. Reactor Vessel: Hot & Cold Leg 0.15 Simpiined equation for free convection to air from Area borizontal pipe, OD = 5.75" Reactor Vessel: DVI Nozzle Area 0.21 Simplified equation for free convection to air from horizontal pipe, OD = 1.66" Reactor Vessel: Bottom Flange 0.09 Simplified equation for free convection to air from Surface Area heated plate facing downward (~\ , (_ m:Wecauu4411.non:151oo395 4,11 11 REVISION: 1 I I
\
TABLE 4.11-6 O OSU TEST ANALYSIS PLOT PACKAGE FOR SECTION 4.11 8 Plot Number Component Variables Units Description 1 Reactor Vessel N/A N/A Reactor Vessel Geometry 2 Reactor Vessel N/A N/A Reactor Vessel Level and Pressure Instrumentation 3 Reactor Vessel LDP-112 in. Uncorrected Levels Results of Test SB18 LDP-113 4 Reactor Vessel Delta-P = psi Pressure Gradients for Tests SB01 and SB18 l'F-108 - FT-107 i 5 Reactor Vessel N/A N/A Reactor Vessel Model l i 6 Reactor Vessel N/A N/A Modeling of Core Steam and Flow l Calculations 7 Reactor Vessel Core Temperature *F Core Temperature Gradient Rise l l l 9 l 9 m4Wmec4u344-41 t.noo:1b.100395 4.I1-12 REVISION: 1
f.~x i ELEVATION IN INCHES THERMOCOUPLES 97.58 gi,33-REGION 8 87.t2 UPPER WAD
~ BO.ll 77.42-REGION 7/ UPPER SUPPORT PLATE 74.42-
- 67.73 REGION 6 - 65.33 UPPER PLENUM 58.50 -
53 50 - 53.00 d 75-- REGION 5/ UPPER CORE PLATE 50.86 - REGION 4/ CORE PLENUM - 51.06 i 49.13 - - 49.13 i 55 Nhio l l
- 43.l3
<3.t3-5hw I N 37.13 - - 37,13 6$+"
i si. is - REGION 3 - 3t,13 CORE
$$m I N
- 25. is - - 25.13 6 N Z n 19.13 - - 19.13 5$-
EN 13.13 - 10.50 9.91 - REGION 2/ LOWER CORE PLATE
~ 0.22 l
6.41 REGION I LOWER PLENUM
- 2.00 0.M L INSIDE BOTTOM OF LOWER PLENUM
. ;"% l
\. . Figure 4.11-1 Reactor Vessel Geometry l m:\apN%ec4\2344 411.non:It>.100395 4.11 13 REVislON: 1
O r REGION O M UPPER HEAD ()LT-IIS REGION 7/ UPPER SUPPORT PLATE ()0Pil4 REGION 6 (-)LN t13 UPPER PLENUM LDP il2 REGION 5/ UPPER CORE PLATE y,g,g REGION 4/ CORE PLENUM OF
~ i $$w I N Eym I N
()LOPil0 w REGION 3 3h I N CORE OLDP-ise e 3gN g () LOP-109 Q-REGION 2/ LOWER CORE PLATE LT 106 REGION I LDP 107 LOWER PLENUM Figure 4.112 Reactor Vessel Level Instrumentation r mu emsecau344-4 t.noa:I M oo393 4,11 14 REVISION: 1
4 i
- i
- i. '
4 f .. i iO : 1 1. 1 . I 2 4 i
;. FIGURES 4.113 AND 4.114
' ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT i-c.
1 i i' a i , i l i, 1 4 s i i i i 5 3' 1-4 i i 1. ,t i
- i 1
i-i l 1 i l i i a !~
- - m
- W4unuti.nas:15im995 4.11-15 REVISION: I
- j. o ,
p
i
.\ l l
l ELEVATIm IN INCHES TEMPERATURE ZONE
- 97. 2 15 l
' ~
REGION 8 l 87.12 UPPER HEAD l l 14 77.42-
, , REGION 7/ UPPER SUPPORT PLATE 13 REGION 6 UPPER PLENUM i2 58.50 - ,
l l1 ; h,'3,
~
REGION 5/ UPPER CORE PLATE ( REGION 4/ CORE PLENUM 10 ( 49.13 - g "fe I N ' 43.13 - 8 I 6 hu, 37.13 - En 1 l 7 6 Z h, N 31.13 - REGION 3 6 g CORE I N 25.13 - 5
~
19.13 - 4 6 - l Z n i 13.13 - 3 9.91 - REGION 2/ LOWER CORE PLATE 2 ; REGION i I LOWER PLENUM I---INSIDE BOTTOM OF LOWER PLENUM D (V Figure 4.11-5 Reactor Vessel hlodel . m:wpuonc4s234ui t.non:th ioo395 4,11 17 REVISION: 1
O l
$1 = total flow rate A, = steam flow rate Proportise keliEE90 0
N ha X, as core exit Pu Po P. h,IeN g k - sat line i P. i kt core entrance il U Figure 4.116 Modeling of Core Steam and Flow Calculations m%Wesec4u14+4 1.noo:ts100395 4.11 18 REVISION: 1
{ FIGURE 4.117 IS NOT INCLUDED IN THIS ' NONPROPRIETARY DOCUMENT ; 1 i. i i l t i i
- l
- i. I i l t
4
- ( ,
h r i i !' m:We4u344-411. :ib.noims 4.11 19 REVISION: 1 i i ew. - . _ m -_--ma .,,,r ..%.,m... . . . . w - - _,v- - - - - m,.- ,... . --c, e.%--4 .e- + e-- .*rr<- - W*"
4.12 Downcomer
'Ihe downcomer consists of the annular volume in the RPV surrounding the core and upper plenum.
c Its component interfaces include two DVI lines, four cold legs, reactor vessel lower plenum, and
. reactor vessel upper head bypass gaps. The downcomer is modeled separately from the other reactor vessel regions. Parameters calculated for the downcomer include liquid level, mass, flow from the DVI lines, and flow in the lower plenum. ,
Downcomer instrumentation used to analyze the test includes two LDPs that measure liquid level, and thermocouples that provide fluid and metal temperatures.' The LDPs are positioned at different circumferential locations and span nearly the full height of the downcomer.
, 4.12.1 Downcomer Level and Mass The downcomer fluid temperature distribution is defined by dividing the volume into a vertical array of temperature zones. Where multiple thermocouples are provided at a given elevation, the average l temperature is used. Axial midpoints between the thermocouple elevations define the intermediate zone boundaries. The top and bottom zone boundaries are set equal to the top and bottom of the downcomer. Fluid and steam properties are assumed to be constant in each temperature zone. Zone propestles are based on local pressure and temperature. Local pressure is based on the pressure at the l top of the reactor vessel, the liquid levels, and the density effects of water. I I
The downcomer liquid level is determined by averaging the two LDP readings and applying density j corrections for the density difference between the reference leg and the measured water column. l Downcomer liquid and vapor mass are determined using lookup tables that relate liquid levels to l volumes. These tables account for axial variations in the cross-sectional areas. Mass calculations are done on a temperature-zone basis and employ the temperature-zone densities. 4.12.2 Fluid Stored Energy For use in the overall system energy balance calculations, the fluid stored energy in the downcomer is given by the following: Ugo,nc = U,,oc + U ,oc 4.12-1 where: ! l
~_
U,,oc = {M coc xcg x (T.oc r - T,,) 4.12-2 m i n nwec4u344-412. o :ib-too395 4.12-1 REVISION: 1
and: 1 i U,, oc = { M,nc x c,,, x (T,x - T,,,) 4.12-3 J=4 a 4 J 4 4 5 3 i 1 0, f [ m:\ap6aossec4u144-412.noo: s ti-loo 395 4.12-2 REVISION: 1
O . TABLE 4.121 VOLUME VERSUS HEIGHT TABLE FOR DOWNCOMER FLUID VOLUME CALCULATIONS Ileight Volume (in.) (in.') _ 0 0 4.41 0 6.0 281.0 22.0 3108.0 38.0 5936.0 78.0 13,0N.0 78.11 13,998.0 83.62 13,998.0 98.0 13 998.0 O l' O m:Wec4\2344-412.non:lb-100395 4.12 3 REVISION: 1
j 4.13_ Steam Generator Primary Side The AP600 includes two SGs, which under normal operating conditions provide the heat sink for reactor core heat removal and steam source to power the turbine. The OSU test facility incorporates two simulated SGs. For the purposes of developing mass and energy balance equations to relresent the
$~ ~a1 b$aulic performance of the RCS primary side, the pri nary side of the simulated SGs is considered as three interconnected sections:
- Inlet plenum
. Tubes (both uphill and downhill sides)
. Outlet plenum The mass and energy balance equations are developed by considering the equations for each of these sections.
4.13.1 Inlet Plenun I l A general mass balance equation for the inlet plenums may be wTitten as:
, = Kf gt x -M e-Tuss %J l where the subscripts:
IIL X = Hot leg X = 1 (SG 1) X = 2 (SG 2) IP = SG inlet plenum IP-TUBE = Interface between inlet plenum and tube bundle of SG Similarly, the energy equation for the SG inlet plenum is written as: d(M T)* * " NI st 4.13 2 Qsa " C r x h gt uix - S Tuss r sa x hy ,3,3, y, g
- Om so x MerAL ~O F So x AMB i
I O l m%oxbu4u)44an3. o :ib.noo395 4.13-1 REVISION: 1
where the subscripts: SG X = Steam generator X = 1 (SG 1) X = 2 (S0 2) AMD = Ambient conditions associated with inlet plenum METAL = Metal mass of inlet plenum MIX = Fluid mixture conditions 4,13,1,1 Mass Halance We mass stored in the steam generator inlet plenums is calculated from level measurements in the _ plenums. De level measurements must be compensated for temperature differences between the fluid in the plenums and that in the sense lines of the instruments. De data channel ids of the level, pressure, and temperature instruments to be used in calculating fluid mass in the inlet plenums are limd in Table 4.13-1. The following approach was used in performing the temperature compensation of the level instrument output and calculating the fluid mass in the SG inlet plenums. . Step 1: First, compensate the readings of die inlet plenum level transducers to account for temperature differences between fluid in the plenum and fluid in the reference legs of the instrument lines. As noted above, the channel ids of the two level transducers to be compensated, one for each SG, are identitled in Table 4.131. He instruments used to measure local pressure and fluid temperatures to be used to accomplish the compensation are also identified in Table 4.131. The LDPs, which provide the SG plenum levels, provide incorrect readings when the pumps are running.
- 1) Ec plenum level is to be based on the density-corrected LDP, no pump-flow corrections to be included in the level calculations.
- 2) ne plenum liquid volume and mass is to be based on two alternatives:
a) Plenums assumed to be liquid solid during pump flow, b) Liquid volume and mass to be based on density-corrected LDP after pump flow stops.
- 3) The time period of pump flow to be defined as:
a) Starting at the beginning of the test. magawc4u344 413.netti loons 4,13 2 REVISION: 1
b) The time period ends a user-input number of seconds after the flow in both cold legs attached to a SG drops below a user-selected value. Step 2: The local pressures and temperatures measured by the instruments identified in Table 4.13-1 are used as inputs to calculate the thermodynamic properties of the fluid in the in:et plenums: p, = p, (P. T) p, = p, (P, T) 4.13-3 h, = h, (P, T) h, = h, (P,T) where the subscripts: f = Liquid phase of water g = Steam 1 The data channel ids of the instruments to be used in the calculation of the thermodynamic properties of the fluid in the SG inlet plenums are given in Table 4.13-1. Step 3: Using the compensated liquid level and the inlet plenum volume as a function of height, ( the volume of liquid in the plenum is calculated: -
% )')' \
I V,, , so x = V(1),3c x x LDP-XXXcour 4M l where: I V = Volume, ft.' . V(1) = Volume as a function of elevation, ft.8/ft. (See Table 4.13-2) LDP-XXXcoup = Compensated fluid levels data from level transducer identified as
- XXX = 209 (SG-1 inlet plenum) (see Table 4.13-1) l XXX l = 214 (SG-2 inlet plenum) i Step 4
- The liquid mass inventory in the SG inlet plenum is calculated as:
I Me so x t " Pr. e so x x V, so x 4.13-5 l l [J mWec4u344-413.nelb-too395 4.13-3 REVISION: 1
he mass of vapor in the SG inlet plenum is then calculated as: 4.13-6 M s. e sa x " Ps. e sa x x (VrorAt sc x - Y .r msa x) where: VTOTAL, = Total volume of the SG inlet plenum ofinterest Step 5: The total fluid mass in the inlet plenum of the SG of interest is then calculated by summing the mass associated with each phase in the plenum: M, sa x = (M,,,+ M,,,)30x 4.13-7 Step 6: The rate of change in mass inventory of the SG inlet plenum may be approximated by differencing two consecutive calculated values ofliquid mass: dM r sa x M e sa x (Mesax.,- Me sa x, i-i) 4.13-8 dt At t, - t, i where the subscript: 1 = Index of the data and time arrays 4.13.1.2 Energy Balance Equation 4.13-2 defines the general form of an energy balance for the SG inlet plenum. 'Ihe plenum may contain both vapor and liquid at the same time. The energy balance must account for both phases of the working fluid. Thus, the left hand side of Equation 4.13-2 may be expanded as: d(Me sa x(T-T,y)) d(T,) d(M r. & sa x) cp = c,, e M.escx r dt " ' ~ "" dt 4.13-9 d(T ) d(M )
+ c p,, M , e sa x "' 8 8 ~ ""
dt d where: Tag = Reference temperature,32*F m:\,sp60(Avec4\2344-413.noa:1 b.100395 4.13-4 REVISION: 1
. . .- - .= . . -
- p. i r
(, ) Writing the preceding equation in its difference form, the rate of change of energy associated with the -
- fluid in the SG inlet plenum is calculated:
d (M,, sa x (T-T,3,)) AT, AM,,,,3ax P gg P. f f. IP So x g p, f f ~ REF} g , 4.13-10 AT AM "*
+ c,, , M,, i, sa x y + c,, , U, - T,3,) _.
Re data channel ids of the instruments to be used to evaluate the thermal transport properties for this calculation are listed in Table 4.13-1. Note that the value for T, will be taken as the minimum of the measured temperature and saturation temperature. The value for T, will be taken as the maximum of the measured temperature and saturation temperature. The energy rate of change in stored energy in the SG inlet plenum metal components is expressed as: l dTMETAL
- OrP so x werAL =MiP so x werar. x c,, y3.ru x 4.13-11 O \._)
1
\
Representing the preceding equation in its difference form: 1 1 i AT " OrP so x seral - M,,so ,y,7,t x c,, y37, x 4.13-12 For these calculations, the outside surface temperature of the SG inlet plenum will be used to represent the average temperature of all the inlet plenum metal. The specific data channel ids to be used for these calculations are listed in Table 4.13-1. De heat loss from the inlet plenum of each SG may be represented in the form: i 1 Q IP SG X AMB = A ,p 3a x xq 4.13-13 where: A = Effective area of SG inlet plenum being considered, ft.2
,3 q = Flux, Bru/(hr-ft.2)
(Flux meters HFM-301 and HFM-302 are used.) m:pec4us4+413.non:ib.noo395 4.13-5 REVISION: 1 L =
The rate of energy loss to the ambient may be calculated using data collected from the inlet plenum surface temperature instruments identified in Table 4.13-1. The quality in the inlet plenum is calculated as: X= M s. e sa x 4.13-14 M,, ,3a x + M . re so x For use in the overall system energy balance calculations, the fluid stored energy in SG inlet plenums is given by the following: U m sa x = U r e sa x + U,, ,3a x 4.3-15 where: U,, , 3a x = Mr . e sa x xce,, e sa x x (T.r e sa x - T,,,) 4.13-16 and: Us. e sa x = M s. e so x
- Cp,, e sa x x(T,,,3ax - T,,,) 4.13-17 4.13.2 Steam Generator Tubes The general mass balance equation for the tubes in each SG may be written as:
dM rter.s so x
"*' S * """ S
- dt where the subscripts:
TUDE = Tube bundle of an SG TUBE-OP = Interface between ttibe bundle of SG and SG outlet plenum 4.13.2.1 Mass Balance 'lhe mass stored in the tubes of each SG is calculated from level measurements across the hot leg (up-hill side) and cold leg (down-hill side) of the tubes. Like other level measurements taken with the test facility, these measurements must be compensated for temperature differences between the fluid in m:wuouc4u344-413.non:15too395 4.13-6 REVISION: 1
--.. . .~ .. .- .
J ( the tubes and in the sense lines of the instruments. The data channel ids of the level, pressure, and l' temperature instruments to be used in calculating fluid mass in the inlet plenums are listed in Table 4.131. I I The installation of the common reference leg for the level and pressure transducers on the SG tubes l provides for the fluid inventory of the leg to be at the same temperature as the SG secondary side at l
; the initiation of a test. Once the primary side pressure drops below that of the SG secondary side, the reference leg begins'to boil off. This boil-off continues until all fluid in the reference leg that was at l or above primary loop saturation temperature has flashed. The boil-off of the reference leg causes the j- level and pressure transducers to give false indications of water in the tubes and a change in pressure i at the top of the tubes. Furthermore, refilling the SG tubes with water is not possible, as the SG J
secondary side acts as a heat source that is above the saturation temperature of the primary side. Thus, logic has been implemented in the code that, once drained, precludes the tubes from refilling. a This approach was taken as data indicated the tubes drain before the primary-side pressure decreases below that of the secondary side. I i l l 4 The following approach is to be used in performing the temperature compensation of the levels
- instrument output and calculating the fluid mass in the SG inlet plenum.
} Step 1: Compensate the levels readings associated with the tube level transducers to j
'O account for temperature differences between fluid in the tubes and fluid in the U reference legs of the instrument lines. As noted above, the channel ids of the four level transducers to be corr.pensated, two for each SG, are identified in 4- Table 4.13-1. The instruments used to measure local pressure and fluid temperatures to be used to accomplish the compensation are also identified in Table 4.13-1.
d
- Step 2
- The local pressures and temperatures identified in Table 4.131 are used as inputs to the ASME steam tables to calculate the thermodynamic properties of the fluid in
- the inlet plenum
p, = p,(P, T) p, = p , (P, T ), 4.13-19 h, = h,(P, T ) h, = h, (P, T ), Step 3: Using the compensated liquid level and the volume in the SG tubes as a function
]
of height (see Tables 4.13-3 and 4.13-4), the volume of liquid in the tubes of a ; given SG is calculated as: l V,, w,g3 xi, = V(1) ,,,n x LDP-XXX,, 4.13-20 V 1 4-. . mAap60owec4\2344-413.non: t h 100395 4.13-7 REVISION: 1
}
where the subscript: XL = Side of the tubes of a given SG being considered: XL = IIL (hot-leg side) XL = CL (cold-leg side) The total volume of liquid in the tubes of a given SG is then calculated as: V,, .rusas " 4' Vr Tuses ct + V. rTuses at It follows that the volume of vapor in that SG tubes is calculated as: V s. Tunes
" Vtubes TOTAL f. tubes where the subscript:
TOTAL = Total volume of the tubes bundle of a given SG, ft.'
- Step 4
- Accounting for a difference in fluid temperature between the hot leg and cold leg sides of the SG tubes, the liquid mass inventory in each side of the SG tubes of interest is calculated as:
l l M r. Tunes xt Pr x V. Tuses r xt 4'OU The total 11guld mass in the SG tubes is then calculated as: M,, 7u,33 = 4.13-24 M r. Tunss et + M .r Tunes ut The mass of vapor in the tubes of the SG of interest is then calculated as: hi .s Tusss
" 4.13-25 P s. Tunss x (V.rOTAL - Vr . Tusss )
Step 5: The total fluid mass in the tubes of the SG ofinterest is calculated by summing l the mass associated with each phase in the tubes: hi russs
- IM .r Tunes + M .s Tusss) 4ED l
9 mAapMXAsec4\2344-413.non: l b.100395 4.13-8 REVISION: 1
t . 's Step 6: The rate of change in mass inventory may be approximated by differencing two consecutive calculated values ofliquid mass: b8ES ( BES. i
~
BES. i-
, ES , 4.13-27 dt At t, - t,,,
where the subscript: 1- = Index of the data and time arrays ne total mass flow rate from the inlet plenum into the tubes of a given SG is calculated from Equation 4.13-1. With the change in mass storage of the tubes of a given SG calculated from Equation 4.13-26, the terms of the general mass balance on the SG tube bundle, Equation 4.13-18, may be rearranged to solve for the mass flow from the tubes into the outlet plenum of interest. 4.13.2.2 Energy llalance Equation 4.13-15 defines the general form of an energy balance for the SG tubes. The tubes may contain both vapor and liquid at the same time. He energy balance must account for both phases of p the working fluid. Thus, the left hand side of Equation 4.13-15 may be expanded as: i d(M T)russs d(T f. Tusu) Cp
*C.f P f. TUBES - T,37) d (M,, . ruses) + c,, f , M . Tunes g 4.13-28 d(T,, . ruses)
+ C . a hs. TUBES ~ REF) p g
d(M,,p .rusa)
- C , g gM,, . ruses where:
Tat, = Reference temperature; 32'F l l O V maaraxwc40344 413.noa:ib.ioo395 4.13-9 REVISION: 1
Equation 4.13-28 may be further expanded to specifically address the liquid volume on the hot leg (" uphill" side) and cold leg (" downhill" side) of the SG tubes, and written in its difference form to operate on the data: d(MT)793g3
+ C . t (T,AM . Tusss r
e, - c, , Mr . Tuses ryAT. Tusss p - Tm) 4.13-29 AM,, 7y3,3
+ c,, , M, . ruses AT5 rvass + pC , (T, - Tm) y De data channel ids of the instruments to be used to evaluate the thermal transport properties are listed in Table 4.13-1. Note that the value for T, will be taken as the minimum of the average measured temperature and saturation temperature. De value for T, will be taken as the maximum of the average measured temperature and saturation temperature.
He equation for the rate of change in internal energy of the metal of a tube bundle is wTitten as: dTruss so x MErAL 4.13-30 Oruss sa x Merrt =MTuss SO X METAL Cp, METAL g Representing the rate of change in energy stored by the metal of the tubes of each SG in difference form: TTuss SG X MsTAL
= h 4.13-31 Q Tuss, SG X McTAL '8E, SO X McTAL x cp, METAL x g
De data channel ids to be used for the metal energy storage calculations are listed in Table 4.13-1. He quality in the hot-leg side of the tubes is calculated as:
. Tuss- P X= 4.13-32 M ,7y3s.op + M,7e,so, For use in the overall system energy balance calculations, the fluid stored energy in the SG tubes is given by the following:
UTuss so x = U.rTuss sa x + U ,sTuss sa x 4OU O m:pc4us44-413..on: wino 395 4.13-10 REVISION: 1
i. 4 I f] Q where: 4
~
U,,7u3,3a x = M .r was so x x c,,, x (T.r ruas so x - T,,,) 4.13-34 and: i j U s. Tuss so x = M s. Tuss so x x c,,, x (T s. Tuss so x - T,,.) 4.GM 4.13.3 Outlet Plenum A general mass balance equation for the SG outlet plenums may be written as: 4 dM I " -
'asers x e.CL x/Y dt where the subscripts:
C CL X/Y = Designates cold legs where:
- X/Y = 1/3 (SG-1) l X/Y = 2/4 (SG-2) l OP = SG outlet plenum !
TUBE-OP = Interface between tube bundle and outlet plenum of a SG l r : Similarly, the energy equation for the t,G outlet plenum is written as: d(MT) P 4 c, =s ! - - - 4.13-37 4.13.3.1 Mass Halance The mass stored in the SG outlet plenum is calculated from a level measurement in the plenum. Like other level measurements taken with the test facility, these measurements must be compensated for temperature differences between the fluid in the plenum and that in the sense lines of the instruraents. j The data channel ids of the level, pressure, and temperature instruments to be used in calculating fluid mass in the inlet plenums are listed in Table 4.13-1. The approach taken to compensate these level
- readings is identical to that of the inlet plenum and will not be repeated. The volume versus height j values for the SG outlet plenum are provided in Table 4.13-5.
t*
'd i
m:pc4u3444ano : b too395 4.13-11 REVISION: 1 l
Similarly, calculation of fluid mass in the outlet plenum, and mass flow rates into and out of the outlet plenum are the same as those performed for the inlet plenum and are not repeated here. 4.13.3.2 Energy Italance The calculation of an energy balance for the SG outlet plenums is similar to that performed for the inlet plenums and is not repeated here. The data channels used in the calculation of the energy balance for the outlet plenums are listed in Table 4.13-1. Since there is no instrumentation on the outlet plenums, heat loss is calculated by scaling the heat loss from the inlet plenum. He metal energy and heat loss to ambient calculations are performed using the inlet plenum surface temperature measurements. The heat loss from the outlet plenum from each SG is calculated as: T -T Qor sa X AMS " P SG X AMB
- T ,f -Tamb 1P where:
T, = Ambient temperature Toe.t
= Fluid temperature in the outlet plenum T,,, = Fluid temperature in the inlet plenum l
l 9 I
- m
- vesecau344-413.non:Ib too395 4.13-12 REVISION: 1 l
l
l l D
\
TABLE 4.13-1 DATA CIIANNEL ID FOR SG INLET l'LENUM MASS AND ENERGY CALCULATIONS Data Data Channel Channel ID Data Channel ID for for Pressure ID for Fluid I Location Liquid Levels Transducers Thermocouples Notes SG-1 inlet plenum LDP-209 PT-205 TF-205 P= i-301 TF-205 Hot-leg side LDP 215 Irr-201 TF-211 TF-217 SG-1 tubes TF-217 Cold leg side LDP-219 FT-201 , TF-203 CL-1 LDP-213 l'T-201 TF-201 Subtract elevation SG-1 outlet head of water in plenum tubes to calculate CL-3 LDP-211 l'T-201 TF-203 local pressure SG-2 inlet plenum LDP-214 i'T-202 TF-206 P= i-302 V TF-206 )
}Iot-leg side LDP-218 I'T-2(M TF-212
, TF-218 I I SG-2 tubes TF-218 I Cold-leg side LDP-222 i'T-204 TF-2N CL-2 LDP-210 l'T-2(M TF-202 .ubtract elevation i SG-2 outlet head of water in CL-4 LDP-212 PT-2M TF-2M u s to dcubte kx:al pressure
)
\
i , l l 2 l
)
1 O mh;wwcc4u344-413.non:ib-loo 395 4.13-13 REVISION: 1 i
TABLE 4.13-2 O VOLUME VERSUS IIEIGIIT TABLE FOR STEAM GENERATOR INLET PLENUM Height (in.) Volume (in.8) l 0.25 105.8
]
125 1973 2.25 205.8 3.25 427.7 4.25 - 560.2 5.25 700.3 6.25 830.9 6.75 897.0 7.25 963.1 8.25 1095.4 9.25 1227.7 10.25 1359.9 11.25 1492.2 12.25 1624.5 13.25 1756.7 14.25 1889.0 15.25 2021.2 ) O m:\ap600\sec4\2344-413. con:1b-100395 4,13.]4 REVISION: 1
TABLE 4.13-3 VOLUME VERSUS HEIGHT TABLE FOR STFAM GENERATOR TUBES (DOWN HILL SIDE) , Height (in.) Volume (in.') 0 0 102.0 3918.0 TABLE 4.13-4 VOLUME VERSUS HEIGHT TABLE FOR STEAM GENERATOR TUBES (UP HILL SIDE) Height (in.) Volume (in.5) 0 0 102.0 3918.0 O\ V 1 1 4 i l O mWwc4u344 413.noa:Ibloo395 4.13-15 REVISION: 1 1
l TAllLE 4.13-5 O VOLUME VERSUS HEIGIIT TABLE FOR STEAM GENERATOR OUTLET PLENUM Ileight (in.) Volume (In.') 3.45 42.73 4.45 78.74 5.45 136.99 6.45 214.72 7.45 309.01 8.45 416.90 9.45 535.39 10.45 661.46 11.45 792.08 l 1.95 858.15 12.45 924.32 13.45 1056.59 14.45 1188.85 15.45 1321.12 16.45 1453.38 17.45 1585.64 18.45 1717.91 19.45 1850.17 20.45 1982.43 O m:We44344-413.non:lt> 100395 4,]3 16 REVISION: 1
l 1 (" 3 4.14 Steam Generator Secondary Side N 1 4 The design of the AP600 includes two SGs. Under normal operating conditions, the SGs are a heat sink for reactor core heat removal and a steam source to power the turbine. During a transient, the secondary side becomes a potential heat source to the primary side when the primary side depressurizes and cools below the temperature of the secondary side. This subsection presents the equations for SG-1 and SG-2. The calculations are the same for both - components. The notation is as follows: SG SS X = SG secondary side in general SG SS 1 = SG-1 secondary side specifically SG SS 2 = SG-2 secondary side specifically Other notation in this section is as follows: f = Liquid phase of water g = Steam 4.14.1 Inputs and Assumptions A general mass balance on the SG secondary side may be expressed as: dM" "' NI S so ss x 4'I4'I r, sa ss x dt For small time-steps, the time rate of change of mass in the SG secondary side may be approximated as: dM sa ss x Mi sa ss x Mso ss x. , - M 8 88 * '-' 4.14-2 di At t, - t,.i where the difference terms represent data from two consecutive calculations or data scans. In general, the secondary-side liquid inventory mass is calculated in three steps:
- 1) Determine the temperature-compensated collapsed liquid level in the secondary side
[N i m:\np600%sec4\2344-414.non: I b- 100395 4,]41 REVIsICN: 1
- 2) Calculate the liquid volume from the collapsed liquid level
- 3) Calculate the liquid mass using the local thermodynamic properties of water determined from local pressure and fluid temperature measurements The pertinent data channels associated with the SG-1 secondary side are shown in Table 4.14-1, and the pertinent data channels associated with the SG-2 secondary side are shown in Table 4.14-2. In addition to the flow meters for the individual SGs, another flow meter provides information for the flows out of the secondary-side system. This flow meter is attached to the line that contains combined flow from the two SGs and the data channels associated with this combined flow meter are shown in Table 4.14-3.
Flow from the flow meter data channels (that is, FMM and FVM data channels) use a lower bound of 0.0. Negative values from these flow meters were not used. The geometry of the SGs is indirectly defined by Table 4.14-4, showing volume as a function of fluid height. In the following equations, this function is expressed as the function V(level,). Although narrow-range and wide-range data channels were available at the OSU test facility, the calculations assume that the wide-range channels are used. To avoid out-of-range errors, the fluid level readings are bounded to the values shown in Table 4.14-4. Here are no direct interfaces to other components that provide values to equations for this comporrnt. 4.14.2 Mass llalance Calculations 4.14.2.1 Steam Generator Secondary Components The equations listed below hold for SG-1 and SG-2. Only the channel names differ, as shown in Tables 14.4-1 and 14.4-2. The fluid mass stored in the SG secondary side is calculated from a wide-range level measurement of that volume. Like other level measurements, these readings must be density-compensated. The following approach is used in performing the density compensation of the SG secondary-side level measurements. He temperature channels are located at equal elevations on opposite sides of the component. For calculations and for LDP compensation, use the average of these two channels: T+T 2 T' = 4.14-3 2 O mwww.c4u344 414.noo:twoo395 4.14-2 REVISION: 1
e
.l 1,--
l = ('
. The density of the fluid in the SG secondary side is obtained from a standard steam table specific volume call: j l
. 4.14-4 i p, = VCL(P, , T,) i The specific heat of the fluid in the SG secondary side is obtained from a standard steam table call:
c,,, = CPL (P, , T,) 4.14-5 The mass of fluid is calculated: M, = p, x V(level,)' x C 4.14-6 i where: g V(level,) = Volume as a function of SG secondary side liquid, in.' t C = Conversion constant, in.' to ft.' The energy flow rate for this mass is defined: l AT' 4.14-7 Q, = M, x c,,, x ht , Since the facility also has flow meters associated with this component, additional calculations can be defiried for mass flows in and out of the component. The enthalpy of the fluid flowing into the SG secondary side is obtained from a standard steam table enthalpy call: hg, = HCL (P,, , Tg,) 4,14 8 m4*hc4u344-414.no :ib too395 4.14-3 REVISION: 1 J
l The density of the fluid flowing into the SG secondary side is obtained from a standard steam table specific volume call: l I 1.0 ; p g' = 4.14-9 ' VCL(Pg , , Tg,) j The mass flow rate into the SG secondary side is calculated: Nig, = W,, x p g, x C 4.14-10 where: W= Volumetric flow rate of liquid, gpm C = Conversion constant, gpm to ft.8/sec. The energy flow rate of the fluid flowing into the SG secondary side is defined: Qg, = hg , Nig, 4.14-11 Likewise, the enthalpy and the specific volume of the vapor flowing out of the SG secondary side are O obtained from a standard steam table enthalpy call: h our, Voer, = HSS(Pour,, Tour,) 4.14-12 The density of the vapor flowing out of the SG secondary side is defined: 1 4.14-13 Pour =oer," v .0 The mass flow rate out of the SG secondary side is calculated: Ki our, =Woer, x pour, x C i 4.14-14 9 m:v.couc4u144-414.o :isco395 4.14-4 REVISION: I
l
- (} sU ~ where: \m/
W:= Volumetric flow rate of steam, ft.8/ min. Ci= Conversion constant, min. to sec. Then, the change of the mass for the secondary side is defined: AM 4 so , g g' ,, g 9g 4.14-15 At s The energy flow rate of the vapor flowing out of the SG secondary side is defined: Qour, = hour,Se, 4*I4~I6 4.14.2.2 Steam Generator System The flow meter that is associated with the combined flows from both secondary-side components f, provides values that can be compared to the calculated masses from the other data channels.
~ N.) l A mass-weighted enthalpy at the site of this flow meter is calculated.
l (S our.s. sci xhour . sci) + (Socr. sm xhour. .. sm) 4,14 17 Sour.$ son + N our.s. sm l The specific volume of the vapor flowing out of the SG secondary side is obtained from a standard l steam table call: voer , = SSSISSS ( Pour., , hour ,) 4.14-18 m m The density of the vapor flowing out of the SG secondary side is defined; 1 l l 1.0 i Pour , " y 4.14-19 i our., l l [ s V m:We4u144-414.no : Stoo395 4,14 5 REVISION: I
The mass flow rate out of the SG secondary side system is calculated: 1 M our* =W Otbo
- Pour"" X C' 4.14-20 where:
Ci= Conversion constant, min. to sec. I 4 O t 4 i i mvmuc4c344-414.noa:tstoo395 4.14-6 REVISION: 1
l t
\ TABLE 4.14-1 INSTRUMENT CHANNEL ids FOR SG 1 SECONDARY SIDE MASS AND ENERGY CALCULATIONS Notation in this ChannelID Section Description I'r 301 P, Pressure of the fluid (psig)
TF-305 Ti Temperature (1) of the fluid (*F) TF-307 T Temperature (2) of the fluid ('F) LDP-301 level, Level of the fluid (wide range) (in.) FMM-001 w ,, Volumetric flow rate of the fluid into the component (gpm) TF-311 Temperature of the fluid into the component (*F) T,,g FT-001 p,,, Pressure of the fluid into the component (psig) FVM @l w,,' Volumetric flow rate of the vapor out of the component
- (ft.'/ min.)
4 I'F-301 Pressure of the vapor out of the component (psig) p Po ,. V i i i j . O
,G m:upumc4u344 414..on:isioo395 4.14-7 REVISION: 1
1 I TABLE 4.14-2 INSTRUhtENT CIIANNEL ids FOR SG 2 l SECONDARY SIDE h! ASS AND ENERGY CALCULATIONS Notation in this ! Channel ID Section Description FT-302 P, Pressure of the fluid (psig) TF-306 Ti Temperature (1) of the fluid ('F) TF-308 't .2 Temperature (2) of the fluid ('F) LDP-302 level, Level of the fluid (wide range) (in.) FMM-002 W, Volumetric flow rate of the fluid into the component (gpm) TF-312 Tg Temperature of the fluid into the component (*F) PT-001 P,, Pressure of the fluid into the component (psig) FVM-002 w,,r' Volumetric flow rate of the vapor out of the component Ct.'/ min.) I'F-302 pg Press,'re of the vapor out of the component (psig) l l l Ol m:Wawec4u)44-414 noo:ib-too395 4.14-8 REVISION: 1 1 1
\ 'V TABLE 4.14-3 INSTRUMENT CilANNEL ids FOR SG SYSTEM SECONDARY-SIDE MASS AND ENERGY CALCULATIONS Channel ID Notation in this Section Description FVM-003 w oc r_ Volumetric flow rate of the vapor (ft.hin.)
4 F'T-002 pg Pressure of the vapor (psig) TABLE 4.14-4 FLUID IIEIGilT VERSUS VOLUME FOR SG SECONDARY SIDE MASS AND ENERGY CALCULATIONS Secondary Side-1 Secondary Side-2 Fluid I,evel Volume Fluid Level Volume (in.) (in.') (in.) (in.') 0.0 963.0 0.0 963.0 i [ 46.5 9100.0 46.5 9106.0 ( 94.5 17,984.0 94.5 17,957.0 99.25 19,I48.0 99.5 I9,317.0
- i19.0 25,885.0 119.0 26,738.0 I
l I
'%./
m:\ap60lAsec4\2344 -414. coa : l b- 100395 4,]49 REVISION: 1
'4.15 Pressurizer V.
The pressurizer is a tank-like structure through which mass flows from the primary system to the first three. stages of the ADS valves under accident mitigation. At steady-state, the pressurizer has an initial liquid and steam volume. During a transient, the initial pressurizer and primary-side inventory are vented by the ADS through a pipe running from the top of the pressurizer to the IRWST. Subscript notation in this section is as follows:- f = Liquid phase of water g = Steam PRZR = Pressurizer SL-PRZR = Connection between the pressurizer and the surge line ADS 13 = Connection between the pressurizer and ADS 1-3 AMB = Ambient environment METAL = Pressurizer metal 4.15.1 Inputs and Assumptions
.( \
A general mass balance for the pressurizer may be expressed as: I raza , gst-eaza - S ADS 13 4.15-1 Similarly, a general energy balance on the pressurizer may be expressed as: d (MT ) C r dt 8 ""*"
~
^8' ~ " ^' ~ ^"'
The pertinent data channels associated with the pressurizer are shown in Table 4.15-1. In addition, the following values are provided by the ADS 1-3 module:
= Energy rate from pressurizer by way of ADS 1-3 (Btu /sec.)
.Q ,
Q = Mass flow rate through the ADS 1-3 (lbm/sec.) The geometry of the pressurizer is indirectly defined by Table 4.15-2, showing volume as a function
-f7 of fluid height. In the equations below, this function is expressed as V(level,).
b ,
' ma p60&uc4u344-415..on:tt>. loo 395 4.15-1 REVISION: 1
Table 4.15 3 provides the data channels and metal structure information needed for the pressurizer metal segments. In the calculation of the heat loss to the metal, the heat capacity of a segment is defined as a function of the metal temperature (heat _ cap (T,(j))). Table 4.15-4 provides this function. 4.15.2 Mass Balance Calculation The fluid level reading from LDP-601 needs to be compensated to account for temperature differences between fluid in the pressurizer and fluid in the reference leg of the instrument line. Since two temperature sensors are defined for the pressurizer, the temperatures are averaged for the fluid temperature: T, = T1 + T2 4.15-3 2.0 The density of the fluid in the pressurizer is obtained from an ASME steam table specific volume call: J 1.0 4.15-4 p, = VCL(P, , T,) The mass of the fluid is then calculated: Mr = p, x V(level,) x C 4.15-5 where: C = Conversion constant, in.' to ft.' The fluid specific heat, c p,,, is defined from a steam table call: c,,,= CPL ( P, , T,) 4.15-6 9 l I m:\arWXNec4\2344-415.noo:lb 100395 4.15-2 REVISION: 1 i
,-~>_ ! ' The enthalpy of the fluid from the surge line into the pressurizer is defined from a steam table call: h3ty,z,= M( P, , T,) 4. N
' The' void fraction of the pressurizer is defined:
V(level,) vf,=1.0 - 4.15 8 V, where V, is the maximum value provided by the fluid height versus volume data in Table 4.15-2. This void fraction is bounded: 0.0 5 vf, s 1.0 4.15-9 The vapor specific heat, c,,, and specific volume, y,, are defined from a steam table call: O v,, c,,,= CPV( P, , T,) 4.15-10 The vapor density is defined: p,= 4.15 11 8 The volume of the vapor in the pressurizer is defined: V,= Vm - V, 4.15-12 The mass of the vapor is calculated: M, = p, x V, x C 4.15-13 m:pec4u344-4ts on:n>.ioo395 4.15-3 REVISION: 1 l
The total mass (Ibm) in the pressurizer is defined; hi p ,2, = M, + M, ASM The rate of change in mass inventory of the pressurizer may be approximated by differencing two consecutive calculated values of the liquid and vapor masses: dhi,, paz, , M ,,,,z, , M ,, - M,,,, dt At t, - t, i 4.15-15 dM, ,,2, , Ahi,, ,,2, , M,,- M, dt At t, - t, ,, where the difference terms represent data from two consecutive calculations or data scans. The total rate of change of the mass inventory of the pressurizer is calculated: dM,,z, , AMp,2, , AM,, p,2, + AM . sPa2R 4.15-16 dt At At The mass rate of flow through ADS 13 is calculated in the ADS 1-3 module. The rate of flow from the surge line into the pressurizer may now be solved by rearranging Equation 4.15-1, and solving for the rate of flow from the surge line into the pressurizer: Nf,p,2, = " + M,33,3 4SU 3t
'Ihus, the mass flow into the pressurizer from the surge line is calculated using the pressurizer mass storage term and the total ADS mass flow rate, in addition, the enthalpy of the fluid in the line is provided by a steam tabic call:
h3t p,2 ,, = IICL ( P, , T, ) 4.15-18
. O maap60wecau34&415.aoa:ib.ioo395 4.15-4 REVISION: 1
h l p_ 4.15.3 Energy Italance The energy flow associated with the pressurizer consists of the following components: 4
- Rate of change in stored energy of the pressurizer inventory
- Rate of energy transport to the pressurizer from the surge line
- Rate of energy transport from the pressurizer by way of ADS 1 . . ' Rate of change in stored energy of the pressurizer metal
- Rate of energy loss to the environment Expressions for each of these components is developed in the following subsections.
( _ _4.15.3.1 Rate of Change of Energy in the Pressurizer Fluid Inventory 1
- The stored energy of the pressurizer inventory must account for both the liquid and vapor phases of the inventory:
i i d (M T ),u, , d (M, T, )_ d (M, T, ) Pun 4.15-19 P P. f P, g gg gg gt "Ihe equation for energy chenge associated with the liquid phase of the pressurizer '.sventory may be expressed: 1 1 d( M, (T,-T,,p) ),,,, c,, p
=
4.15 20 cp, , (T, - Tacp) d( M "" ) + cp, , M '"2" d( T, ) dt l l l where: T, ag = Reference temperature, 32*F 7% f 4-
%)
m:wwwec4u3*415.noa:isioo395 4.15 5 REVISION: 1 e
t l Similarly, the energy change associated with the vapor phase of die pressurizer inventory may be expressed as: d( M, (T,-Tay) ) , P, 3 g 4.15 21 d( M, ) d( T, ) c,, p U,-T,y) dt P. s g g Expressing the preceding two equations as differences: d( A1,,paz, U, 4 ,yD c,' ' ~ dt 4.15-22 cp, , (T,-Tagp) AM' " + c e, , M,, paz, AT and: g.PRZR g~ RE , CP,3 gt 4.15-23 AM 8
- AT "P. g 6 3-Tagp) + cp, , hi,, p,2, y 1
Expanding the two terms on the right-hand side of the preceding two equations, the liquid phase expression becomes: AM' ""2" M(, PRZR. i - M fPRZR el cp, , (T,-Tay) At
- g, q gg t, - t,,,
4.15-24 AT T -T H c e, , M,, pg7., = c ,,e M ,,p ,za,, , O manswxnsec4u344-4 5..ce:Ib-ioo395 4.15-6 REVISION: 1
(/p) x. and the vapor phase expression becomes: ce, , (T,-T,g,) AM*
= c , , (T,, ,-T,sp) .
p M* * -M"" 4.15-25 AT T -T ' Cr ,3 M 8- '"2" #' 8 M8-'"2E ' At - (3 where the subscript: i = ~ Inden of data and time arrays Combining the expressions above, the energy rate of the pressurizer is defined: Qeu,=c,,M, T- T'"' T- T'"' + e + c, ,M, t, - t,,, t, - t,,, 4.15-26 l M,'- M M, - M Cr .,(T,-T,,,) + p.g( g~
.(N L, i -t i-i REF i , 6-i j
l 4.15.3.2 Energy Transport Rate of ADS Flow The energy transpon associated with the flow through ADS 13 is calculated in the module for the ADS 1-3 separator. 4.15.3.3 Heat Loss to Metal he heat loss to the ambient surrounding from the pressurizer is calculated:
""' l Qw={ (heat _ cap (T,,fj)) x Mm(T,,/j)) x (T**,(j)'-T*,,(j)" ) 4.15-27 l 3.i t, - t,,,
where: nsurf = Number of metal segments, as defined by Table 4.15-3 OV De metal segment masses and heat caprities are defined by Table 4.15-4. mAap60thec4VM4-415.noa:Ibl00395 4,15 7 REVISION: l'
4.15.3.4 lleat Loss to Ambient In the calculation of the heat loss to the ambient surroundings, the pressurizer tank is modeled as having a number of segments, where the number of metal segments is defined as shown by Table 4.15-3. These divisions are chosen to coincide with the avai.lable temperature instrumentation. The heat loss for each metal segment is calculated below and then summed for the component. In the following calculations, the following constants are used: E = Emissivity (0.8) x = Insulation thickness (2 in.) k = Mean thermal conductivity of insulation (0.31 Btu-in/hr.-ft.2) C=i Conversion of 'F to degrees Rankine (459.6) CSB = Stefan-Boltzmann constant (1.713E-9 Btu /hr.-ft.2- R') C2 = Conversion of hours to seconds (3600.0) The heat loss to ambient heat transfer calculation methodology for the pressurizer is analogous to that employed for the other insulated components of the system. Each tank metal segment is treated as an insulated vertical cylinder. The calculations for the pressurizer energy loss to ambient are similar to those for the sumps, as presented in Subsection 4.9.5.5. This similarity is based on the fact that both components are insulated tanks, and both are modelled as vertical cylinders. The turbulent range form of the free convection heat transfer correlation from Reference 16 applies, based upon the conditions of the OSU test facility air at atmospheric pressure and ambient temperature (averaged over the tests). In the radiation heat transfer correlation, an emiselvity of 0.8 is assumed. The free convection heat transfer coefficient (11,,) for a metal segment j is defined: 11, = 0.09 x hT,fj)-T,,y, 4.15-28 The coefficient 0.09 in Equation 4.15-28 is obtained from the general form of heat transfer coefficient for alt: II, = 0.19 [ 0.10 AT l' = 0.09 [ AT ]t 4.15-29 O m:upcowc4u344-415 on:ib.too39s 4.15-8 REVISION: 1
.- - .- . - - _ - .- . -- . ~~
( ( The assumed insulation surface temperature (T,,) for a metal segment j is defined: T,,= (0,10 x (Tu, T,ua) + T,u,) + C, 4.15-30 l The use of the constant 0.10 la Equation 415-30 is based on the assumption that the temperature drop in the insulation is 90 percent'of the overall drop in temperature between the metal temperature and the ambient temperature. 1 The ambient temperature (T,) ('R) is defined: ; 1 l T, = Tau, + C, 4.15-31 l The radiant heat transfer coefficient (H,) , for a metal segment j is defined: H,,- CSB x E x (T,2+T,2) x (T, +T,) 4.15-32 b) V The surface resistance (Rg) for a metal segment j is defined:
= 1.0 Ra 4.15-33 (H,, + H,,) x A(j)
I The insulation resistance (R,(j)) for a metal segment j is defined:
- 4.15 34 R , = k x mean area (j)
/"%
4 h mwwmc4sn44-415..onsloo395 4.15-9 REVISION: 1
Given all of the above, the heat loss to the ambient (Btu /sec.) over all metal segments can be defined: T,,0)-T,u,
== f 4.15-35 R,4 +R,,
O AMB E J*l g2 4.15.3.5 Energy Transport Rate from the Surge Line The energy rate of surge line flow (Btu /sec.) is defined as:
~ ~
Ost-raza " Oraza - OADSI) metal ~ AMB ' O O m:Wecau344-415.non;1b-100395 4,13 10 ktX mN: 1
s (' 1 TABLE 4.15-1 INSTRUMENT CHANNEL IDS FOR PRESSURIZER MASS AND ENERGY CALCULATIONS .f l 1 1 Channel ID Notation in this Section . Description l 1 IT-604 - p, Pressure of the fluid (psig) i J IT-604 - P8 Pressure of the vapor (psig) I TF-605 Tl Temperature of the fluid ('F) 1 SC-608 72 Temperature of the fluid (*F) i 1 LDP401 level, Level of the fluid (in.) TF-602 T& Temperature of the vapor (*F) TF-006 T, Ambient temperature ('F) i i i l
- )
- i. ;
iO I 1 1 1 1 j 1 I l .f I i 1 1 I 1 ! l 4 i l i
\
b L m:Wec4u344-415.non:1b-ioo395 4.15-11 REVISION: 1 l
>-.s-n.,. ., ,.w , . . _ . , ,
T iBLE 4.15 2 FLUID IIEIGliT VERSUS VOLUME FOR l'RESSURIZER MASS CALCULATIONS Fluid Level Volume (in.) (in.') 0.0 0.0 4.625 537.1 12.5 688.7 24.5 2324.9 44.5 4634.9 M.5 7968.0 84.5 9326.8 1N.95 11549.1 TABLE 4.15-3 METAL DATA FOR PRESSURIZER METAL ENERGY CALCULATIONS Surface Mean Metal Temperature Metal Mass of segment Area Surface Area Segment Data Channel Mw 0) AU) mean_ area (J) Number j for T,,,(j) (Ibm) (ft.') (ft.') 1 TFM-6N 30.0 3.0 2.28 2 TFM-602 2551.0 17.41 15.25 3 TFM-605 2551.0 17.41 15.25 4 TFM-607 36.0 3.0 2.28 O maar60hc4u3nals.non:tb-too395 4.15-12 REVISION: 1
i
- l 1
. p. 1 TAllLE 4.15 4 TEMPERATURE VERSUS HEAT CAPACITY FOR PRESSURIZER METAL ENERGY CALCULATIONS Temperature Heat Capacity ('F) (Btu /lbm 'F) l 70.0 0.1085 100.0 0.1109 200.0 0.1175 l 300.0 0.1223 400.0 0.1256 500.0 0.1279 600.0 0.1297 O d 6 e j a i A m:uiswecau344-415.noa:1b ioo395 4.15-13 REVISION: 1 I
4 4.16 Pressurizer Surg Line
- De pressurizer surge line is piping that connects the pressurizer to the primary system. During a transient simulation where the ADS is actuated, the surge line becomes part of the relief path from the primary system to the pressurizer, ADS, and containment.
Subscript notation in this section is as follows: f- = Liquid phase of water g = Steam SL = Surge line HL-SL = Junction between the hot leg and the surge line SL-PRZR . = Connection between the pressurizer and the surge line ADS 13 = ADS 1-3 AMB = Ambient environment METAL = Pressurizer surge line metal A general mass balance on the pressurizer surge line may be expressed as: 8' = Mat3t - s .paz, 3t 4.16-1 4 Similarly, a general energy balance on the pressurizer surge line may be wTitten as: 4 cp 8' = (sh,)gt.3t - M ,)3tp,2, - Qug7,o - Q,y, 4M 4.16.1 Inputs and Assumptions i De pertinent data channels associated with the pressurizer surge line are shown in Table 4.16-1. In addition, the following values are provided by the pressurizer calculations: ! M st _paz, = Mass rate of junction between surge line and pressurizer (lbm/sec.) h 3t = Enthalpy of fluid in the junction between surge line and pressurizer (Bru/lbm)
.paz,,
He geometry of the pressurizer surge line is indirectly defined by Table 4.16-2, showing volume as a function of fluid height. In the equations below, this function is expressed as V(level,). He highest
-( volume value in this table is referenced in subsequent equations as (,
I arW4um416.aoe:151oo395 4.16 1 REVislON: 1
Table 4.16-3 provides the data channels and metal structure information needed for the pressurizer surge line metal segments. Only one metal segment was used to model the pressurizer surge line. In the calculation of the heat loss to the metal, the heat capacity of a segment is defined as a function of the metal temperature (heat _. cap (T.,fj))). Table 4.16-4 ; rovides this function. 4.16.2 Mass llalance The fluid level reading requires compensation because the sensor is outside the actual component. Since only one temperature sensor is defined for the pressurizer surge line, no temperature averaging is required for the LDP compensation. The density of the fluid in the pressurizer surge line is obtained from a standard steam table specific volume call: 1.0 4.16-3 p' = VCL(P, , T,)
'Ihe mass of the water in the pressurizer surge line is then calculated:
M, = p, x V(level,) x C 4.16-4 where: C = Conversion constant, in.' to ft.' The specific volume (v,) and the heat capacity (c,,,) of the gas are both obtained from a steam table call: cg ,, y, = CPV ( P, , T, ) 4.16-5 The density of the gas in the pressurizer surge line is: p, = 4.16-6 O m.\are00\su4u344-416.non:it> loo 395 4.16-2 REVISION: I
l i
.' l The volume of the gas in the pressurizer surge line is defined:
l V , = V , - V ,. 4.16-7 The mass of the gas in the pressurizer surge line is calculated: M, = p, x V, x C 4.16-8 ]J The total mass in the pressurizer surge line is defined: 4 Mst = M, + M, 4.16-9 i The mass flow rate of the pressurizer surge line from time step 1 1 to time step i is defined:
- dM st , M3t,- M841 4.16-10 dt t, - t,,,
, The mass flow rate from the surge line into the pressurizer is calculated from the mass balance on the j g pressurizer; the ADS 13 mass flow rate is measured and the rate of the fluid mass stored in or depleted from the pressurizer is calculated. From Equation 4.16-10, the rate that fluid inventory is either stored in or depleted from the surge line is calculated. Therefore, Equation 4.16-1 may be rearranged to' solve for the mass flow rate of the junction between the hot leg and the surge line:
s gust =M3vpaz,+ t
- Note that the flow from the hot leg into the surge line is calculated starting with the ADS 1-3 flow r
- and working in to the primary system to solve for the mass flow rates and fluid mass inventory changes, f
4.16.3 Energy Balance 4.16.3.1 Energy in Surge Line The heat capacity (c,,) of the fluid is obtained from an ASME steam table call: c,, = CPL ( P, , T,) 4.16-12 A b +- m:\ap600%sec4\2344-416.aoe:lb 100395 4,16-3 REVISION: 1 i
Assuming that the variation of specific heat of both the liquid and vapor masses of water over the range of pressure conditions experienced in the test is negligible, the left side of Equation 4.16-2 is expanded: d(M3t(T-Tay)) = Ost = c, dt dT dT8 4.16-13 c,ff, + c,.,M, + dM c,,,(T,-T,gp) dM' + c,,,(T,-T ag ,) Given the above equation, the energy rate of the pressurizer surge line is calculated: d(M3t(T-T,g,1) = Ost = c, dt T- T'"' T - T""i + 4.16-14 c,pf, + c, ,M, t, - t, ,, t, - t r3 c,,,(T,-Tagg) "' + cy (T,-Tag,) t, - t, ,, t, - t, ,, where: T,u = Reference temperature,32 *F 4.16.3.2 llent Loss to Surge Line Piping The energy rate of the metal is defined as a sum over all the segments of the surge line piping: mt T"'- T"-' 4.16-15 Omt= M,,(i) heat _ cap (T,') t,-t,3 4.16.3.3 IIcat Loss to Ambient Surroundings from Surge Line Piping in the calculation of the heat loss to the ambient surroundings, the pressurizer surge line is modeled as a horizontal segment, where the number of metal segments is defined as shown by Table 4.16-3. These divisions are chosen to coincide with the available temperature instrumentation. The heat loss is O maapmucm4&416.no :tb-too395 4.16-4 REVISION: 1
l i i l
,m l
calculated as shown below for the component. In the following calculations, the following constants ;
. are used: l l
.E. = ~ Emissivity (0.8' x- = Insulation thickness (2 in.) l k = Mean thermal conductivity of insulation (0.31 Btu in/hr ft.2)
Ci '= Conversion of 'F to degrees Rankine (459.6) d CSB = Stefan-Boltzmann constant (1.713E-9 Bru/hr-ft.2-R ) C2 = Conversion of hours to seconds (3600.0) The heat loss to ambient heat transfer calculation methodology for the pressurizer surge line is analogous to diat employed for the other insulated components of the system. Each metal segment is treated as an insulated horizontal cylinder. The calculations for the pressurizer surge line energy loss to ambient are similar to the cold-leg balance lines, as presented in Subsection 4.4.6. This similarity is based on the fact that both components are insulated pipes, and both are modelled as horizontal ! cylinders. I De laminar range form of the free convection heat transfer correlation from Reference 16 applies, ! based upon the conditions of the OSU test facility air at atmospheric pressure and ambient temperature
]
(averaged over the tests). In the derivation of the coefficient (0.18) in the convection heat transfer ! O \ correlation, an outside diameter of 6-in. is used. In the radiation heat transfer correlation, an emissivity ) of 0.8 is assumed. The free convection heat transfer coefficient (H,,) for a metal segment j is defined: Hg = 0.18 x hTJ)-T,y, 4.16-16 The assumed insulation surface temperature (Tg) for a metal s'egment j is defined: Tq= (0.10 x (Tg -TAwa) + TAus) + C, 4.16-17 I i The use of the constant 0.10 in Equation 4.1617 is based on die assumption that the temperature drop in the insulation is 90 percent of the overall drop in temperature between the metal temperature and the ambient temperature. O V m:Wec4u344-416.non:tsioo395 4.16-5 REVISION: 1
The ambient temperature (T,) is defined as follows: T , = T,u, + C, 4.16-18 The radiant heat transfer coefficient (H,,) for a metal segment j is defined: 11,= CSB x E x (T,2 + T,2) x (T,, + T,) 4.16-19 The surface resistance (R,,) for a metal segment j is defined: 1.0
= 4.16-20 R '"'i (11,, + H,,) x A(j)
The insulation resistance (R,(j)) for a metal segment j is defined: X 4.16-21 R"' =i k x mean_ area (j) Given all of the above, the heat loss to the ambient (Btu /sec.) over all metal segments can be defined: T,,,(J)-Tiy, ! n.wf 4.16-22 R,,,,+ R ,, QAMB" it g2 4.16.3.4 llot Leg / Surge Line Junction Assuming saturated vapor, the enthalpy of the gas in the junction area is provided by a steam table call: hgt.si,, = m ( Pge_st ) 4.WD l The enthalpy of the fluid in the junction area is provided by a steam table call: l hut-st., = IICL ( Pat , Tgt.3t ) 4.M-M O mAap600wc4uu4-416.non:ib loo 395 4.16-6 REVISION: 1
)
1he energy rate of the hot leg surge line area can be calculated: Ont-st " Ost + M -raza st h3t p,2,, + Q,g7,t + QA y, 4.43 O v E 1 i s 0 O Q n*pso,ecau 344-416.aoa : l t>. t oo395 4.16-7 REVISION: 1
TAllLE 4.161 INSTRUhlENT CIIANNEL IDS FOR PRESSURIZER SURGE LINE hlASS AND ENERGY CALCULATIONS Notation in Channel ID this Section Description IT 602 p, Pressure of the fluid (psig) TF-603 T, Temperature of the fluid (*F) FT-W P, Pressure of the gas (psig) TF-601 T, Temperature of the gas ('F) LDP-602 level, Level of the fluid (in.) IT-202 PgL ,,L N m em & h m h d k w g h W h kg m W O 1F-603 Twi,3t hymm a h Ndm M & wp h W Wq m m TAllLE 4.16-2 FLUID IIEIGIIT VERSUS VOLUhlE FOR PRESSURIZER SURGE LINE hlASS CALCULATIONS Fluid Level Volume (in.) (in.') 0.0 0.0 5.0 71.38 23.0 835.14 27.0 1327.01 28.5 1752.14 31.25 2266.26 32.325 2454.I1 35.124 2778.14 45.35 2879.24 48.48 2907.06 9 maagwenc4u344-416.non:1b-loo 395 4.16-8 REVISION: 1 l
f
\ TABLE 4.16 3 hilWAL DATA FOR PRESSURIZER SURGE LINE h!ETAL ENERGY CALCULATIONS Surface hiean hietal Temperature hletal Mass of Segment Area Surface Area Segment Data Channel h!,,w(j) A(j) mean_ area (j)
Number j for Tm /J) (ibm) (ft.2) (ft.2) 1 TFht-603 224 3 38.53 31.71 TABLE 4.16-4 TEhlPERATURE VERSUS IIEAT CAPACITY FOR PRESSURIZER SURGE LLNE h1ETAL ENERGY CALCULATIONS Temperature Heat capacity ('F) (Btu /lbm *F) 70.0 0.1085 100.0 0.1109 200.0 0.1175
\ 300.0 0.1223 400.0 0.1256 500.0 0.1279 600.0 0,1297 4
)
l l [ t s I mhp60tAiecau344 416.non:Ib-too395 4.16 9 REVISION' l
4.17 Cold Legs
'Ihe CMT piping design provides for the pressure balance lines for the CMTs to be connected to two of the four cold legs; specifically, the balance line piping is connected to CL-1 and CL-3. CL-2 and CL-4 do not have CMT pressure balance lines associated with them.
l 4.17.1 Cold Leg with Core Makeup Tank Balance Lines (CL-1 and CL-3) The break is located at CL-3; the break term applies only to CL 3. The mass balance on the cold legs with CMT balance lines, (CL-1 and CL-3), may be written as:
""' 4*
= Ni sa i
- Ni st - Moc.ct . at -M oax where the subscripts:
BL = CMT balance lines CL w BL = Cold legs with balance lines (CL-1 and CL 3) DC = Downcomer
; SG 1 = Steam generator-1 N
BRK = Break, applies to CL 3 for selected tests
'Ihe change in mass inventory in CL-1 and CL-3 given by the left-hand side of Equation 4.17-1 may be expanded to:
dM et . at dM et , dM a)
, 4.17 2 dt dt dt where the subscripts:
CL 1 = Cold leg-1 CL3 = Cold leg-3 4.17.1.1. Mass Terms Direct measurement of mass flow is provided in the primary system cold legs of the OSU test facility only during operation of the simulated reactor coolant pumps (RCPs) Overall primary system mass p balance calculations (Section 4.21) do not evaluate the cold leg as a separate component. Cold-leg
.V flow rate calculations are, therefore, limited to the initial period of pumped flow, mAap600sec4\2144 417.non:lb-100395 4,17 1 REVISION: 1 i
Prior to the start of testing and until the RCPs coast down, cold-leg Dow is a single-phase liquid and may, therefore, be measured with the magnetic Cow meters located in CL 1 and Ci,3. NI ct . ut.
" NI ntu-20: +NI ntu-203 where the subscripts:
FMM-201 = Instrument channel ids for magnetic Dow meters in CL 1 and CL-3, , FMM-203 respectively l i The single phase liquid flow rate is calculated as: Ni nty. " \Y atu- x p, (P3 ,y.m. Tn,y.m) x C 4 17-4 m m l where: l l l W = Volumetric flow rate of liquid, gpm C = Conversion constant, gpm to ft.3/sec. and the subscripts: FMM-XXX = data for flow meter XXX where: l XXX = 201 (CL 1) XXX = 203 (CL-3) f = Liquid phase of water The local liquid density is calculated from the ASME steam table routines using pressure, differential pressure, and temperature values from the data channels listed in Table 4.171. The local pressure used to evaluate local fluid density is calculated as: Patu- m = Per xxx + Por- m 4.17-5 l l where the subscripts: PT XXX = Pressure instrument per Table 4.171 DP YYY = Differential pressure instrument per Table 4.171
- O m:Wecau344-417.noo:Id.ioo395 4.17 2 REVISION: 1
Flow for tests with inoperative flow meters may be inferred from the cold. leg differential pressure cells. Test SB01 has functional flow meters and differential pressure cells and can be used to calibrate this method. The resultant single-phase liquid flow rate is calculated as: AP 4.17-6 gg , gj 2 ant-xxx nat-xxx ur-xxx gp t $ ur-xxx where: Ni ng.xxx = Average flow in Ibm /see through FMM XXX during the calibration period for Test SB01. The calibration penod is the period of full F power operation prior to the start of tM transient. AP3 m ,xxx
= Differential pressure corresponding to FMM-XXX per Table 4.17-1 APur-xxx = Average value of APai,y.xxx during the calibration period for Test SB01 O Once flow in the cold legs becomes two-phase, the magnitude of the output from the flow meters is not indicative of the actual flow in those components. 'Ihe liquid mass in CL-1 and CL-3 is then calculated using local level transducers as follows:
Step 1: Compensate the readings from the downcomer level transducer listed in Table 4.17-2 to account for temperature differences between fluid in the downcomer and fluid in the reference leg of the instrument line. The local pressure and temperature transducers to be used to accomplish the compensation I are also identified in Table 4.17 ' t Step 2: Tables of level versus volume were developed for the cold legs. From the I collapsed liquid level calculated from Step 1, calculate the liquid volume in the cold leg using linear interpolation: l V,, ci, x = f(LDP-YYYcos,,) 4.17-7 l i where the subscript: CL X = CL-1 or CL-3 I
' /^T t/
mw.c4u344 417.mab.ioons 4.17 3 REVISION: 1 I
The elevation of the cold leg relative to the downcomer LDP span extends from [ l in. Thus, the table of volume versus elevation for the cold legs is mapped onto the downcomer levels readings so that: Downcomer-Compensated Water Level Water Level in CL-1 and CL-3 (in.) (in.) [ ) b.e [ j=b.e [ j b.e [ ) b.e Step 3: The local pressures and temperatures measured using the instruments identified in Table 4.17 2 are used as inputs to the ASME steam tables to calculate the density of the liquid and vapor in the pressurizer in the cold legs: P r. ct x
= p(P,,r-xxx, T1,.xxx)
" P(Pi r-xxx, TTv-m) f P,. ct x where the subscripts:
l PT-XXX = Channel ID for local pressure measurement in downcomer TF-XXX = Channel ID for local temperature measurement of liquid temperatures TF ZZZ = Channel ID for local temperature measurement of vapor temperatures Step 4: Using the local thermodynamic properties of water as determined from local pressure and fluid temperature measurements, the liquid and vapor mass in the pressurizer is calculated as: i
- P r. ct x x V,, a x hI r. ct x hi s. et x Ps. ct x X (V Tm. a x - V .r ct x) where the subscript:
TOT = Total volume of component 9 m:Niramc4u344-4n.non:ib too395 4.17-4 REVISION: 1
i
- b. Step 5: The rate of change in mass inventory of the cold leg may be approximated by differencing two consecutive calculated values of the liquid and vapor masses:
dM,, ci, x , AM,, et x , M,, ct x, , . - M,, et x, ,., dt At t, - t, . , 4.17-10 dM,,ctx , AM s. ct x , M,, a x, , - M,, ct x, , , I dt At t, - t.,i where the subscript: 4 i = A specific value in the time and corresponding data array The total rate of change of the mass inventory of CL-1 or CL-3 is then calculated as: 2 dM ct , 3e AM ct et
! dt At 4.17 11 AM,, ct , + AM,, ct i + AM .r ct 3 + AM .s ct 3 l
2 4.17.1.2 Energy Terms I l The fluid in these two cold legs may, depending on the time and nature of the test, be in either a ! liquid or a vapor phase. Therefore, the change of energy for either CL 1 or CL-3 may be written as: l d( M ct x T) d( T, ) d( Mct x, , ) ! Cr dt "C.: r M ct x. , dt
+ c,, , T, dt 4.17 12 d( T ) d( M )
+C.:
r M c' ** 8 "' 8 8 dt dt J l l O l mwx4um 41ta:it,toons 4,17 5 REVISION: 1
l t Writing the previous equation in its difference form: d( M ct x T) AT, AM cL x, f P, gg C,f P CL x, f g C,f P f At 4.17-13 AT8 AM *8
+C,P M ct x, , + c,,, T, g ,
The data channel ids for the instroments to be used to define the thermal transport properdes are idendfled in Table 4.17-2.
%e rate of energy change of the metal mass for CL-1 and CL-3 is calculated as:
Oc., w BL kErAL QCL1NErE + Q CL3htETAL 4*l7 l4 expanding, Tw-20x 4,17 15 QnxkEra = Mxcxp At where the subscript: TFM 20X = Temperature-sensing element to be used for this calculation: X = 1 CL-1 X = 3 CL-3 Similarly, the heat flux from the surface of the cold leg is calculated as: Oct . at Asta Oct i Asia + Oct 3 Asta 4.17-16 he evld legs and CMT balance lines are similar in that they are predominantly horizontal pipes. The cold leg surface heat flux calculations and equadons are, therefore,idendcal to those of the CMT balance line. The balance line equadons are provided in Subsection 4.4.6 and the associated cold-leg instrumentadon is listed in Table 4,17 2, O m Np60 msec 4V 144-417.non: l b.100395 REVISION: 1 4.17 6
4 4.17.2 Cold Leg without Core Makeup Tank Halance Lines (Cold Leg 2 and Cold Leg-4) The mass balance on the cold legs without CMT balance lines (CL-2 and CL-4) may be written as: 4 dM a ,, ,t 4,37,37
, g1 ,, g1 g SG 2 DC, CL wk BL where the subscripts:
i CL w/o BL = Cold legs without balance lines (CL-2 and CL-4) DC = Downcomer SG 2 = Steam generator 2
'Ihe change in mass inventory in CL-2 and CL-4 given by the left hand side of Equation 4.17-17 j- above may be expanded to:
dM a we at dM ct dM cu
, 2
. 4.17-18 f dt dt dt i
!\ where the subscripts: 1 CL2 = Cold leg-2 l CL4 = Cold leg-4 l 1 l 4.17.2.1 Mass Terms , Direct measurement of mass flow is provided in the primary system cold legs of the OSU test facility only during operation of the simulated reactor coelant pumps (RCPs). Overall primary system mass balance calculations (Section 4.21) do not evaluat: the cold leg as a separate component. Cold leg flow rate calculations are, therefore, limited to trm laitial period of pumped flow. I Prior to the start of testing and until the RCPs coast down, cold leg flow is a single-phase liquid and e' may, therefore, be measured with the magnetic flow meters located in CL-2 and CL-4. Therefore: 4 bl ct we BL
- b atu-202 +b a m -204 '
O m:pec4u34&4itnon:tb.100395 4,17 7 REVISION: 1
1 where the subscripts: i FMM 202 = Instrument channel ids for magnetic flow meters in CL-2 and CL4, l FMM-204 respectively I All other parameters and subscripts are as previously defined. De single-phase liquid flow rate is calculated as: 1 S am-m " Wnu-m x p, (Pnm-m, 3T , m) x C 4.17-20 where: W = Volumetric flow rate of liquid, gpm C = Conversion constant, gpm to ft.8/sec. and the subscripts: FMM XXX = Data for flow meter XXX where XXX = 202 CL-2 204 CL-4 f = Liquid phase of water %c local liquid density is calculated from the ASME steam table routines using pressure, differential pressure, and temperature values from the data channels listed in Table 4.17 3 De local pressure used to evaluate local fluid density is calculated as: Pnu.m = PT-XXX + DP-YYY 4.17 21 where the subscripts: FTXXX = Pressure instrument per Table 4.17-3 DP-YYY = Differential pressure instrument per Table 4.17 3 O nWacau)44-417.aoe:tb-too395 4.17-8 REVISION: 1
l
\
r Flow for tests with inoperative flow meters may be inferred from the cold leg differential pressure cells. Test SB01 has functional flow meters and differential pressure cells and can be used to calibrate this m'e thod. The resultant single-phase liquid flow rate is calculated as: APemi-xxx 4.17 22 g4 , tmtaxx gy:sr-xxx a 3p
$ REF-xxx where:
1 I Ni ggy,xxx = Average flow in Ibm /see through FMM XXX during the calibration period for Test SB01. The calibration period is the period of full power operation prior to the start of the transient. AP,,xxx = Differential pressure corresponding to FMM XXX per Table 4.17 3 l AP,3,,xxx = Average value of APruu-xxx during the calibration period for Test SB01 A . Once flow in the cold legs becomes two-phase, the magnitude of the output from the flow meters is (] not Indicative of the actual flow in those components. The liquid mass in CL 2 and CL-4 is then ' calculated using local level transducers as follows: 1 1 Step 1: Compensate the readings from the downcomer level transducer listed in Table 4.17-4 I to account for temperature differences between fluid in the downcomer and the fluid in l the reference leg of the instrument line. The local pressure and temperature transducers to be used to accomplish the compensation are also identified in l Table 4.17-4. Step 2: Tables of level versus volume were developed for the cold legs. From the collapsed liquid level calculated from Step 1, calcillate the liquid volume in the pressurizer, V,,ci,x, as follows: V,, a x = f (LDP-YYYeos,p) 4.17 23 where the subscript: i CL X = CL 2 or CL-4 magewcau344-417.no.:n too39s 4.17 9 REVISION: 1
l l The elevation of the cold leg extends from [ ]** in. The table of vo ume versus elevation for CL 2 and CL-4 is mapped onto the downcomer levels readings, such tlut: Downcomer-Comnensated Water I.evel Water Level in CL-2 and CL-4 (in.) (in.) b., [ ja. [ j
; ja, g ja.
Step 3: 'The local pressures and temperatures measured using the instruments identified in Table 4.17 4 are used as inputs to the AShiE steam tables to calculate the density of the liquid and vapor in the pressurizer in the cold legs:
- P(Pn-xxx, TTr-xxx)
P r. ct x p ,, et x
= p(P n.xxx, TTv-zzz) where the subscripts:
IT-XXX = Channel ID for local pressure measurement in downcomer
'IF-XXX = Channel ID for local temperature measurement of liquid and vapor temperatures TFZZZ = Channel ID for local temperature measurement of vapor temperatures Step 4: Using the local thermodynamic properties of water as determined from local pressure and fluid temperature measurements, the liquid and vapor mass in the pressurizer is calculated as:
h!,, ci, x = p ,, ct x x V .r et x bl i, et x " - V,, et x) Ps. ct x x (Vior, et x where the subscript: TOT = Total volume of component 9 mAnn'aAsec4\2344 417.noe:lb.100395 4.17 10 REVISION: 1
O Q Step 5: The rate of change in .iiass inventory of the cold leg may be approximated by differencing two consecutive calculated values of the liquid and vapor masses: dM,, a x AM,, ct x , M,, et x, , - M,, ci, x, ,,, dt At ti - t .,i 4.17 26 dM,, et, x , AM,, ci, x , M,. ct x, , - M,, ct x, ,,, dt At t, - t i , where the subscript: i = A specific value in the time and corresponding data array The total rate of change of the mass inventory of CL-2 or CL-4 is then calculated as: dM et ,,, ,e AM ct ,,, et dt At 4.17 27 AM,, ct , + AMs. ct : + AM r. ct , + AM,, ct . At 1 1 4.17.2.2 anergy Terms ! The fluid it these two cold legs may, depending on the time and nature of the test, be in either a , liquid or a vapor phase. The change of energy for either CL 2 or CL-4 may be written as: 1 d( M et T) d( T, ) d(hict x, , ) ; cp x
=c,, e Met x. t '
dt dt i 4.17 28 l
- C * '
+ c.:r Mct x. :
dt +c.s Ta P dt l l 1 l m:www.e4um ait=:1b.100w5 4.17.I1 REVISION: I l l l
1 Writing the previous equation in its difference form: d( M co x T) AM et x, , cp = cp,, MCL x. t AT, + cp , , T, dt At At g,g AT AM
+
r C.: M'*8 C At "* 8 8 A The data channel ids for the instruments to be used to define the thermal transport properties are identified in Table 4.17-4. He rate of energy change of the metal mass for CL-2 and CL-4 is calculated as: 4.17-30 O CL w/o BL METAL OCL 2 MErAL + OCL 4 METAL expanding,
= Mxcx p 2cx 4.17-31 QCLx MErAL where the subscript:
O TFM-20X = Temperature-sensing element to be used for this calculation: X = 2 CL - 2 X = 4 CL - 4 Similarly, the heat flux from the surface of the cold leg is calculated as: QCL .io BL AMB QCL 2 AMB # Q CL 4 AMB 4.17-32 The cold legs and CMT balance lines are similar in that they are predominantly horizontal pipes. The cold-leg surface heat flux calculations and equations are, therefore, identical to those of the CMT balance lines. The balance line equations are provided in Subsection 4.4.6 and the associated cold-leg instrumentation is listed in Table 4.17-4. O maarwnsecau344-417.non: b-too395 4,17 12 REVISION: 1
v TABLE 4.171 DATA CIIANNEL ids USED TO CALCULATE LOCAL FLUID PROPERTIES FOR FLOW METERS Channel ids Data ChannelID for Data Channel ID for Data Channel ID Applicable Liquid Flow Meter for Differential Pressure for Conditions and DP Pressure Transducers Transducers Fluid Thermocouples Cold leg inventory is FMM 201 FT-101 DP-121 TF-107 all liquid DP-203 l FMM 203 IT 103 DP-123 TF-103 DP-205 TABLE 4.17 2 DATA CllANNEL ids USED TO CALCULATE FLUID PROPERTIES FOR LEVELS TRANSDUCERS Data ChannelID Data Channel Data Channel Data Channel for Energy Applicable ID for ID for ID for Balance Conditions Location Liquid Imels local Pressure Fluid Temperature Calculations Compensated water Downcomer LDP-140 TF-152 level when water annulus at PTIll TF-155 t T level in dowocomer - 180*az TF-158 V is above bottom of cold leg (= 270*az for backup) LDP-il6 Density and enthalpy CL-1 17-101 TF-107 (hquid) of hquid in cold legs SC-105 (vapor) CL-3 PT-103 TF-103 (liquid) SC 101 (vapor) Change in stored CL-1 TFM-201 energy of cold legs Heat loss to ambient CL-1 HFM 201 CL-3 HFM-203 Ambient temperature All CL TF-006 l l i O O maarwesec40344-417.non:Ib-100395 4.17-13 REVISION: 1
O TAllLE 4.17 3 DATA CilANNEL ids USED TO CALCULATE LOCAL FLUID PROPER ~1tES FOR FLOW METERS Channel ID Data ChannelID for Data Channel ID for Data Channel ID Applicable Liquid Flow Meter for Differential Pressure for Conditions and DP Pressure Transducers Transducers Fluid Thermocouples Cold leg inventory is FMM-202 I'r102 DP-122 TF-108 all liquid DP-202 FMM-204 I'T-104 DP-124 TF-IN DP-204 TABLE 4.17-4 DATA CIIANNEL ids USED TO CALCULATE FLUID PROPERTIES FOR LEVELS TRANSDUCERS Data Channel Data Channel ID ID Data ChannelID Data ChannelID for Energy Applicable for for for Balance Conditions Location Liquid Levels Local Pressure Fluid Temperature Calculations Compensated water Downcorner LDP-140 TF 152 level when water annulus at FT-I l l TF-155 } level in downcomer - 180*az TF 158 is above bottom of (= 270*az for LDP-il6 cold leg backup) Density and enthalpy CL-2 l'T-102 TF-108 (liquid) of liquid in cold legs SC-106 (vapor) CL-4 l'T-IN TF-IN (bquid) SC-102 (vapor) Change in stored CL-2 TFM-202 energy of cold legs , Heat loss to ambient CL-2 HFM 202 CL-4 HFM-2M Ambient temperaturr All CL TF-006 9 m:\ap60thsec4\2344-417.non: 1 h 100395 4,17 14 REVISION: 1
'O- 4.18 Hot Legs The OSU test facility incorporates the two hot legs of the AP600. Each hot leg has an ADS-4 line associated with it. In addition, one hot leg, designated as HL-2, also has the pressurizer surge line and the PRHR feed line associate with it. The pressurizer surge line connection to HL-2 provides a flow path from the primary system to ADS 1-3. Considering the various flow paths, the mass balance for HL-1 and HL-2, respectively, may be written as:
dM ' " M cP-ilt i
-b ADS 4-1 ~
SG I dt and dM * = S ur-ta. 2 -M SL -b ADS 4-2 ~ PRmt ~ SG 2 dt where the subscripts: ADS 4-X = ADS-4 X = 1, HL-1 q = 2, HL-2 b IIL 1, HL 2 = HL-1 or HL-2, respectively SG 1, SG 2 = Steam generator-1 or steam generator-2, respectively SL = Junction between HL-2 and the pressurizer surge line UP - HL 1, = Junction between the RPV upper plenum and HL-1 UP - HL 2 = Junction bewteen the RPV upper plenum and HL-2 There is no direct measurement of mass flow in the primary system hot legs of the OSU test facility; hot leg flows may be inferred by coatinuity and/or energy balances performed on the system. Overall primary system mass balance calculations (Section 4.21) do not evaluate the hot leg as a separate component. Ilot-leg flow rates are, therefore, not calculated. Hot-leg calculations include fluid inventory, the rate of change in fluid and metal energy, and heat loss to ambient. O. m:\ap600Wec4\2344 418.noo:Ib100395 4,jg.1 REVISION: 1
i l l 4.18.1 Mass Storage in the flot Legs The hot leg may be in one of three states; all liquid, all vapor, or a mixture of liquid and vapor. The output from the appropriate level transducers listed in Table 4.18-1 is used to first calculate the liquid inventory and mass in the hot leg, and then corresponding vapor inventory and mass. This calculation is accomplished as follows: Step 1: Compensate the readings from HL-1 level transducers licted in Table 4.18-1 to account for temperature differences between fluid in the hot leg and fluid in the reference leg of the instrument line. The instruments used to measure local pressure and fluid temperatures to be used to accomplish the compensation are also identified in Table 4.18-1. Step 2: The local pressures and temperatures measured by the instruments identified in Table 4.18-1 are used as inputs to the ASME steam tables to calculate the density of the fluid in the hot legs: Pr. a x
= p, (P,T) 4.18-3 p,, g x = r (P,T) where the subscripts:
f = Liquid phase O g = Steam Step 3: Tables of level versus volume were developed for the two hot legs. Using the compensated liquid level and the table of hot-leg volume as a function of height, the volume,in ft.3, of liquid in the hot leg is determined using linear interpolation: V,, a i = f(LDP-XXXcoup) 4. M Step 4: 'Ihe liquid mass inventory in the hot leg is calculated as: M,, y x = p ,, a x x V,, y x 4.18-5 and the vapor mass is calculated as: M,, a x = p ,, y x x (Vror, a x - V,, y x) 4.18-6 O m:bp600eec4'i2344-4 l 8.non:1 tw 100395 4,Ig.2 REVISION: 1
_. _ _= . . ~ where the subscript: g- = Vapor conditions at the local hot-leg pressures and temperatures
. TOT- = Total, or liquid + vapor
, he total fluid mass in the hot leg is calculated as: M gx = M,, a x + M,, a x 4.18 1 i' Step 5: ne rate of change in mass inventory of the hot leg may be approximated by differencing i two consecutive calculated values ofliquid mass: ) 1 l i 1 dM gx AM gx Ma x, , - Mg x, ,_,
, , 4.18-8 4 dt At t, - t i.,
where the subscript: 1 = Index of the data and time arrays s 4.18.2 Energy Terms The fluid in the two hot legs may, depending upon the time and nature of the test, be either in a liquid i or a vapor phase. Thus, the change of energy for the fluid in either HL-1 or HL-2 may be written as: i d( M gx T) d( T, ) d( Ma x, , ) cp = ce, , M,,t x, , + ce, , T, dt dt dt 4.18-9 d( T ) d( M )
*C.:r M ut x. : dt "' ' d 4
i where the subscript: g = Steam f = Liquid phase of water HL X = X = 1 for HL 1 X = 2 for HL 2
- A. ;
maniamc4u3n41s.nno:tb-too395 4.18-3 REVISION: 1
l l Writing the above equation in its difference form: d( Mm. x T) ~ AT, + M .mx. f P g P. f lu. X f g P. f f g 4.18-10 AT AM *8 ce, , M,,o x, , y + c ,p , T, _ Re data channel ids for the instruments to be used to define the thermal transport properties are identified in Table 4.18-1. He rate of energy change of the metal mass for IIL-1 and IIL-2 is calculated as: AT
- awl Qn.xurra i
= M urrn xc p x where the subscript:
TFM 20Y Designates the temperature sensing element to be used for this calculation: where Y = 5 => HL-1 6 => HL-2 Re hot legs and CMT balance lines are similar in that they are predominantly horizontal pires. The hot-leg surface heat flux calculations and equations are, therefore, identical to those of the CMT balance line. The balance line equations are provided in Subsection 4.4.6 and the associated hot-leg instrumentation is provided in Table 4.18-1. O m:\apsoc%ec4u 344-4 t s.noa:151oo395 4.18-4 REVISION: 1
1 l O TABLE 4.18-1 DATA CIIANNEL ids USED IN IIOT LEG MASS AND ENERGY CALCULATIONS Data Channel Data Channel ID Data Channel Data Channel ID for ID ID for Energy Applicable Isr for Fluid Balance Conditions Location Liquid Levels Local Pressure Temperature Calculations Fluid conditions in HL-1 HPS-2050) - hot leg HL-2 HPS-206"> % Compensated levels HL-1 LDP-207 frr-202 TF-205 readings in hot legs HL-2 LDP-208 Irr-202 TF-206 Density and IIL-1 % frr-202 TF-143 (liquid) , enthalpy of liquid in SC-141 (vapor) HL-2 Irr ri TF-142 (liquid) SC-140 (vapor)
^
Change in stored HL-1 TFM-205 I l energy of hot legs l HL-2 TFM 206 j O ifcat loss to ambient IIL-1 HFM-205 i kb , HL-2 HFM 206 Ambient HL-1 TF-006 HL-2 temperature Note: (1) Heated phase switch is used to determine state of fluid in hot legs: liquid solid = 100, full scale reading ) two-phase = 0 < x < 100 vapor solid = 0 l t v m:WMhec44344 418.noo:1tr100'95 4,]g.5 REVISION: 1
V - 4.19 Pressure Conversions
' A number ofinput pressures need to be converted from psig to psia and placed onto the results plot file. As with other calculations involving input pressures in the various modules, the ambient pressure i: ; data channel (PT-003) is used for the conversion, as opposed to a constant of 14.7 psi. The units of the resulting pressure values is psia. A list of pressures converted from psig to psia is given in Table 4,19-1, i
t l: d t 4 4 O ,b i 4 f 4 e i d l m:Wahc4u344-419.non:15too395 4.19-1 REVISION: 1 e
}
TABLE 4.191 PRESSURE CONVERSIONS Data Channel Channel Description DP-905 Break Separ Ent DP PT-001 MFP Dischrg Pressure PT-002 MS Header Pressure PT-101 CL1 Press @RV Flange PT-102 CL2 Press @RV Flange FT-103 CL3 Press @RV Flange PT-104 CL4 Press @RV Flange IYT-107 RV Upper Head Press PT-108 RV Bottom Pressure PT-109 DVil Pres @RV Flange PT-110 DV12 Pres @RV Flange PT-111 RV Dwncmer Press-Top PT-112 RV Dwncmer Press-Bot I'l 113 Retr Blw Mid Spc Grd PT-201 SG1 Long Tube Press irr-202 IfL-2 Pressure
~
PT-203 PR Upstrm of Break-1 PT-204 SG2 Long Tube Press PT-205 HL-1 Pressure PT-206 PR Upstrm of Break-2 PT-301 SG1 Sec Steam Press PT-302 SG2 Sec Steam Press PT-401 ACC-1 Pressure PT-402 ACC-2 Pressure PT-501 CMT-1 Pressure PT-502 CMT-2 Pressure m:Wec4u344-419.noa:its too395 4,19 2 REVISION: 1
l l l O TABLE 4.191 (Continued) PRESSURE CONVERSIONS Data Channel Channel Description FT-602 PZR NR Pressure f*T-603 PZR NR Pressure
=
PT-604 PZR WR Pressure FT-605 ADSl-3 Separtr Press IrT-606 IRWST Sparger Press , I'T-610 ADS 4-2 Separtr Press PT-611 ADS 4-1 Separtr Press . PT-701 IRWST Pressure PT-801 CVSP Discharge Press PT-802 RNSP Discharge Press I'T-901 Primary Sump Press I'T-902 BAMS Header Pressure PT-905 Break Separatr Press l l l l 4 l l i m%WDsec4V344-4l9 noa:Itw100395 4,19-3 REVISION: 1
l l s
]
' ) ..A i
cated. The impact of the downcomer liquid on the solid surface of the core barrel flange was heard during the test.
During accumulator injection, PRIIR flow decreased to near [ ]'b' and the PRIIR level decreased. He majority of the nitrogen was discharged into the DVI lines at about [
J'A' seconds, respectively with small subsequent periodic outflows marked by spikes in the DVI line flow The PRilR IIX iniet fluid temperature became subcooled while the ADS-4 valves opened at
[ ]"' seconds and over the next [ l'** seconds dropped to and paralleled the PRIIR llX outlet fluid temperature. Again, this is an indication that there was no flow through the PRilR IIX during this time frame. From [ ]'** seconds, both DVI nozzle temperatures increased from essentially ambient conditions to as high as [ l'A* *F and then returned to ambient condition. De DVI fluid temperature increase was caused by two factors. First, there was rapid heating of the remaining water injected from the CMTs. Second, the reactor vessel downcomer level was at the DVI nozzle level during this period, possibly partially uncovering the nozzles. De DVI fluid temperature transient was terminated when IRWST injection began refilling the reactor vessel at about [ ]'** seconds, and temperatures returned to ambient conditions when the CMTs were empty of hot liquid.
Steam generation in the reactor vessel was reestablished at about [ l*** seconds and the reactor vessel inventory began to decrease. At [ l'** seconds, the ADS 4-1 and ADS 4-2 valves opened automatically when the CMT-1 level reached its low-low level setpoint. ADS-4 actuation started a decline in RCS inventory that could not be overcome until IRWST injection began. There was initially maamw44 50.non:Ib ioo395 5.1.1-4 REVISION: 1
an excess of mass in the system to be vented through ADS-4 before IRWST injecdon could occur.
; CMT 1 and CMT-2 were completely empty at [ ]** seconds, respectively.
The collapsed reactor vessel level reached a minimum value of about [ ]** in. at [ ]** seconds. Although this level is below the top of the heater rod heated length, the actual level of the top of the two-phase mixture is much higher, as shown in the behavior of the core outlet thermocouples (Figure 5.1.1-4), which do not exceed saturation temperatures. At about [ ]** seconds, the RCS system pressure decreased to about [ ]** psig, which was sufficiendy low that the IRWST static head was greater than RCS pressure, and IRWST injection began. 5.1.1.4 In Containment Refueling Water Storage Tank injection Phase
~ IRWST injecdon was split between the two DVI lines beginning at [ ]** seconds and continually diminished (Figure 5.1.12) as the differential head between the IRWST and the RCS decreased with draining of the IRWST. IRWST injection was sufficient for the primary system to refill. De ;
pressurizer and pressurizer surge line emptied a second time at about [ ]*' seconds, respectively. O When the pressurizer had a liquid level, ADS 13 separator and sparger pressure became negadve by as much as [ ]** psig. Rese values remained negative from [ ]** seconds as steam in the ; ADS lines was condensed in the IRWST. The negative pressure was broken as the level in the IRWST l decreased below the sparger nozzles. No vacuum breaker was installed on the sparger line inside the l I IRWST for this test. A vacuum breaker is included in the AP600 plant design, thus this OSU test response may not be typical of the AP600. The surge line then began to reflood almost immediately at [ ]** seconds and the pressurizer at about [ ]** seconds. De reflood was caused by RCS levels increasing above the reactor vessel nozzles because IRWST injection exceeded the inventory losses, and by the condensadon in the ADS 1-3 lines. The maximum pressurizer level attained was , about [ ]** In at [ ]** seconds, but it immediately began to decrease. The pressurizer was empty at [ ]*' seconds and remained empty for the remainder of the test. The surge line stayed full until [ ]^* seconds when the level decreased to about [ ]** In. and remained there until [ ]** seconds. The level again decreased to about [ ]**In.at[ ]** seconds and oscillated between about [ ]** in. for the remainder of the test. Both CMT balance lines began to refill at about [ ]** seconds when the IRWST injection increased the reactor vessel level sufficiently to begin covering and refilling the cold legs. At about [ ]** seconds, when the CMT-2 balance line had completely refilled, CMT 2 began to rapidly refill and reached the [- ]** in. level (about two-thirds full) at about [ ]** seconds. He CMT refill is discussed in more detail in Subsection 6.1.1. After the CMTs were panially refilled, there was no injection flow from the CMTs because the higher stade head of the IRWST held the CMT discharge ,i line check valves closed.
. ma p6coamso.noo:isioons 5.1.1-5 REVISION: 1 i
Steam generation started again at about [ ]*' seconds and continued for the remainder of the test (Figure 5.1.13). CMT 1 and CMT-2 remained at essentially constant levels for several thousand seconds and then began slow draindowns at about [ ]** seconds, respectively. The draindown for both CMTs was slow and did not occur until the IRWST relative level was [
]*' in. below that of the Chits. Data indicate that the CMTs drained for a while, and then the differential head between the IRWST and the CMTs again closed the CMT discharge check valves, terminating draining until the differential shifted the other way and draining recommenced. Both CMTs were completely empty at about [ ]** seconds, which coincides closely with the primary sump injection valve opening at [ ]** seconds. A possible correlation is that when the primary sump valve opened, the IRWST had just reached its minimum level of about [ ]** In., which is about
[ ]** In. below the instrumented level for the CMTs, and that there was still a slight negative pressure remt.ining in the CMTs. Also, when the primary sump injection valves opened, there was a short period in which the IRWST and primary sump levels equalized, causing a decrease in RCS fluid levels and resulting in a rapid drop in CMT levels from about [ ]** in. Starting at about [ ]** seconds, there was a series of pressure, level, and flow oscillations that occurred throughout the components of the facility lasting until about [ ]** seconds. These oscillations will be discussed in detail in Subsection 6.1.3. At[ ]** seconds, the PRilR llX inlet fluid temperature instantly increased from [ ]** 'F to saturation temperature. The temperature increased at about [ ]** seconds after pressure, level, and flow oscillations began in the facility and was possibly caused by the inlet line burping, which once again allowed the line to fill with saturated steam. Following the burp, all of the PRHR IIX temperatures began to slowly approach saturation. The break separator level began to increase at the same rate as the primary sump at about [ ]** seconds. This increase occurred when the sump level reached the height of the break separator k)op seal. As a result of this increase, the break separator level reached the height of the break in CL-3, causing break flow to reverse and flow from the break separator into the RCS through the break at about [ ]** seconds. Break flow then remained essentially zero or slightly negative 1 throughout the rest of the test. 5.1.1.5 Sump Injection Phase Primary sump injection (Figure 5.1.1-2) began through the check valves around the sump injection valves at about [ ]*' seconds, when primary sump and IRWST levels were essentially equal. At [ ]** seconds, the primary sump injection valves automatically opened when the IRWST reached its low low level setpoint of [ 1** In. In the LTC mode of operation, primary system inventory was lost through the ADS 4-1 and ADS 4 2 valves to the primary sump. System inventory was made up through primary sump and IRWST injection through the DVIlines and some small flow from the primary sump through the break mwm. 50. ib.ioo395 5.1.1 6 REVISION: 1
O V separator and into the break. - The driving force for this flow was the difference between die liquid head in the downcomer and the corresponding head in the simulated core. 'Ihe two-phase mixture produced in the core flowed out through the hot legs and ADS-4 to the primary sump, and die somewhat cooler fluid in the sump returned from the bottom of the sump to the reactor vessel downcomer via the DVI lines. When sump injection began, the reactor vessel downcomer fluid temperatures rapidly increased to the
. sump flow injection temperature. The core steam generation increased (Figure 5.1.13) due to injection of hotter sump fluid. When the primary sump injection valves opened, the DVI flow decreased, and the sump and IRWST levels equalized. During the reduced DVI flow period, there was an upward spike in reactor downcomer temperatures. The upper downcomer fluid temperature, indicated by thermocouples located above the DVI nozzle elevation, increased to saturation at this time and remained at saturation for the rest of the test. Figure 5.1.13 shows a corresponding increase in steam generation rate.
Overflow from the primary sump to the secondary sump started at about [ l*' seconds. When primary sump injection started through the check valves, flowing around the sump injection valve lines, IRWST injection from each line was about [ l"' lbm/sec. and another [ -) 'lbm/sec. from each sump line to each DVI line. With the opening of the primary sump valves [ ]"' seconds, injection flow increased to approximately [ ]"' lbm/sec. through each DVI line. About one-third of the flow from sump-l was diverted back into the IRWST with the remainder flowing to the DVI nozzle (k due to the smaller pressure drop in the IRWST to sump line (2.5-in diameter versus 1.5 in. diameter for sump 2). One fourth of DVI 2 flow was provided by flow from the IRWST and the remainder from the sump 2 line. The PRHR HX outlet temperature remained subcooled in the range of [ l'"' 'F during most of the test, but after sump injection, it began to rise and was just reaching saturation temperature at the end of the test. The PRHR HX was inactive during this phase, since the IRWST had drained. 4 O mAap600\2344w 50.noo:Ib-100395 5.1.1 7 REVISION: 1 a
TAllt.E 5.1.11 O OSU TINT ANALYSIS PLOT PACKAGE FOR SUllSECflON 5.1.1 Plot No. Component Variables Units Description i Pressurizer CI'T4M psia System pressure and event history 2 Water WWTDVil+WWTDVI2, Ibm /sec. Total of CMT, accumulator, IRWST, injection WOUTACCl+WOLTTACC2, and sump injecdon flows WWTIRWil+WWTIRW12, WWTSMP!T 3 Reactor RPVASOU2 lbm/sec. Steam generation in reactor vessel vessel 4 Reactor 1V8RPV, IITMXRPV, TSAT *F Rextor vessel outlet temperature, vessel maximum clad temperature and fuel exit saturation temperature O l O mh twwn4w.50.noa:tb loons 5.1,l.8 REVISION: I
O TIIE FIGURES LISTED IN TABLE 5.1.1 1 ARE NOT INCLUDED IN Tills NONPROPRIETARY DOCUMENT I I b l 1 1 I m:Wan2%.so noa:tb-wo395 5.1.1 9 REVISION: I l { 1 J
f k 5.1.2 Short Term Transient For the 2 in., cold leg break, Matrix Test SB01, the short-term tramient covered the first [ ]** seconds. As can be seen from Figures 5.1.21 and 5.1.2 2, this period included full depressurization of . the facility through all four stages of the ADS together with CMT and accumulator injection plus the inidal stages of IRWST injection. The mass and energy distribution for this phase of the transient is discussed here based on the plot package detailed in Table 5.1.21. These plots concentrate on die primary system including the accumulators, CMTs, IRWST, sumps, and flow from the primary system via the ADS, break, and IRWST overCow. 5.1.2.1 Maintenance of Core Cooling Reactor Pressure Vessel and Downcomer Mass Distribution For the short term transient, the most important criteria was the maintenance of sufficient core inventory to supply adequate cooling of the heater rods. Figure 5.1.2 57 slows that there were no significant excursions in heater rod temperatures and, therefore, sufficient core inventory was maintained through I this phase of the transient to remove the decay heat from the rods, flowever, for significant portions of the transient a two-phase mixture is present in the core and upper plenum regions. The following i discussion tracks the variation in water level and mass througlout the reactor vessel and downcomer. O The total fluid mass in the reactor pressure vessel (RPV) is shown in Figure 5.1.240. liere, and throughout the report the overall RPV inventory is described in terms of the eight regions shown in Figure 4.11 1 plus a separate downcomer region. The hdtlal RPV inventory is [ ]** lbm. During j the course of the short term transient, the vessel inventory experiences two mass minimums: [ ]** lbm before accumulator injection and [ ]*'lbm before IRWST injection. Steam generation was near i maximum at these times (Figure 5.1.2 55). By the end of the short term transient, vessel inventory I recovered to a steady [ ]** lbm. Sindlar variations were seen in the core fluid mass and water level i- shown in Figures 5 l.2-44 and 5.1.2-45. The minimum core level occurred before IRWST injection. It . . can be seen from Figure 5.1.2 45 that during this phase of the transient, the collapsed liquid level l dropped to [ ]** in, below the top of the heated rod length. By the end of the short term transient, the
]
effect of IRWST injection ended all core boiling (Figure 5.1.2 55), and the core was again water solid. , 1 it was noticed in the analysis of the SDLOCA simulations on the SPES 2 facility
- that following the I pump trip there were short period oscillations in primary system flow, temperature, and pressure. A small number of oscillations were also observed in the OSU test response after the end of the initial blowdown (see, for example, the pressure response in Figure 5.1.21 and the reactor vessel mass in Figure 5.1.2 40), in the SPES 2 tests, the oscillations were clearly driven by power to-flow mismatches in the core due to high core power levels, which were needed in SPES 2 to compensate for Idgh ambient losses. The short term oscillations observed in the OSU results are believed to be a result of ;
p pressure oscillations following the initial blowdown. 4 U l m%wmmso.amtb.toons 5.1.21 REVISION: 1
i The Guld mass in the core region is shown in Figure 5.1.2-44. Once again, the two minimums occur prior to accumulator injection and prior to IRWST injection. De ndnimum core inventory is [ l** lbm. He collapsed liquid level in the upper plenum region spanned by LDP-138 and the associated fluid mass are shown in Figures 5.1.2-49 and 5.1.2-48. It can be seen that, during the period before accumulator injection, the collapsed liquid level in the upper plenum dropped below the hot leg elevation. During accumulator injection, the steam bubble in the upper plenum partially condensed and the water level briefly rose above the hot legs. Following the end of accumulator injection, flow from the CMTs was not sufficient to maintain the upper plenum level, and the region of LDP 113 completely drained of water. IRWST injection caused the upper plenum to refill, and this region became water-solid again at approximately [ ] seconds. He fluid mass and collapsed liquid level for the head region are given in Figures 5.1.2 50 and 5.1.2-51. During the first [ l'* seconds, the head inventory reaches a minimum of [ ] lbm of water. Accumulator injection was sufficient to maintain a level in the head region. At the end of accumulator injection and before IRWST injection began, the upper head drained. Once IRWST injection began, liquid level in the upper head was reestablished. De mass of fluid and collapsed liquid level in the reactor vessel downcomer are shown in Figures 5.l.2-41 and 5.1.2-42. During the blowdown phase of the transient, the level dropped to the elevation of the cold legs. Tlds elevation was maintained undl the cold legs were fully drained. Following this time, the collapsed level remained between the DVI and hot leg elevations until IRWST injection once again raised the level above the cold legs and cold leg refill commenced. l oop Mass Distribution For th s discussion the loop was considered to consist of the hot and cold-leg pipe work, the SG primary side, and the pressurizer plus surge line. De total fluid mass and water level for the pressurizer are shown in Figures 5.1.2 34 and 5.1.2-35. During the blowdown phase of the transient ([ l'* seconds), the pressurizer drained rapidly, becoming completely empty of water at about [ ]'* seconds. De pressurizer remained empty until ADS-1 actuation at [ ]'* seconds. At this time, water was drawn back into the pressurizer as steam and water flowed out of the ADS. A fluid inventory of over [ l'* lbm was maintained until ADS-4 actuation at [ l'* seconds. His caused an initial outsurge through the surge line, followed by a more gradual draining of the pressurizer as mass flowed out of the hot legs via the ADS 4 valves. The pressurizer fully drained at [ l'* seconds and remained empty for the remainder of the transient. Mass data for the SG U tubes and their associated inlet and outlet plena are shown in Figures 5.1.2 32 and 5.1.2 31. %e SG tubes gradually drained until ADS actuation when all the tubes and plena were empty of water. SG 1 on the broken loop drained before SG 2. Any flow tiuough the SGs ceased mvouch50.noo:Ib.100N5 5.1.2 2 REVISION: 1 I l
7 i once the tubes drained and steam trapped the U tubes became superheated. Once the SG tubes drained,
- l. natural circulation around the primary loop circuit ceased. He SG U-tubes remained empty for the remainder of the short term transient.
He mass of water and vapor in the hot legs a c shown in Figures 5.1.2 58 and 5.1.2 59. He water j mass calculated for HL 1 is not considered valid for this test as it did not indicate the appropriate level of draining. He hot legs maintained their water inventory until [ ]** seconds into the transient when they started to drain (Figure 5.1.2-58). The hot legs completely drained within [ ]*' seconds. i Actuation of ADS 1 caused a rapid increase in void fraction in both hot legs. A larger void fraction was maintained in HL-1 as steam was preferentially removed from IIL 2 by PRHR. Figure 5.1.2 59 l reveals that, at around [ ]** seconds, the small amot.nt of steam remaining in HL-2 was removed, i and the water level in that hot leg increased as a result of condensation in the PRHR HX drawing a vacuum and raising the level on loop 2. De liquid and vapor mass for the four cold legs are shown in Figures 5.1.2-60 and 5.1.2 6L
- Following initial blowdown, all four cold legs became two-phase, although there was a greater void j fractic'1 in CL 1 anxi CL-4 compared with CL-2 and CL-3 (Figure 5.1.2-61). The mass variation in all
- four cold legs appears very similar. De mass variations were derived from levels in the reactor vessel
- downcomer because the level instmments on the cold legs are unreliable. All the cold legs were L completely dralned after ADS actuadon and refilled at [ ]** seconds when flow from the IRWST refilled the reactor vessel downcomer to the level of the cold legs (Figure 5,1.2-42). The cold legs did
not drain uniformly, but rather, CL 3 (with the break) drained first, followed by CL 1 with CL-2 and CL-4 oclayed. Figure 5.1.2 31 shows that the SG oudet plenum ou loop 1 drained at about
[ ]** seconds before that on loop 2, confirming the expected asymmetry in cold-leg behavior, i Mass Isdected to the Primary System
- De CMTs trrnsitioned from a recirculation to a mass injection mode at approximately [ ]^* seconds 4 when the cold leg started to drain. Draindown of the CMTs continued until the CMT check valves I were closed by flow from the accumulators. CMT draindown restarted at the end of accumulator injection, continuing until IRWST injection began (Figures 5.1.2 5 and 5.1.2 6).
J The accumulators drained about [ ]** seconds before activation of the ADS. He accumulators started discharging into the DVI line when the system pressure dropped below the pressure in each 4 accumulator. Accumulator injection began at approximately [ ]** seconds and continued until the accumulator emptied at approximately [ ]** seconds (Figure 5.1.2-23). Complete discharge from the
- - accumulators was indicated by a sharp decrease in the temperature of the fluid exiting each accumulator due to the discharge of expansion-cooled nitrogen cover gas, which was released into the primary system (Subsection 6.1.4). Flow from the CMTs was significantly reduced during the discharge of the
- accumulators and increased again once accumulator discharge was completed.
mwmso..ons.ioo395 5.1.2-3 REVis!ON: 1 1 _ _ _ _ _ - . . - , , , , . , . . - , , . . , , , , , , . . , . ., , ,.n.. , . . - , . . - _ , . . - . .
It should be noted that, for Matrix Test SB01, the indicated flow measurement from ACC 2 was not valid due to an error in the data acquisition for that instrument. The integrated mass flow has, however, been included in the mass balance since the total indicated outflow was only in error by a small amount (Figure 5.1.2 26). De IRWST injection valves opened when the reactor vessel pressme low low level setpoint was reached. Injecdon flow only started when the reactor vessel pressure became less than the stade head fmm the IRWST. Figure 5.1.2-16 shows that the IRWST injection began at approximately [ ]** seconds after CMT flow ceased. IRWST flow gradually increased to a peak value of [ ]** lbm/sec. ([ ]*' lbm/sec. per injection line) at [ ]** seconds before gradually decreasing. Mass Ejected from the Primary System At dme zero in the transient, a 2-in. break was initiated at the bottom of CL-3. De mass flow rate from the primary system through the break is shown in Figures 5.1.2-67 and 5.1.2-68. For the first [ ]** seconds following the break, [ ]** lbm of steam and water left & primary system via the break (Figure 5.1.2-62). During this period, the primary system depressurized to around 300 psi (Figure 5.1.21). By the onset of ADS actuation, the cold legs drained, and there was almost no water flow out of the break. Between [ ]** seconds, ADS 1-3 activated and the system depressurized rapidly. Break flow significantly decreased once the ADS activated, since the ADS valve area was significantly larger than the break. At around [ ]** seconds, ADS-4 was initiated, and the primary system depressurized until IRWST injection commenced at [ ]** seconds. Actuation of ADS 1-3 rapidly terminated the flow of steam from the break, although this was replaced by steam flow through the ADS 1-3 valves for the next [ ]** seconds (Figure 5.1.2-63). His steam flow was accompanied by an outflow of water from ADS l-3 at a peak rate of over 4 lbm/sec. (Figure 5.1.2-66). After [ ]** seconds into the accident simulation, the mass flowing through ADS 1-3 was composed almost entirely of water. De rate of flow through the ADS continued at a gradually reducing rate until [ ]** seconds when the ADS-4 valves opened, causing flow through ADS 13 to terminate and be replaced by flow through the lower resistance ADS-4 paths. For ADS 4-1, there was a near-steady water flow rate of [ ]** lbm/sec. from the time of inidadon. Ilowever, for ADS 4-2, there was an initial outsurge at [ ]** lbm/sec. followed by a drop to near zero and an increase to over [ ]** lbm/sec. (Figure 5.1.2-66). He integrated mass flow out of the primary system via the ADS and the break are shown in Figures 5.1.2-62 to 5.1.2-64. During the first [ ]** seconds of the transient, over [ ]** lbm of water left the primary system. Of this, the [ ]** lbm flowing through ADS 1-3 was deposited in the IRWST. He [ ]** lbm leaving the ADS-4 system and the liquid part of the [ ]** lbm flowing through the break were added to the liquid overflow from the IRWST and deposited nearly [ ]** lbm of water into the primary sump (Figure 5.1.2 28) By the end of the short-term transient. l the water level in the primary sump reached nearly [ ]** In. (Figure 5.1.2-29). m%swmso.m:tb-ioons 5.1.2-4 REVIs10N: 1 l l
i i l l- At [ )*' seconds into the transient, the cold legs refilled enough to allow a restart of approximately ( )*'lbm/sec of water flow through the break, so that the total rate of water How from the primary
- i system to the sump was approximately [ ]*' lbndsec. At this time, the flow rate into the reactor
- vessel through the DVI lines was approximately [ ]** lbm/sec.
Mass llalance l Figure 5.1.2 70 presents the variadon in the total system inventory during the short-term transient. ] Following the initiadon of the break, there was an increase in system mass of approximately } [ ]** lbm. By the end of the short-tenn transient, the system inventory decreased by approximately ! [ ]*'lbm compared with the post-break value. During this period of the transient, about ( )** 1bm of water was lost from the primary system as steam (Figure 5.1.2-63). f 4
- in addition td the overall reduction in system inventory, some reductions and recoveries (dips) were j observed. There was a general decrease in system inventory following the initiadon of ADS 1-3 until the start of ADS 4 vendng (around [ ]** seconds). Two other marked dips were also j observed at around [ ]** seconds. These two dips resulted from corresponding changes in the primary sump inventory, which have been observed in the level data from which the masses were
] calculated. The readings on the load cells for the sump did not show corresponding dips, and there l were no indications of How out of the sump at this time. It is, therefore, believed that these resulted j from corresponding increases in pressure above the sump water level and lead to an artificially low j differendal pressure of the sump and a reduced level indicadon. The necessary pressure increase could j have resulted from condensadon in the exhaust line causing a build up in steam until the pressure was ! great enough to drive the venting through the check valve. i j A mass balance analysis has been performed on the primary system. Figure 5.1.2 71 plots the ! measured primary system mass determined by summing the contributions from the reactor vessel, y downcomer, hot and cold legs, SG primary side, pressurizer, and surge line plus the PRHR HX. The i second curve on Figure 5.1.2 71 provides an alternative primary system mass determined from the mass 2 balance, that is given by: i 3 M ' (t) = My(0) + M,(t) - M,(t) 5.1.2-1 i a where:
- M'y(t) = Mass calculated primary system mass j My(0) = ' Measured primary system mass at the start of transient
] M,(t) = Total integrated mass injected from all sources (i.e., accumulators,
, CMT, IRWST, and sumps) to time t M (t) = Mass lost from primary system to dme t via CMT balance lines,
- ADS 1-3, ADS 4, and break ,
)
mAap6000344w 50.nco:Ib-100395 5.1.2-5 REVISION: 1
- 1
l The difference in the two primary system mass curves is shown in Figure 5.1.2 72 as the mass balance Crror. During the short term transient, there was, in general, an overestimate of the mass in the primary system from the measured data of up to [ ]** lbm relative to that calculated from the mass balance, although there was an underesumate following ADS 13 Irdtladon. There are two main contributions to this excess mass. First, the measured primary system inventory did not include all of the pipework in the system, and there was approximately [ )** lbm of mass missing from the inidal inventory. This mass was deposited in and lost from the measured system as pipes drained and refilled. Some of the addidonal mass was subsequendy lost via one of the leakage paths. Second, the instrumentation on the hot legs is believed to have given erroneous level measurements during certain portions of the transient. Figure 5.1.2 58 shows that, for this test, only one hot leg appears to drain, and this contributes an overestimate of ( ]** lbm. By the end of the short term transient, the apparent mass-balance error was approximately [ ]*" lbm. Figure 5.1.2 73 shows the total integrated mass How from and to the primary system, together with the water inventory remaining in the sources of cooling water. During the short term transient, there was a net loss of water from the primary system of approximately [ ]** lbm of which only a small quandty was deposited in the sources. Of the lost primary system inventory, ( )** lbm was lost as steam. The rest has been added to the water stored in the ADS and break separators. Pressure Decay O Figure 5.1.21 shows the primary system pressure during the test. Throughout the LOCA portion of this test, the pressure was controlled by the saturation pressure of the hottest fluid in the primary system. At initiation of the break, the controlling fluid volume was the pressurizer and surge line; however, within the first [ ]** seconds (after the initial blowdown phase), this shifted to the reactor vessel. Figure 5.1.2 3 shows that the temperature of the upper plenum was equal to the saturation temperature corresponding to the primary system pressure measured in the upper head during die natural circuladon phase and into the ADS phase. The pressure stabilized at the saturation pressure for the upper plenum and then continued a slow pressure decrease responding to the cooling caused by CMT injection. Figure 5.1.21 shows an increase in the pressure decay rate that occurred at approximately [ ]*' seconds when the CMTs transitioned from natural circulation injection to draindown injection, which essendally doubled the injection rate of cold water into the DVI line. The higher injection rate resulted in a more rapid temperature drop in the upper plenum (core oudet in Figure 5.1.2 3), which was reflected in a more rapid pressure decay With the actuadon of ADS-1 at approximately [ ]** seconds, the pressure dropped rapidly due to the increased rate of mass ejected from the system, and the increased flow of cold water into the downcomer and the core (Figure 5.1.2 63). 'Ihis continued to reduce the power channel inlet plenum temperature and subcool the heater rods in the core due to the higher flow. Since the reactor vessel outlet plenum became subcooled at about [ ]** seconds, the hottest fluid in the system was in the pressurizer, the cold legs, and the CMTs, and the pressure was partially supported by the flashing of fluid in one or several of mvoocmso.nonab ioows 5.1.2-6 REVis!ON: 1
. . - -_ - .. -. ._ .- = . _. ..
4 these locations. When accumulator discharge ended at between [ ]"' seconds, the reactor vessel temperature again increased to the saturation temperature and took control of the system pressure j for the remainder of the LOCA phase. l 5.1.2.2 Energy inventory lleat removal from the reactor core follows a sequence similar to pressure decay for SBLOCA tests. Before the reactor trips, nearly all the energy generated in the core is removed by the SGs and out of the break with a small fraction lost to surrounding heat sinks. When the reactor trips, the primary system pumps trip, and flow through the SG tubes is sharply reduced. Coupled with the isolation of the SG secondary side, the effect is to significantly reduce heat removal by the SGs. At that time, the PRIIR HX isolation valves open, and energy is temoved to the IRWST as well as out of the break. As the system drains, primary system pressure is reduced, and the sensible heat of the coolant and i metal add to the core heat load. The CMTs start to drain, and the ADS is activated. At that tirne, heat removal is accomplished through ADS flow, and the PRHR HX becomes less effective. Finally, ADS-4 is actuated, the primary system is completely depressurized, and the IRWST is actuated. The q LOCA phase of the test is then completed. The behavior of the components involved in the energy removal is discussed below. 1 I i V Core I 1 The power output of the core is shown in Figure 5.1.2 2. After reactor trip, the core power is representative of decay heat levels expected in the AP600 core. Flow through the core is shown in ; Figure 5.1.2 56 and the steam generation rate is given in Figure 5.1.2 55. As discussed in Section 4.11, the steam production rate has been calculated by two methods, the Tsat method and the DVI line flow method. Figures 5.1.2-53,5.1.2 54, and 5.1.2 55 reproduce the saturation line elevation power split above the
- saturation elevation and steam generation rates from the two methods. Both methods give similar predictions, but neither method gives valid predictions before [ ]** seconds because of flow oscillations during natural circulation in the primary system.
Bath methods for calculating the steam generation predict that the maximum steam generation rate during the LOCA phase will occur at approximately [ ]** seconds, just prior to IRWST injection. l The peak cladding temperature is shown in Figure 5,1.2 57 and indicates that the core is adequately cooled at all times during the test. I
/N U
l l 1 nw oraunsw.so.n=tb tcons 5.1.27 REVISION: 1 l
Steam Generntor IIcat Transfer SGs remove most of the heat from the gimary system during normal operation; however, heat transfer from the primary to secondary side is significantly reduced after the pumps trip. Etis is due to reduced flow in the tubes, which causes a sharp reduction in the tube side heat transfer coef0cient, in addhlon, die secondary side is isolated, which causes the temperature and pressure to remain high as primary-side pressure rapidly decreases. Passive Residual lleat Rrmoval lleat Tran;fer He PRIIR is designed to remove heat from the primary system from the time when the SGs become thermally isolated prior to the initladon of the ADS. One measure of the effectiveness of the PRilR is the increase in the Duld internal energy in the IRWST, which serves as the heat ek for the PRiiR. Figure 5.1.2-33 shows the SG primary and secondary pressure together with the PRl!R integrated heat transfer as regesented by the IRWST Duld energy after allowing for the contribution from ADS 13 inflow. PRilR flow began at about [ ]** seconds when the SG heat removal ended. We heat removal rate was approximately [ l*' Btu /sec, until [ j"' seconds when ADS 1 activated. At that time, the heating rate in the IRWST increased to ( }** Btu /sec. due to corxlensation of steam vented through ADS 13 and heating from the PRilR continues. PRIIR heat transfer rate decreased when the accumulators discharged at about [ ]** seconds as the core became subcooled. Automntic Depressurization System and lireak O Energy removal from the ADS and break are shown in Figure 5.1.2 69. Fluid energy exiting the break increased at a constant rate until ADS 1 actuated. For ADS 1, ADS 2, and ADS 3, compared to the break, the energy removal occurred at a somewhat lower rate due to the reduced system pressure. When ADS 4 actuated, energy removal switched from ADS 13 to the larger How path of ADS-4. De ADS effectively reduced the primary system pressure to allow gravity injecdon How from the IRWST. His flow sufficiently subcooled the primary system, ending core boiling, partially collapsing the steam bubble in the upper plenum and bringing the system to near atmospheric pressure. Overall Energy llalance Figure 5.1.2-74 shows all the energy components in the heat balance for the system during the LOCA phase ([ ]** seconds). Eroughout the event, the heater rod bundle power was the dominant heat input to the system, and before the stan of the event, the SGs pmvided the dominant heat removal. At the start of the event, the SG secondary side was isolated, and the RCPs tripped. During this period, heat removal by the SG was reduced as natural circulation flow occurred in the primary system. As the primary system depressurized, the pressure reached the secondary-side pressure, and heat transfer through the SGs effectively ended. O m.wmuw sowib toows 5.1.28 REVISION: 1
l l l l 1 De steam component of the break flow was leaving the control volume at the start of the event until ADS actuation, lleat loss via this path was nearly [ }"' Bluhee., which is far greater than the reactor decay power ([ ]"* Btu /sec.). The remaining energy lost through the break exhaust resulted from a decrease in the fluid internal energy as the primary system depressurized. l After ADS 13 actuated, the break flow effectively ended, and ADS hea' was deposited into the l IRWST. nus, the fluid internal energy in the control volume increased from this time until the end of the SBLOCA phase as the IRWST water temperature increased. Also at that time, the metal masses in the control volume lost energy as primary system temperature decreased. His resulted in energy being , lost at a rate of nearly ( )*' Btu /sec. between [ ]** seconds. After [ ]*' seconds, the metal masses lost energy at a much lower rate, which is consistent with the primary system fluid temperature decay. Ambient losses were reduced from a maximum of [ ]*' Btu /sec, at full power condittoru to ( ]*' Bru/sec, at the end of the LOCA phase. A large increase was observed in the deficit between the rod bundle power and the various sources of i energy dissipation from the control volume after the initiation of ADS 13, and again after the initiation of ADS 4 at approximately [ ]** seconds. Steam exhaust included steam from the IRWST and the steam portion of the flow from ADS 4. Ilowever, relatively little steam was measured by the vapor flow meters. Further discussion of this steam flow is provided in the mass and energy balarwe in Sections 6.2.2 and 6.2.3. O ; l i I l pv l mwnni..samwtoons 5.1.29 REVISION: 1
TAllt.E 5.1.21 OSU TFST ANAIJSIS STANDARD PLOT PACKAGE FOR SUllSECTION 5.1.2 Plot No, Component Variables Unks Description i Pressuruer CIT 4M psia System pressure 2 RPV RPVPWR kW Core power 3 RPV TOIRPV T08RPV, 'F Core inlet / outlet temperature, ST08RPV saturation temperature 4 SG CIT 201, CIT 2N, psia Primary and secondary pressures in SG Cli 301, CIT 302 5 DVI l WWTDVILI, IbnVsec. Individual components and total flow in WWTIRWil, DVl 1 WOUTACCI, WWTIRW13 6 DVI2 WWEVIL2, IbnVsec. Individual components and total flow in WWTIRWI2, DV!2 WOUTACC2, W%TIRW14 7 ChtT AhtChtTIII, Ibm Fluid mass in ChtTs (excludes balance AhtChtT2B 1mes) H Ch1T CLDP 502, CLDP 507 in Collapsed liquid level in Ch1Ts 9 ChtT hilWDVILI, Ibm Integrated mass out of ChtTs htIWDVIL2 10 ChtT WWTDVILI, IbnVsec. Flow out of CNtTs WW"IT)VIL2 11 ChtT WWTCLilLI, IbnVsce. Flow into ChtTs WW"1CLBL2 12 ChtT CLDP 509, CLDP510 in. Level CL-ChtT balance lines 13 ChtT UChtT1, UChtT2 Illu Fluid energy in ChtTs 14 IRWST IRWST lbm hlass of Guld in IRWST 15 IRWST CLDP 701 in. Collapsed liquid level in IRWST to IRWST WWTlRWil, IbnVsec. Flow from 1RWST to DVI lines WWTIRW12 17 IRWST IRWSTOR lbm/sec, OverCow from IRWST to sump 18 IRWST ADSl3Th!R lbnVsec. Total ADS flow mto IRWST 19 IRWST ADSl3TIR hillRWil, Ibm integrated mass out of IRWST hillRW12, ht!!RWlO 20 IRWST UIRWST Iltu Fluid energy in IRWST 21 PRilR CLDP 802 in. Collapsed liquid level in PRHR lIX 9 m % w o n
- 50 m :th.ta 095 $,1,2 10 REVISION: I
w V TAllt.E 5.1.21 (Continued) USU TEST ANAL,YSIS STANDARD P1,Ol' PACKAGE FOR SilllSF.CTION 5.1.2 l' lot No. Component Variables Units Description 22 PRl!R WWOTPRilR lbm/sec. hicasured outlet 00w from PRilR tube 23 Accumulator Ah!ACCl, Ah!ACC2 lbm htass of Guid in accumulators 24 Accumulator CLDP 401, CLDP 402 in. Collapsed liquid level in accumulators 25 Accumulator WOUTACCI, Ibm /sec. How from accumulators W0lITACC2 26 Accumulator htOUTACCI, Ibm Integrated mass out of accumulators htOUTACC2 27 Accumulator UACCI, UACC2 Blu Ruid energy in accumulators 28 Prirnary sump AhtPShtP lbm Primary sump Ouid mass 29 Primary sump CLDP 901 in. Primary sump level 30 Primary sump UPShtP Btu Primary sump Guid energy 31 SG htSSGIPI, htSSGIP2, Ibm hiass of Guid in SG primary side htSSGOPI, htSSGOP2 inlet / outlet plena 32 SG htSSGIITI, htSSGIIT2, Ibm blass of Guid in SG primary side hot htSSGCTI, htSSGCT2 and cold tubes 33 SG/PRilR CF'r.201, CirT 301, psia & SGI pressure and PRl(R integrated heat QPRIIRI Blu output 34 Pressurizer PZht Ibm Ruid mass in pressuriier '%~) 35 Pressurtier CLDP ft)! in. Collapsed hquid level in pressurizer 36 Pressurizer UPZ Blu Fluid energy in pressuriier 37 Surge line PLht Ibm Ruid mass in surge line 38 Surge line CLDP-(i)2 in. Collapsed liquid level in surge line 39 Surge hne UPSL Btu Ruid energy in surge line 40 RPV NtWRPV lbm Total Guid mass in reactor vessel 41 RPV DCht Ibm Ruid mass in downcomer 42 RPV LDFMitX' in. Collapsed liquid level in downeomer compared to various reference elevations 43 RPV htWOIRPV lbm Ruid mass in lower plenum 14 RPV htWO3RPV lbm Ruid mass in core region 45 RPV LDP03RPV in. Collapsed liquid level in core 46 RPV RPVAVDF2 Core exit void fraction 47 RPV RPVAQOU2 Core exit quality 48 RPV htWO6RPV lbm Fluid mass in the upper plenum 49 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 50 RPV htWO8RPV lbm Fluid mass in the upper head Q 51 RPV LDP03RPV in. Collapsed liquid level in the upper head mvm2144w.so nalb-ioo195 5.1.2 11 REVISION: I
l l l i TAllt.E 5.1,21 (Continued) ONU TFST ANAINSIS STANDARD Pl OT PACKAGE FOR SUllsECT10N 5.1.2 9 l' lot No. Component Variables Units Description 52 RPV URPV litu Total Guid energy in reactor vessel 53 RPV RPVXE, RPVASL2 ft Level of Tsat line 54 RPV RPVPab, RPVAPab2, kW lleated rod power above and below RPVPWR Tsat level and total 55 RPV RPVRXV, RPVASOU2 lbnVsec. Core steam generation rate 56 RPV RPVALIN2 lbnVsec. Calculated core now 57 RPV IIThtXRPV, ST08RPV *F htaximum clad temperature and saturauon temperature 58 liot leg htWl(LI, h!WilL2 lbm Water mass in bot legs 59 Ilot leg htVIIL1, htVi!L2 lbm Vapor mass in hot legs 60 Cold leg CLIWhtS, CL2WhtS, Ibm Water mass in cold legs CL3WhtS, CL4WhtS 61 Cold leg CLI VhtS, CL2VhtS, Ibm Vapor mass in cold legs CL3VhtS, CL4VhtS 62 ADS and break BRKSTIR, ADSl3TIR, Ibm Total discharged mass for ADS l 3, ADS 41TIR, ADS 42TIR ADS 4s, and break 63 ADS and break BRKTIVF, ADl3TIVF, Ibm Totalintegrated vapor How for ADS AIMii;VF, AD42TIVF and break 64 ADS and break BRKTILF, ADl3TILF, Ibm Total integrated liquid now for ADS AD41TILF, AIM 2TILF and break 65 ADS and break ADS l3SVR, IbnVsec. Vapor tiow out ADS 13 and ADS 4 ADS 41SVR, ADS 42SVR 66 ADS and break ADS l3SLR, Ibm /sec. Liquid 110w out ADS 13 and ADS-4 ADS 41SLR, ADS 42SLR 67 ADS and break BRKSSVR lbnVsec. Vapor How out of break 68 ADS and break BRKSSLR lbnVsec. Liquid now out of break 69 ADS and break BRKSPEl, ADSl3EI, Blu Integrated Guid energy for ADS 13, ADS 4tEl, ADS 42EI ADS-8, and break 70 htass balance TOThtASS lbm Total system mass inventory 71 hiass balance PRIhthtASS, Ibm bicasured primary system inventory and PRIhtASS2 value from tnass balance 72 htass balance htERROR lbm hiass balance error 73 htass balance htIN,h!OUT lbm Integrated mass now in and out of SRChtASS primary system and source mass 74 Energy balance Various Btu Components of energy bahnce O m w m m w .$o m ist M 9s 5.1.2 12 REVIS10N: I
l 1 D-( i Tile FIGURES LISTED IN TAllLE 5.1.21 ARE NOT INCL.UDED IN Tills NONPROPRIETARY DOCUMENT O o mh inw)44w.50.noa:tb toons 5.1.2 13 REVI$lON: 1
M
'O
!d 5.1.3. Long Term Transient
%e long term transient covers the transition from IRWST to sump injection and provides information
- on the LTC response of the AP600. For the 2 in, cold leg break, Matrix Test SB01, the long-term transient encompasses die time frame of [ ]*' seconds to the end of the test near
[ ]*' seconds. De behavior of the test facility during this period of the transient is discussed in this subsecdon using the plot package detailed in Table 5.1.31. Dese results concentrate on the
; components of the primary system that remain active during the LTC phase, that is, the RPV, the hot legs, ADS-4, the sumps, and the IRWST.
5.1.3.1 Maintenance of Core Cooling j Reactor Pressure Vessel and Downcomer Mass Distribution For the long term transient, the passive core cooling systems must supply sufficient flow to prevent any overheating of the heater rods. During the time frame of [ ]** seconds, the decay heat simulation of the heated rods reduced power from 175 to 120 kW (Figure 5.1.31). As seen in Figure 5.1.3 38, there were no significant excursions in heater rod temperatures and, therefore, sufficient core flow was maintained throughout the long-term transient.
%e mass of water in the reactor pressure vessel is shown in Figure 5.1.3 25. After an initial decline, the reactor vessel water mass settled at an average value of [ l# Ibm until the sump injection ]
valves opened at around [ ld seconds. From [ l'*d seconds, oscillations in vessel i .I inventory were observed. Rese oscillations can be seen in measurements throughout the primary system, and they are discussed further in Subsection 6.1.3 De onset of sump injection caused a drop , in vessel inventory to a constant value of [ l'*d Ibm, which is 67 percent of the initial vessel water I l Inventory. At approximately [ ]** seconds, the water mass increased to [ ]'6d lbm in response to a partial collapse of the steam bubble in the head region (Figure 5.1.3 33) and remained at j this level during the last [ ]** seconds of the translent. ! 1 k Core water mass and the collapsed liquid level are shown in Figures 5.1.3-28 and 5.1.3 29. From i [ ]'6d seconds, the core remained near water solid with only a very low level of boiling ) l (Figure 5.1.3 36); however, after the sump injection valves opened, the increase in core temperature I l resulting from the influx of hot sump water (Figures 5.1.3 2 to 5.1.3-5) reduced the core collapsed liquid level to just below the top of the heater rods. At this time, the level at which the core fluid temperatures reached saturation dropped (Figure 5.1.3 34), and the amount of boiling increased along the upper regions of the heater rods. De level of bolling continued at a constant rate until the end of , the accident simulation (Figure 5.1.3-36). Note that the long term transient core steam generation calculation is based only on the DVI line flow method, since the Tsat method is very sensitive to small changes in local pressure. .O mwea2m sanon:thioo395 5.1.31 REVISION: 1
The collapsed liquid level in the upper plenum region is shown in Figure 5.1.3-32. Right before sump injection began, the collapsed liquid level in the upper plenum remained at the top of the hot legs. Following the influx of hot water from the sumps, the level dropped to the top of the DVI lines and remained there for the remainder of the transient. The mass of fluid and collapsed liquid level in the reactor vessel downcomer are shown in Figures 5.1.3-26 and 5.1.3-27. Before the start of sump injection, the collapsed liquid level in the downcomer was at the level of the center of the cold legs. The start of sump injection through the injection valves at [ ]^* seconds caused a readjustment of the levels around the primary system, and, subsequently, the downcomer level remained at the bottom of the hot legs for the remainder of the transient. Figure 5.1.3-4 shows the response of the downcomer water temperatures during the LTC phase of the transient. Initial sump injection via the check valves caused a sudden rise in downcomer water temperrture as the hotter sump water entered the downcomer. The opening of the main sump injection valve momentarily increased downcomer temperature as DVI flow temporarily decreased (Figures 5.1.3-6 and 5.1.3-7). During the long-term phase of the transient, the downcomer water temperatures increaseu gradually, but it can be seen that the temperature remained well subcooled. The downcomer, therefore, provides a potential site for steam condensation. Such a process is believed to be involved in driving the oscillations observed between [ ]** seconds into the transient (Subsection 6.1.3). Loop Mass Distributbn During the LTC phase of the transient, there was a low level of Miling in the upper regions of the core. The following are five potential paths for the steam:
- To the hot legs and out of the ADS-4 valves To HL-2, through the pressurizer and ADS 1-3 valves to be deposited in the IRWST
- To IIL-2, through the PRIIR HX to the SG outlet To the hot legs, through the SGs to the cold legs to return to the reactor vessel or flow out of the break To the head region and into the downcomer, potentially condensing in the downcomer, or, if the downcomer ,/ater level is below the cold legs, venting through the cold legs to the CMTs and break l Figures 5.1.3 39 and 5.1.3-40 show that during sump injection, two-pim. flow occurred in the hot legs. Figure 5.1.3-23 shows that throughout this phase of the transient, die pressurizer surge line mAap600G344w-50.non:1b 100395 5.1.3-2 REVISION: 1
___s_.
4 remained filled with water. Figure 5.1.3-21 shbws that the inlet plenum on SG-2 also remained ( plugged with water. It was expected that the inlet plenum on SG 1 also contained a small amount of water, although for this test, the level measurement was over-ranged 'and SG-2 results have been used
- for SG-1. There is no evidence of flow through the PRHR HX during the long term-transient.
'Iherefore, for steam enteriug the hot legs, the only available flow path was out of the primary loop to the sumps via the ADS-4 valves. However, there was no evidence on the steam vortex meters for vapor flow out of the ADS. Note that the vortex meters have a deadband, and low flow was not
. detected. 'Ihe expected rate of loss of steam through each ADS-4 valve during the long-term transient is of that order. As discussed in Subsection 6.2.2, examination of the fluid thermocouples in the steam lines from the ADS-4 separator shows the temperature at or above saturation for all of the transient beyond 15,000 seconds, which indicates that steam was leaving the ' system by this path at a low flow rate.
Mass Ejected from the Primary System Integrated mass flow out of the primary system via the ADS and the break is shown in Figure 5.1.3-43. By the end of the accident simulation, approximately [ j'6# lbm of' vater had flowed out of the primary system. During the LTC phase of the transient, the only significant outflow was through the ADS-4 valves, with a small apparent flow through tue break. The apparent break flow does not represent flow out of the primary system, but indicates continued interaction between the break separator and the sump. The most marked manifestation of this interaction is at [ j'6' seconds when the primary sump began to flow into the secondary sump causing oscillating flow indications in the liquid flow out of the break separator. This is confirmed by Figures 5.1.3-44 and 5.1.3-45, which show flow through the ADS and the break. During the sump injection phase of the transient, outflow in the form of liquid exited out of the ADS-4 valves. Water flowed through each of these at an average rate of [ ]'6# lbm/sec. During the long-term transient, there was no evidence from the vortex meters for steam flowing out of
- the primary system via the ADS-4 valves, it is, therefore, m. obvious where the steam generated in the core was ducted. As discussed in Subsection 6.1.3, there is evidence for steam flowing from the ;
upper head to the downcomer via the bypass holes, but this estimated flow is not enough to account , for all the steam generated in the core. The system mass inventory (Subsection 5.1.3.5) indicates a i ilow of [ j'6"lbm/sec. and is required to explain the fall in measured system mass inventory. As noted above and in Subsection 6.2.2, fluid temperatures in the vapor lines out of the ADS-4 stage separators indicate that steam left the primary system via the ADS. At approximately [ ]'6# seconds into the transient, the level in the primary sump (Figure 5.1.314) reached the point at which overflow to the secondary sump occurred. At this time, there was approximately [ j'6# lbm of water collected in the primary sump. From the start of primary sump overflow to the end of the transient, [ ]*6# lbm of water was transferred to the secondary sump (Figure 5.1.3-16). I ; mAap600\2344w-50.nos:ll>100395 5,],3 3 REVISION: I b
Mass Injected to the Primary System
'Ihe total DVI line flow, CMT flow, and IRWST flow are shown in Figures 5.1.3-6 and 5.1.3-7; flow from the pdmary sump is shown in Figure 5.1.3-19, From around [ ]'*' seconds, there was a contribution to the DVI flow from the CMTs as they finished post-refill draindown.
During the presump injection phase of the transient, IRWST flow proceeded at a gradually reducing rate vdth the effect of the primary system oscillations superimposed. At [ ]'** seconds, flow from the primary sump began through the check valves around the main injection valves. After the initial outsurge, a flow rate of [ ]'*d Ibm /sec. was achieved. At around [ ]'*# seconds, the level in the IRWST fell to [ ]'A'in and the primary sump injection valves opened, allowing additional flow to the primary system from the sump with an initial surge of water at a peak rate of [ ]*** lbm/sec. (Figure 5.1.3-19). As the levels in the primary system adjusted, flow through the DVI lines ceased temporarily before restarting again. Following the start of sump injection through the main valve, flow through DVI-l resulted entirely from aump injection (Figure 5.1.3-6) while for DVI-2, flow was mainly from the IRWST (around [ ]'6' lbm/sec.) with a small contribution from the primary sump flow (Figure 5.1.3-7). Both DVI lines delivered water to the reactor vessel at a near-constant rate of just under [ j'*' lbm/sec. for the entire period of the transient following the opening of the sump injection valves. Note that IRWST flow rates were based on measured positive flows. During sump injection, there was reverse flow from the primary sump to the IRWST via the DVI 1 line. During sump injection, [ ]'*# lbm of water was delivered from the primary sump to the primary system in addition to [ ]'** lbm from the IRWST (Figure 5.1.3-20). During this time, a total of around [ ]'** lbm of water left the primary system via the ADS-4 valves. The inventory of water available for injection to the primary system was not reduced significantly by the end of the transient (Figure 5.1.3-13). During the long-term transient, the water level in the break separator was sufficient to allow flow back into the primary system through the break. This is representative 01'the behavior expected in the AP600 where the level of water in the sump would reach the break elevation. Mass llalance Figure 5.1.3-46 shows the variation in the total system mass inventory during the entirety of Matrix Test SB01. Following the short-term transient, the total inventory remained between [
]'** lbm below the initial value. Initially, there was an increase in inventory until
[ ]'** seconds when the inventory fell. From [ ]'b' seconds to the start of sump injection, the inventory remained essentially constant. After sump injection began, there was an increase in inventory followed by a gradual reduction, so that by the end of the LTC phase of the transient, [ ]'*# lbm of water was also lost. This reduction was consistent with a flow rate of steam out of the system of [ ]'*d Ibm /sec. Note that the vortex flow meters have a deadband, and thus, m:\ap600Zl44w 50.noo:Ib-100375 5.1.3-4 REVis!ON: 1
1 l
/
f' ( ' the vortex meters would not register the low flow of steam passing through each ADS-4 valve; therefore, this inventory loss was not measured directly. Figure 5.1.3-46 shows dips of about [200] lbm superimposed on the general trends in mass inventory.
. As in the short-term transient,' these dips resulted from corresponding changes in the sump inventory, I observed in the level data from which the masses were calculated (Figures 5.1.3-13,5.1.3-14,5.1.3-16,
'and 5.1.3-17). %c readings on the load cells for the two sumps did not show corresponding dips, and l
there were no indications ofincreased flow at this time; thus, the dips are believed to be the result of . i corresponding pressure variations in the differential pressure taps and not mass changes. De mass balance calculation has been performed for the entire transient, and the results are presented in Figures 5.1.3-47 to 5.1.3-49. During the long-term phase of the transient, the measured water inventory in the primary system remained approximately constant at [950] lbm. De mass balance calculations do not fully allow for reverse flow from the break separator to the primary circuit. . i 5.1.3.2 Energy Halance i Figure 5.1.3-50 shows all the energy components in the heat balance for the system during the LTC phase. During this phase, the heater rod bundle power was the dominant heat input to the system. De SG heat transfer ended during the LOCA phase and did not contribute to the overall energy l Q
\/
balance during the LTC phase. Bus, for the L'IE phase, the active components in the overall energy balance were rod bundle power, change in the fluid internal energy, change in the metal internal energy, ambient losses, and steam exhaust from the control volume. he fluid energy in the control volume increased steadily until sump injection began at (approximately [ ]* seconds). At that time, fluid throughout the system approached saturated conditions, and the rate of increase was reduced. At the same time, metal mass temperatures increased and the primary system temperature increased in response to the hotter water from the sump. Also, ambient losses increased slightly as the control volume temperatures increased. De sum of these increases is much less than the rod bundle power and must be a result of steam exhausting from the control volume. His assumption is consistent with the mass balance deficit discussed previously. De steam component of the ADS-4 flow left the control volvme from actuation of ADS-4. For the LTC phase, this quantity is essentially zero, as measured by the vapor flow meters. It is concluded that the vapor flow meters in the ADS-4 separators did not measure the relatively small steam flow rates accurately. Further discussion of this steam flow is provided in the mass and energy balance in Sections 6.2.2 and 6.2.3. D). k,
. ma.peom2mso. o.:isioo395 5.1.3-5 REVIslON: 1
'4
TABLE 5.1.31 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.I.3 LONG TERM TRANSIENT Plot No. Component VariaUes Units Description i RPV RPVPWR kW Core power 2 Primary sump TSMPI1, TSMPI2 *F Sump injection line temperatures 3 DVI TDVIL1, TDVIL2 'F DVIline temperatures 4 RPV TOIDC, T02DC, f03DC, 'F Water and saturation tempera:ures in ST01DC downcomer 5 RPV TOIRPV, 'II)8RPV, "F Core inlet / outlet temperature, ST08RPV saturation temperature 6 DVI l WWTDVIL1, Ibm /sec. Individual components and total flow WWTIRWil, in DVI-I WWTIRWI3 7 DVI-2 WWIDVIL2, Ibm /sec. Individual components and total flow WMWI2, in DVI 2 WWTIRW14 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 10 IRWST IRWST lbm Mass of fluid in IRWST 11 rRWST CLDP-701 in. Collapsed liquid level in IRWST 12 IRWST UIRWST Btu Fluid energy in IRWST 13 Primary sump AMPSMP lbm Prtmary sump fluid mass 14 Prunary sump CLDP-901 in. Primary sump level 15 Primary sump UPSMP Btu Primary sump fluid energy 16 Secondary sump AMSSMP lbm Secondary sump fluid mass 17 Secondary sump CLDP-902 in. Secondary sump level 18 Secondary sump U3SMP Blu Secondary sump fluid energy 19 Primary sump WSTSMPET, WWTSMPIT lbm/s:c. Primary sump steam and liquid injection rate 20 Primary sump MISMPII, MISMP12, Ibm Integrated primary sump and IRWST MISMPIT, MIIRWT flows 21 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG side inlet / outlet MSSGOP1, MSSGOP2 plena 22 Surge line PLM lbm Fluid mass in surge line l 23 Surge line CLDP-602 in. Collapsed liquid level in surge line 24 Surge line UPSL Btu Fluid energy in surge line 25 RPV MWRPV lbm Total fluid mass in reactor vessel O m:W344.-50.noo:ib-too395 5.1.3-6 REVISION: 1
I l i TABLE 5..t.31 (Continued) OSU TEST ANALYSIS STANDARD l>i OT PACKAGE FOR SUBSECTION 5.1.3 LONG TERM TRANSIENT Plot No. Component Variables Units l Description 26 RPV DCM lbm Fluid mass in downcomer 27 RPV LDP01DC ic Collapsed liquid level in dowmcomer compared to various reference elevations 28 RPV MWO3RPV lbm Fluid mass in core region 29 RPV LDP03RPV in. Collapsed liquid level in core , 30 RPV RPVAVDF2 Core exit void fraction 31 RPV RPVAQOU2 Core exit quality 32 RPV LDP06RPV in. Collapsed liquid level in the up' . plenum 33 RPV MWO8RPV lbm Fluid mass in the upper head 34 RPV RPVASL2 ft. Level of Tsat line 1 35 RPV RPVAPab2, RPVPWR kW Heated rod power above and below Tsat level and total 36 RPV RPVASOU2 lbm/sec. Core steam generation rate 37 RPV RPVALIN2 lbm/sec. Calculated core flow 38 RPV HTMXRPV, 'F Maximum clad temperature, saturation ST08RPV temperature and delta 39 Hot leg MWHL1, MWHL2 lbm Water mass in bot legs 40 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs 41 Cold leg CLlWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 42 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 43 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS l-3, i ADS 41TIR, ADS 42TIR ADS 4, and break 44 ADS and break ADS 13TLR, ADS 41TLR, Ibm /sec. Liquid flow out ADS 13 and ADS-4 ADS 42TLR 45 ADS and break BRKSTLR lbm/sec. Liquid flow and total flow out of break 46 Mass balance TOTMASS lbm Total system mass inventory 47 Mass balance PRIMMASS, PRIMASS2 lbm Measured primary system inventory and valve from mass balances 48 Mass balance MERROR Ibm Mass balance error 49 Mass balance MIN, MOUT SRCMASS lbm Integrated mass flow in and out of primary system and source mass 50 Energy balance Various Btu Component of energy balance 51 ADS-4 ADS 41TLR, ADS 42TLR lbm/sec. Oscillations in ADS-4 liquid flow 52 Surge line CLDP-602 in. Oscillations in surgeline level , m:up6 coa 3w.50.non:Ib-too395 5.1.3-7 REVISION: I
TAllLE 5.1.3-1 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.1.3 LONG TERM TRANSIENT Plot No. Component Variables Units Description 53 RPV CirT-107 psia Oscillations in upper bead pressure 54 RPV CLDP-113 in. Oscillations in upper plenum level 55 RPV LDP03RPV in. Oscillations in core level 56 RPV LDP0lDC in. Oscillations in downcomer level l O l t t l 9 m:W3h50.non:tvioo395 5.1.3-8 REVISION: 1
,___._._.__....,___._...__r.__._ . . . _ . . . - _ . _ _ _ _ _ _ _ . _ . _ _ . _ _ _ . _ . _ . , _ _ _ _ _ . _ _ .
i, THE FIGURES LISTED IN TABLE 5.1.34 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT l i h i l r l I 1 t i i~ l l ( I i i
- mp344. 50..o
- b.ioo395 5.1.39 REVISION: 1-E
- v. - . . . . . . . , , . - . - . . _ . . . . . . - _ . . _ - - . _ - - . . . . . . _ . - - .. - - - - -
3 (j ( 5.2 Analysis of Matrix Test SB18 Matrix Test SB18 (OSU Test U0018) simulated a 2-in. cold-leg (CL) break SBLOCA with LTC and without operation of nonsafety-related systems. The break was located at the bottom of CL-3 with a simulated failure of one of the ADS 4 lines. CL-3 is on the CMT side of the facility. The analysis of Matrix Test SB18 is divided into three sections: General facility performance (Subsection 5.2.1) describes the overall response of the system throughout the test. The performance of the facility is characterized by the figures listed in Table 5.1.1-1.
+ SBLOCA (Subsection 5.2.2) provides a discussion of the system behavior from the start of the test, through system depressurization, to approximately [ )**' seconds into the transient, and includes the initial system blowdown, the establishment of natural circulation, and the initial portion of the IRWST injection cooling (Figure 5.2-2).
+
LTC (S ibsection 5.2.3) discusses the behavior of the remainder of the test and includes the completion of IRWST injection and the establishment of sump injerfbn. The refill and subsequent recirculadon of the CMT is considered as a separate discussion within
'- Subsection 6.1.1. The period between SBLOCA and LTC is not discussed specifically since the system was behaving in a stable manner.
Matrix Test SBl8 was a duplication of Matrix Test SB01. The purpose of performing Matrix Test
- SB18 was to confirm the ability of the facility to replicate its response to a SBLOCA, with the same 1 configuration from the beginning to the end of the test program.
f The differences between SB01 and SB18 are as follows:
- In SBl8, a vacuum breaker was installed on the ADS 13 sparger line inside the IRWST to eliminate negative pressures in the pressurizer and ADS 1-3 separator.
In SB18, pressurizer heater logic was changed so that the PLC initiated a signal to open the pressurizer heater SCR contactor at [ ]'*# seconds after S signal actuation, thereby ensuring de-energization of the heaters. 1
+
In SB18, to prevent heating at the top of the CMTs prior to break valve opening, CMT i balance line isolation valves were closed and opened by the operator 1 minute after the TEST pushbutton was pressed. 1 ,n l maap6002344w-52.noo:Ib-100395 5.2-1 REVISION: 1 l
. -_ _. .~
l
) -y
( 5.2.1 Facility Performance 1 A flow nozzle simulating one ADS-4 valve was installed in the ADS 4-1 line, hot leg-1 (HL-1) to the ADS 41 separator, to provide the single failure simuladon. A flow nozzle simulating two valves was
- installed in the ADS 4-2 line (HL-2 to the ADS 4 2 separator). ' Additionally, flow nozzles simulating two lines of flow each were installed in the ADS 13 inlet lines.
The reactor heater control decay algorithm maintained the maximum reactor heater power output for
-! ]'6' seconds, and then the power was programmed to begin to decay, simulating the total decay energy input of the AP600 nuclear fuel. This test was performed with reactor heater rod HTR-C2 317 electrically disconnected to simulate the heater conditions during the performance of Matrix Test SB01.
The facility performance is divided into separate discussions of the five phases of the test:
- Blowdown
- Natural circuladon
- ADS
- IRWST injection '
- Sump injection O
'V 1he overall performance over the [ ]'b* second test is shown in Figures 5.2.1 1 to 5.2.1-4.
Figure 5.2.1 1 shows the pressurizer pressure throughout the test, indicating the various phases and operating components. For clarity, the time scale is split out between 1600 and 13,000 seconds since there was no change in the operating mode during this period. Figure 5.2.12 shows the total injection flow rates into the DVI line from the various systems as a function of time. Figure 5.2.1-3 shows the quantity of steam generated in the core throughout the test. Figure 5.2.1-4 shows the variation in l average measured core outlet temperature and peak clad temperature relative to the core outlet saturation temperature. i l Figures 5.2.1 1 and 5.2.12 show that a continuous flow of water into the reactor vessel is provided by passive safety-related systems as the primary system was depressurizing. The operation passive safety-related systems overlapped so that as one system drained or empded, another provided flow into the simulated reactor vessel for continuous core cooling. 4 Sufficient flow to the core was maintained so that the average measured core outlet temperature is just saturated or subcooled for significant portions of the transient, and the core steam flow is less than the passive safety-related system injection flow, as seen by comparing Figures 5.2.1-2,5.2.1-3, and 5.2.1 l- 4.- As the system transitions into L1C, the water injected from the sump was b;t since it originated in the primary system. The hot injection water temperature after [ l# seconds resulted in continuous
~
steam generation in the heater rod bundle as seen in Figure 5.2.1-3. The resulting steam generation is vented primarily through the ADS-4 valves, m:\ap60062344w 52. con:lt>100395 5.2.1-1 REVIs!ON: 1 i
S.2.1.1 Blowdown Phase Re blowoown phase corresponds to the first [ ]'6* seconds of Test SB18, as shown in Figure 5.2.1 1, and is completed when the steam pressure reaches the SG safety valve setpoint. As with Matrix Test SB01, the test was initiated at time zero by opening the break valve. The hot leg of the RCS was at the same temperature and pressure (420'F at 372 psig) prior to the initiation of the test as for Test SB01. He simulated S signal was generated 0.5 seconds after the break signal and initiated the following actions. In the first [ ]** seconds, the SG pressure setpoint was raised to 335 psig, the reactor shifted to power (kW) control mode with a programmed power demand for 600 kW total power, the main feedwater pump tripped and feedwater was isolated (at [ ]** seconds), the PRIIR llX outlet valve and CMT discharge valves opened (at [ ]** seconds), and the RCPs tripped [ ]** seconds into the event. Forced flow was continued through the PRIIR llX and the CMTs until the RCPs stopped at approximately [ ]** seconds, at which time the PRHR llX flow changed to natural circulation. As the RCS depressurized and coolant escaped through the break, pressurizer level decreased rapidly and steam formation began in the reactor vessel upper head. At about [ ]** seconds, the level in the upper head of the reactor vessel indicated the vessel was beginning to lose inventory as the vessel drained and some liquid flashed to steam. As the primary system pressure fell to near a steady-state condidon, the system transitioned into the natural circulation phase once the pumps coasted down; the system reached the SG pressure setpoint at 335 psig at approximately [ ]** seconds. During this time period, there was initially forced flow and then liquid single-phase natural circulation in the PRHR and CMT. De CMTs provided recirculadng flow to the reactor sessel, while the PRHR removed energy from the primary system. 5.2.1.2 Natural Circulation Phase ne cold legs developed a void fraction at approximately [ ]** seconds at which time the CMT balance lines began to drain. CMT-1 and CMT-2 levels began to decrease, making the transitjon from recirculation to draindown at about [ ]** seconds, respectively, and the injection flow from the CMTs increased (Figure 5.2.12). As the system continued to drain, the SG tubes started to drain at [ ]*' seconds, and the SGs transitioned into a two-phase recirculation behavior. De mass loss through the break decreased the pressunzer level and emptied the pressurizer at approximately ( ]** seconds. At both [ ]*' seconds, a condensation /depressurization event took place in CMT-1, as indicated by a rapid refill of the CMT-1 beJance line as steam from the balance line was condensed in the CMT. Water from the cold leg was drawn up the balance line into the CMT as the balance line filled. He upper head continued to drain and was empty at [ ]** seconds. Steam generation in the reactor vessel reached its maximum ([ ]*' lbm/sec., Figure 5.2.1-3) at [ ]** seconds. The i maap6coum52..o.:ib.ioow5 5.2.1-2 REVIs!ON: 1 l
1 pressurizer surge line was completely empded at approximately [ j seconds. The primary system was at a pressure above the SG secondary side pressure, which continued to remove energy until approximately [ j seconds when the tubes become superheated. As the system contlnued to drain, the Uebes of both SGs were completely empty by approximately [ ]'6# seconds, and the HL 1 and IE-2 levels began to decrease at about [ ]'6# seconds. The horizontal sections of the hot legs started to drain at about [ j'6d seconds. The hot legs remained at saturation temperature and never superheated, even though they were partially or completely empty due to a small flow of saturated steam from the reactor heater bundle to the SGs. At[ ]'6d seconds, CMT 1 reached the ADS 1 setpoint, and the ADS-1 valve opened to initiate ADS blowdown. 5.2.1.3 Automatic Depressurization System Phase The ADS flow path, in conjunction with the break, decreased RCS pressure at a rapid rate, 2 redistributing the mass inventory of the system. The opening of the ADS 1 valve released two-phase flow through the pressurizer to the ADS 13 separator and into the IRWST through the sparger. The opening of the ADS 1 valve, followed by the ADS-2 valve approximately 1 minute later, increased the rate of RCS depressurization. As the different ADS stages opened, the primary vent path shifted from the cold leg break to the ADS valves through the pressurizer. The collapsed liquid level inside the reactor core barrel reached a near-term mlnimum value at j [ J'6' seconds, but the core remained covered during this time period with a two-phase mixture. l When the RCS depressurized to approximately [ ]'b' psia at about [ ]'** seconds, accumulator injection into the DVI line began (Figures 5.2.1 1 and 5.2.12). The accumulators discharged into the DVI line, which reduced CMT 1 injection flow during , accumulator injection by closing off the CMT discharge line check valves until the accumulators were ! almost empty and depressurized. With the accumulators at their maximum injection rate, the RCS I 1 refilled and the surge line and pressurizer began to reflood at about [ ] seconds; the pressurizer attained its maximum level at [ ]'*d seconds. Once the accumulators were empty, the pressurizer and surge line then drained down and were completely empty at [ ]'6' seconds. 1 When RCS pressure decreased to [ ]'6d psig at approximately [ ]'6d seconds, the two IRWST injection valves automatically opened, but IRWST injection could not occur until the RCS pressure decreased to near atmospheric pressure. As the accumulators completed injection at [ ]'6# seconds i , for ACC-1 and [ l'** seconds for ACC-2, CMT-1 and CMT-2 injection flow started to increase at l [ ]'6' seconds. Approximately 50 percent of the nitrogen gas in the accumulator was injected into the DVI lines (see Subsection 6.1.4), momentarily cooling the injection lines at the end of accumulator injection. .The nitrogen caused a decrease in ACC-1 and ACC-2 outlet temperature of about [ J'6# 'F j at approximately ( ] seconds. There was no indicated change in total DVI flow in either DVI line, maap6am23m 52.=1b 100395 -- 5.2.1 3 REVISION: 1
l During the accumulator injection period, there was sufficient injection of subcooled liquid to suppress boiling in the cose region. The downcomer was filled with subcooled liquid, which resulted in the collapse of the superheated steam bubble in the upper portion of the reactor vessel downcomer annulus. As the pressure decreased in this region, the downcomer fluid accelerated upward and impacted on the bottom of the core barrel flange where the core bypass holes are located. The impact of the downcomer liquid on the solid surface of the core barrel flange was heard during the test. During accumulator injection, the PRHR level decreased. The PRHR llX inlet temperature became subcooled coincident with the ADS-4 valves opening at [ ]'6' seconds, and over the next [ ]'6' seconds, dropped to and paralleled the outlet temperature. From [ j seconds, both DVI nozzle fluid temperatures increased from essentially ambient conditions to as high as [ l F as the CMT fluid heated The DVI fluid temperature transient was terminated when the CMT inventory was exhausted, and the IRWST injection began refilling the reactor vessel at about [ l seconds, temperatures returned to ambient when the CMTs were empty, terminating the hot liquid injection. This temperature transient does not appear to have affe ted any other facility parameters. Steam generation in the reactor vessel (RPV) was re-established in the period between about [ J seconds after accumulator injection ended and the system began to drain again. At [ ] seconds, the ADS 4-1 and ADS 4-2 valves opened automatically when CMT 1 level reached its low-low level setpoint. ADS 4 actuation started a decune in RCS inventory up until IRWST injection began. There was initiahy too much mass in the system to be vented through ADS-4 before IRWST injection could occur. CMT-1 and CMT-2 were completely empty at [ J"' seconds, respectively. The pressurizer was slightly subcooled at about [ l'6' seconds and remained subcooled until the data acquisition ceased at approximately [ l seconds. , At about [ ]'"' seconds, the RCS drained sufficiently to decrease system pressure to about [ l'"# psig, which was sufficiendy low that the IRWST static head was greater than RCS pressure, and IRWST injection began. 5.2.1.4 In Containment Refueling Wnter Stornge Tnnk Injection Phase IRWST injection started at about [ j seconds and proceeded at a continually diminishing rate (Figure 5.2.1-2) as the differendal head between the IRWST and the RCS decreased with the drainage of the IRWST. The IRWST injection was sufficient to begin refilung the primary system. The pressurizer and pressurizer surge line emptied for the second time at approximately [ l'6' seconds, respectively. No reflood of the pressurizer occurred because of the vacuum breaker installed on the ADS 1-3 sparger line inside the IRWST. O m:WaxA2%52.non:ib toons 5.2.1-4 REVISION: I
- ..___ -.- .... __ .--_.. .__ _ _ _ _ _ _ _ _ _ _._. _ _~ .
Both CMT balance lines began to refill about [ ]'6' seconds when the IRWST injection increased , the reactor vessel level, sufficiently covering and refilling the cold legs. At about [ ]** seconds, when the Chfr 2 balance line had completely refilled, CMT-2 began to rapidly refill and reached the [ l in. level (about [ l"' percent full) at about [ l'*d seconds. The CMT refill will be discussed in more detail in Subscction 6.1.1. After the CMTs were pardally refilled, there was no
- injection flow from the CMTs because the higher static head of the IRWST held the Chff discharge
, line check valves closed. 4 Steam generation started again at about [ ld seconds and continued for the remainder of the test (Figure 5.2.13). CMT-1 and CMT 2 remained at essentially constant levels for several
- thousand seconds and then began slow draindowns at about [ ]'6d seconds, respectively.
I The draindown for both CMTs was slow and did not occur undl IRWST relative level was [ ]** In. below that of the CMTs Data indicate that the CMTs drained for a while, and then the differential head between the IRWST and the CMTs again closed the CMT discharge check valves, terminating draining until the differential shifted the other way and draining recommenced. Both CMTs were completely empty at about [ l seconds, which caincides closely with the primary
- sump injection valve opening at [ ]** seconds. Failure of the DAS from [ ]** seconds until the completion of the test stopped any further description of the test events. .
The break separator level began to increase at the same rate as the primary sump at about i [ ]'*d seconds. This occurred when sump level reached the height of the break separator loop seal. A As a result of this level increase, break separator level reached the heNht of the break in CL 3, causing break flow to reverse and flow from the break separator into tl.e RCS through the break at about [ ] seconds. The break flow then remained essentially zeto or slightly negative throughout the rest or the test. At approximately [ ]*' seconds, the PRHR inlet temperature started r. sing to the saturation temperature while the discharge temperature remained steady at approxirrately [ ]""F for the
- remainder of the data collection period.
i 5.2.1.5 Sump Indection Phase Primary sump injection (Figure 5.2.12) began through the check valves around the sump injection valves at about [ l'6# seconds, or some [ ]** seconds earlier in Matrix Test SB18 than in Matrix Test SB01. When sump injection began, the reactor vessel downcomer fluid temperatures rapidly increased to the sump flow injection temperature. Primary sump injection valves opened at [ l'6# seconds when the IRWST reached its low low level setpoint of [ ]** In., and the test was stopped 30 minutes later. Detailed data between [ 1** seconds and the end of the test is missing for one instrument rack and thus, this phase of the transient cannot be analyzed. O msp600s23m.52=tb.ioo395 5.2.1 5 REVISION: 1 I
-- _ _ _ ~ - - . . _ , - , - .
- - . - , . -- .-.mm- . - ,.. .c, ...-v -w. - , - - - , r , mr-*, .---r< - - - - - - -
1 l l TAllLE 5.2.1 1 OSU TFST ANAINSIS l' LOT PACKAGE FOR SUllSECTION 5.2.1 Plot No. Component Variables Units Description i Pressurizer CfT-604 pska System pressure and event history 2 Water WWTDVII+WWTDV12, Ibm /sec. Total of CMT, accumulator, IRWST, injection WOUTACCl+WOIJTACC2, and sump injection flows WWTIRWil+WWTIRWl2, WWTSMPIT 3 Reactor RPVASOU2 lbm/sec. Steam genera' ion in reactor vessel vessel 4 Reactor T08RPV, IITMXRPV, TS AT *F Reactor vessel outlet temperature, vessel maximum clad temperature and fuel exit saturation temperature O O m:vm2ms2.non:th-loons 5.2.1-6 REVISION: 1
. . - . . . . . . - . ~ . - ~ ~ . . _ - - . . - - . . . _ . . - - - - . - ~ _ - - - _ ~ _ . . . . . . _ . - .. ..~ .. ~ -
1 i
-l, THE FIGURES LISTED IN TABLE 5.2.1 1 -
' ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT
-1.
I l l l e f 4 i i t 4 s } ' I l ? a e i !- l I
- . m
- W344w-52.non:Ib-100395 521-7 REVISION: 1
!L l o i $ l i - - - . . - . . - - . _ ..~..._. _ _..,___ _ _,
5.2.2 Short. Term Transient - For the 2-in. cold leg break, Test SBl8 (a repeat of SB01), the short-term transient encompassed the first [' ] seconds. As shown in Figures 5.2.21 and 5.2.2 2, this period included the full } depressurization of the facility through all four stages of the ADS, together with CMT and - accumulator injection plus the inidal stages of IRWST injection. The mass and energy distribudon
. results for this phase of the transient were based on the plot package detailed in Table 5.2.21. Rese plots concentrate on the primary system, including the accumulators, CMTs, IRWST, and the mmps and the flows from the primary system via the ADS, break, and IRWST overflow.
5.2.2.1 Maintenance of Core Cooling
, Reactor Pressure Vessel and Downcomer Mass Distribution For the short term transient, the most important criteria is the maintenance of sufficient core inventory to supply adequate cooling of the heater rods. Figure 5.2.2-57 shows that there are no significant-excursions in heater rod temperatures and, therefore, sufficient core inventory was maintained through this phase of the transient to remove the decay heat from the rods. However, a two-phase mixture will be present in the core and upper plenum regions for significant portions of the transient. He following discussion tracks the variation in water level and mass throughout the RPV and downcomer.
he total fluid mass in the RPV is shown in Figure 5.2.2-40. De initial vessel inventory was [ ] ; lbm. During the course of the short-term transient, the vessel inventory experienced two minimum i I values, one of [ ]' lbm before accumulator injection and one of [ ]'6'lbm before IRWST injection. Steam generadon was near maximum at these times, as shown in Figures 5.2.2 55. By the end of the short-term transient, the vessel inventory recovered to a steady [ l# lbm. Similar i variations are shown in the core fluid mass and water level reproduced in Figures 5.2 2-44 and ) 5.2.2 45. Minimum core level occurred before IRWST injection. During this phase of the transient, the collapsed liquid level dropped to [ lin. below the top of the heated rod length (Figure 5.2.2-45). By the end of the short term transient, the effect of IRWST injection ended all core boiling (Figure 5.2.2 55), and the core was again water-solid. In the analysis of the SBLOCA simuladons on the SPES-2 facility
- following the pump trip, there were short term oscilladons in primary system flow, temperature, and pressure. A small number of oscillations were also observed in the OSU test response at the end of the initial blowdown (see for l example the pressure response (Figure 5.2.2-1) and the RPV mass (Figure 5.2.2-40)). In the SPES-2 I tests, the oscillations were driven by power-to-flow mismatches in the core due to the high core power levels that were needed in SPES-2 to compensate for the high ambient losses. The short-term i oscilladons observed in OSU results are believed to be a result of pressure oscillations following the initial blowdown.
!b
- J i
ms.p600(2h52..o :ib ioo395 5.2.2-1 REVISION: 1 1-i
--.a -. - - - - - . . . , , - ,, . . - - - . , , - ..- ,,, , - , . . . .
1 l 1 He fluid mm s in the core region is shown in Figure 5.2.2-44. Once again, the two minimum values occurred prior to accumulator injection and prior to IRWST injection. The niinimum core inventory is [ ]'6' lbm. The collapsed liquid level in the upper plenum region spanned by LDP-138 and the associated fluid mass are shown in Figures 5.2.2-49 and 5.2.2-48. During the period before accumulator injection, the collapsed liquid level in the upper plenum dropped below the hot leg elevation. During the accumulator injection, the steam bubble in the upper plenum partially condensed, and the water level briefly rose above the hot legs. Following the end of accumulator injection, the flow from the CMTs was not suff1cient to maintain the upper plenum level, and this region completely drained of water. He start of IRWST injection refilled the upper plenum, and this region became near water-solid again at approximately [ ] seconds. He fluid mass and collapsed liquid level for the head region are given in Figure 5.2.2-50 and 5.2.2-51. During the first [ ] seconds, the head inventory drained. Following the end of accumulator injection, the head region aga!n drained. Both accumulator injection and IRWST injection were sufficient to supply a level of water in the head. He mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.2.2-41 and 5.2.2-42. During the blowdown phase of the transient, the level dropped to the elevation of the cold legs. This elevation was maintained until the cold legs were fully drained. Following this time, the collapsed level remained between the DVI and hot-leg elevations, until IRWST injection once again raised the level above the cold legs and cold-leg refill commenced. Loop Mass Distribution For this discussion, the loop was considered to consist of the hot- and cold-leg pipe work, the SG primary side, and the pressurizer plus surge line. De total fluid mass and water level for the pressurizer are shown in Figures 5.2.2-34 and 5.2.2-35. During the blowdown phase of the transient ([ ]#) seconds, the pressurizer drained rapidly, becoming completely empty of water at about [ ] seconds. The pressurizer remained empty until ADS-1 actuation at [ ]'6' seconds. At that time, water was drawn back into the pressurizer as steam and water flow out of the ADS. A fluid inventory of over [ ]'6* lbm was maintained until ADS-4 actuation at [ ] seconds. This caused an initial outsurge through the surge line, followed by a more gradual draining of the pressurizer as mass flowed out of the hot legs via the ADS-4 valves. The pressurizer fully drained at [ ]'6# seconds and remained empty for the remainder of the transient. Mass data for the SG U-tubes and their associated inlet and outlet plena are shown in Figures 5.2.2-32 and 5.2.2-31. He SG tubes gradually drained until ADS actuation, when all the tubes and plena were empty of water. The SG on the broken loop (loop 1) drained before that on loop 2. Any flow m%sxx2344w-52.nonsioo395 5.2.2-2 REVISION: 1
- -- -. . _ _ -. --. . . . = - - - _ - . . _ . .
1 I I l through the SGs ceased onu the tubes were drained and the steam trapped within the U-tubes became superheated. Once the SG tubes drained, natural circulation around the primary loop circuit ceased. 1 De SG U-tubes remained empty for the remainder of the short-term transient. However, both inlet and outlet plena on SG 2 show an influx of water [ ]"# seconds into the transient. This corresponds to the time at which the primary system reached atmospheric pressure. Figure 5.2.2-59 shows that in approximately [ -]"# seconds, the small amount of steam remaining in HL 2 was removed, and the water level in that hot leg increased. His is believed to result from condensation in the PRHR drawing a vacuum and raising the level on loop 2. De mass of water and vapor in the hot legs is shown in Figures 5.2.2 58 and 5.2.2 59. He hot legs maintained their water inventory until [ ]*# seconds into the transient, when they started to drain (Figure 5.2.2-58), and were completely drained by [ ]"# seconds. Actuation of ADS-1 caused a rapid increase in void fraction in both hot legs. A larger void fraction was maintained in HL-1 as steam was preferentially removed from HL-2 by the PRHR. He liquid and vapor mass for the four cold legs are shown in Figures 5.2.2 60 and 5.2.2-61. Following the initial blowdown, all four cold legs became two-phase although there was a greater void
]
fraction in CL-1 and CL-4 as compared to CL-2 and CL-3 (Figure 5.2.2-61). The mass variation in all l four cold legs appears very similar, as these have been derived from levels in the RPV downcomer. I p Derived values were used because the level instruments on the cold legs have been found to be d unreliable. All the cold legs completely drained after ADS initiation and refilled at [ ]*# seconds when flow from the IRWST refilled the RPV downcomer to the level of the cold legs (Figure l' 5.2.2-42). The cold legs did not drain uniformly. Instead, CL-3 (with the break) drained first, followed by CL-1, with CL-2 and CL-4 delayed. Figure 5.2.2-31 shows that the SG outlet plenum on loop 1 drains at [ ]*# seconds slightly more than and before loop 2; therefore, the expected asymmetry in cold-leg behavior did occur. Mass Injected to the Primary System The CMTs transitioned from a recirculation to a mass injection mode at approximately [ ]"# seconds when the cold leg started to drain. Draindown of the CMTs continued until the CMT check valves were closed by the flow from the accumulators. CMT draindown restarted at the end of accumulator injection, continuing until IRWST injection began (Figures 5.2.2-6 and 5.2.2.-7). The accumulators started draining approximately 20 seconds after activation of the ADS. Accumulators started discharging into the DVI line when the system pressure dropped below the I pressure preset in each accumulator. Accumulator injection began at approximately [ ]*# seconds and continued until the accumulator was empty at approximately [ ]*# seconds (Figure 5.2.2-23). Complete discharge from the accumulators was indicated by a sharp decrease in the temperature of the _ , fluid exiting each accumulatcr due to the discharge of expansion-cooled, nitrogen-cover gas, which ms.p600c3w-524on:Ib-too395 5.2.2-3 REVISION: 1
l l l l was released into the primary system (see Subsection 6.1.4). Flow from the CMTs was significantly reduced during the discharge of the accumulators and increased once accumulator discharge ended. De IRWST injecdon valves were opened when the reactor vessel pressure low-low setpoint was reached. Injection flow started when the reactor vessel pressure became less than the static head from the IRWST. Figure 5.2.2-16 shows that the IRWST injection began at approximately [ ] seconds after the CMT flow ceased. He IRWST flow gradually increased to a peak value of[ ]*' lbm/sec. ([ ]*' lbm/sec. per injecdon line) at [ ] seconds before gradually decreasing. Mass Ejected from the Primary System At time zero in the transient, a 2-in. break was initiated at the bottom of CL-3. He rate of flow of mass out of the primary system, via the break, is shown in Figures 5.2.2-66 to 5.2.2 68. For the first [ ]'6" seconds following the break, [ ]"# lbm of steam and water escaped from the primary system via the break (Figure 5.2.2-62). During that period, the primary system depressurized to around [ ]"' psi (Figure 5.2.2-1). By the onset of ADS actuation, the cold legs drained, and there was almost no water flow out of the break.13etween [ ] and [ J'6' seconds, ADS 13 activated, and the system depressurized rapidly. We break flow significantly decreased once ADS activated, since the ADS valve area is significantly larger than the break. At around [ ] seconds, the ADS-4 was inidated, and the primary system continued to depressurize until IRWST injection commenced at [ ]"' seconds. The actuation of ADS 1-3 terminated the flow of steam from the break, although this was replaced by steam flow through the ADS 13 valves for the next [ ]*6* seconds (Figure 5.2.2-63). This steam flow was accompanied by an outflow of water from the ADS 1-3 at a peak rate of over [ ]lbm/sec. (Figure 5.2.2-66). After [ ]"' seconds the mass flowing through ADS 13 was composed almost endrely of water. The rate of flow through the ADS continued at a gradually reducing rate until [ ] seconds when the ADS-4 valves opened, terminating flow through the ADS l 3 and replacing this flow with flow through the lower-resistance ADS-4 paths. For ADS 4-1, there was a near-steady water flow at a rate of [ ]"' lbm/sec. from the time of initiadon.110 wever, there was an initial outsurge at [ ]"' lbm/sec. for ADS 4-2, which was followed by a drop to near zero and then an increase to over [ ]*' lbm/sec. (Figure 5.2.2-66). The integrated mass flow out of the primary system via the ADS and the break are shown in Figures 5.2.2-62 to 5.2.2-66. During the first [ ]"' seconds of the transient, over [ l'6' lbm of water escaped from the primary system. Of this, the [ ]*'lbm flowing through ADS 1-3 was deposited in the IRWST. The [ ]"'lbm leaving ADS-4 and the liquid part of the [ ] lbm flowing through the break were added to the overflow from the IRWST to deposit [ ]'6# lbm of water in the primary sump (Figure 5.2.2-28). By the end of the short-term transient, the water level in the primary sump rexhed nearly [ ]*' in. (Figure 5.2.2-29). O mMr6002h52.non:Ib loo 395 5.2.2-4 REVtSION: 1
- -.- - ._- - -- - ~ - - - - ---. - . - - - - - - . ~ . - - . . - -
l l At [ l'** seconds into the transient, the cold legs refilled enough to restart water flow through the break. His proceeded at a rate of approximately [ J"'lbm/sec. at the end of the short term transient, so that the total rate of water flow fmm the primary system to the sump was approximately [ j lbm/sec. At that time, the llow rate into the RPV through the DVI lines was about [ ] Ibm /sec. Mass Italance Figure 5.2.2-70 presents the variadon in the total system inventory during the short term phase of the transient. In addidon to the random variations associated with the measurement uncertainties and some inventory dips, there is a general reduedon in inventory of around [ l lbm from the l initial value. During this phase of the transient, around [ }'6'lbm of steam was lost from the system (Figure 5.2.3 63). A mass balance analysis was performed on the primary system. Figure 5.2.2 71 plots the measured primary system mass determined by summing the contributions from the RPV, downcomer, hot and ccid legs, SG primary, pressurizer, and surge line plus the PR11R. The second curve on Figure 5.2.2-71 provides an alternative primary system mass determined from the mass balance, diat is l given by: 1 M$,(t) = My ,(0) + M,(t) - M,(t) 5.2.21 where: M'g,(t) = Mass balance calculated primary system mass Mg ,(0) = Measured primary system mass at the start of the transient M,(t) = Total integrated mass injected from all sources (i.e., accumulators, CMT, IRWST, and sumps) to time t M.(t) = Mass lost from the primary system to Ome t via the CMT balance lines, ADS 13, ADS-4, and break
%e difference in the two primary system mass curves is shown in Figure 5.2.2 72 as the mass-balance error.
During the short term transient, there is an apparent systematic overestimate of the mass in die primary system from the measured data of up to [ l lbm, relative to that calculated from the mass balance. Here are two main contributions to this excess mass. First, the measured primary system inventory does not include all of the pipework in the system and there is about [ ] lbm of mass missing from the inidal inventory. His mass will be deposited in, and lost from, the measured system as pipes drain and refill. Some of the additional mass is subsequently lost via one of the leakage ( paths. Second, the instrumentation on the hot legs may be giving erroneous level measurements ( during certain portions of the transient. Figure 5.2.2 58 shows that neither hot leg appears to drain in i mAapoocmw-52.non:ib-too395 5.2.2-5 REVISION: 1
l this test, which contributes to an overestimate of some [ j' lbm. Dy the end of the short term transient, the apparent mass-balance error is about [ )* lbm. Figure 5.2.2 73 shows the total integrated mass How from and to the primary system, together with the water inventory remaining in the sources of cooling water. During the short term transient, there was a net loss of water from the primary system of approximately [ J lbm, of which only a small quantity was deposited in the sources. As noted above, the overall system mass inventory shows that of the lost primary system mass,[ ] Ibm has been lost as steam. The rest was added to the water stored in the ADS and break separators. Pressure Decay Figure 5.2.21 shows the primary system pressure during the test.11troughout the LOCA portion of this test, the pressure was controlled by the saturadon pressure of the hottest fluid in the primary system. At inidation of the break, the controlung fluid volume was the pressurizer and surge line; however, within the first [ J seconds, (after the inidal blowdown phase) this shifted to the RPV. Figure 5.2.2 3 shows that the average temperaturi of the upper plenum was equal to the saturation temperature corresponding to the primary system pressure measured in the upper head during the natural circulation phase and into the ADS phase. The pressure stabilized at the saturation pressure for the upper plenum, and then continued a slow pressure decay responding to the cooung caused by the CMT injecdon. Figure 5.2.21 shows an increase in the pressure decay rate occurred at about [ l seconds, when the CMTs transitioned from natural circulation injection to draindown injecdon, which essendally doubled the injection rate of cold water into the DVI line. The tugher injection rate resulted in a rapid temperature drop in die upper plenum (core outlet in Figure 5.2.2 3), which was reflected in a more rapid pressure decay. With the actuation of ADS 1 at about [ ] seconds, the pressure dropped rapidly due to the ! increased rate of mass ejected from the system (Figure 5.2.2 56), and the increased flow of cold water l injected into the downcomer and flowing through the core. This condnued to reduce power channel I inlet plenum temperature, and subcooled the heater rods in the core due to the higher flow. Since the RPV outlet plenum became subcooled at about [ jd seconds, the hottest fluid in the system was in the pressurizer, the cold legs, and the CMTs, and the pressure was partially supported l by the flashing of the fluid in one or several of these locadons. When the accumulator discharge ended (about [ l' seconds), the RPV temperature again increased to the saturadon temperature and took control of the system pressure for the rest of the LOCA phase. Also at the end of accumulator injection, a large amount of noncondensible gas was injected into the primary system and could have affected the heat transfer performance of the PRilR and the CMTs. O maar60m2%.52.noo: b-ioons 5.2.2-6 REV!slON: 1
q Q 5.2.2.2 Energy Inventory lleat removal from die reactor core follows a sequence similar to the pressure decay for the SBLOCA tests. Before reactor trip, nearly all the energy generated in the core is removed by the SGs and out of the break with a small fraction lost to the surroundings. When the reactor tripped, the primary system pumps tripped, and flow through the SG tubes was sharply reduced, Coupled with the Isoladon of the 50 secondary side, the result was to significantly reduce heat removal by the SGs. At that Ome, the PRIIR isolation valves opened, and energy was removed to the IRWST as well as out of the break. As the system drained, the primary system pressure was reduced, and the sensible heai of the coolant and metal added to the core heat load. The CMTs started to drain, and the ADS activated. At diat time, heat removal was accompushed through the ADS flow, and the PRilR became less effective. Finally, ADS 4 actuated, the primary system completely depressurized, and the IRWST actuated. The LOCA phase of die test was then completed. 1hc behavior of die components involved in the energy removal is discussed below. Core The power output of the core is shown in Figure 5.2.2 2. After reactor trip, the core power is C k representative of decay heat levels expected in the AP600 core. Flow through the core is sNwn in Figure 5.2.2 56, and the steam generation rate is given in Figure 5.2.2 55. As discussed in l Section 4.0, the steam generation rate can be calculated by two methods, the Tsat method and the i DVI line flow method. l Figures 5.2.2 53,5.2.2 54, and 5.2.2 55 reproduce the saturation line elevation, power split above die saturation elevation and steam generation rates from the two methods. Both methods give similar predictions. Note that neither method gives valid predictions before [ jdd seconds because of flow oscillations during natural circulation in the primary system. l The maximum steam generation rate during the LOCA phase occurred at approximately [ ]*' seconds, which was just prior to IRWST injection. The peak cladding temperature is shown in Figure 5.2.2 57, and indicates that the core was adequately cooled at all times during the test. Steam Generator lleat Transfer The SGs remove most of the heat from the primary system daring normal operation. Ilowever, heat transfer from the primary to secondary side was significantiy reduced after the pumps trip. This was due to reduced flow in the tubes, which caused a sharp reduction in the tube-side heat transfer coefficient. In addition, the secondary side was isolatad, which caused the temperature and pressure to remain high as the primary side pressure rapidly decreased. v mNr6000h12 ace: b.too395 5.2.27 REVISION: 1
l Passive Residual llent Removal lleat Transfer 1 He PRilR is designed to remove heat from the primary system from the Ome when the SGs become I thermally isolated due to the inidation of the ADS. One measure of the effectiveness of the PRilR is the increase in the fluid internal energy in the IRWST, widch serves as the heat sink for the PRilR. Figure 5.2.2 33 shows the fluid internal energy in the IRWST during the first [ ]** seconds of the test. De PRilR began operating at about [ l'6' seconds as the SO heat removal ended. De heat removal rate was approximately [ ]** Btu /sec. until [ )** seconds when ADS-1 activated. At that time, the headng rate in the IRWST increased to [ ]** Bru/sec. due to steam condensing from the ADS 13, and heating from the PRilR. De PRilR heat removal decreased when the accumulators discharged at about [ }** seconds as the core became subcooled. Automatic Depressurization System and fireak
%e ene.gy removal from the ADS and break are shown in Figure 5.2.2-69. ne fluid e.vrgy exiting the break increased at a constant rate until ADS 1 was actuated. For ADS 1, ADS 2, and ADS-3, compared to the break, the energy removal occurred at a somewhat lower rate due to reduced system pressure. When ADS-4 was actuated, energy removal switched from ADS l 3 to the larger flow path ADS 4. He ADS cffectively reduced the primary system pressure to start gravity injection flow from the IRWST. His flow was sufficient to subcool the primary system, ending core boiling and partially collapsing the steam bubble in the upper plenum to bring the system to near atmospheric pressure.
Overall Energy linlance Figure 5.2.2 74 shows all the energy components in the heat balance for tir system during the LOCA phase ([ ]** seconds). Broughout the event, the heater rod bundle power was the dominant heat input to the system, and before the start of the event, the SGs provided the dominant heat removal. At the start of the event, the SO secondary side was isolated, and the RCPs tripped. During this period, heat removal by the SG was reduced as natural circulation flow occurred in the primary system. As the primary system depressurized, the pressure reached the secondary side pressure, and heat transfer effectively ended. From the start of the event, the steam component of the break flow left the control volume until the actuation of ADS. IIcat loss via this path was nearly [ )** Blu/sec., which was far greater than die reactor decay power [ ]** Btu /sec. %e remaining energy lost through the break exhaust consisted of a decrease in the fluid internal energy as the primary system depressurized. After ADS 13 was actuated, the break flow effectively ended, and the ADS heat was deposited into the IRWST Rus, the fluid internal energy in the control volume increased until the end of the LOCA phase, as the IRWST water temperature increased. Also, at this time, the metal masses in the cor. trol volume lost energy as primary system temperature decreased. As a result, energy was deposited at a rate of nearly [ ]*' Btu /sec. between [ ]** seconds. After [ ]** mvm23hs2.nc :n>.too395 5.2.2-8 REVISION: 1
(' seconds, the metal masses lost energy at a much lower rate, which was consistent with the primary system fluid temperature decay. Ambient losses were reduced from a maximum of [ J'6# Bru/sec. at full-power conditions tc [ ] Btu /sec. at the end of the LOCA phase. A large increase was observed in the deficit between the rod bundle power and the various sources of energy dissipation from the control volume after the initiation of ADS 1-3, and again after the initiation of ADS-4 at approximately [ ]'6' seconds. ' Die steam exhaust included steam from the IRWST and the steam portion of the flow from ADS-4. However, relatively little steam flow was measured by the vapor flow meters. Further discussion of this steam flow is provided in the mass and energy balance in Sections 6.2.2 and 6.2.3. ( v (/ m:wwe2344w.52.non:isioo395 5.2.2-9 REVISION: 1
TABLE 5.2.2-1 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.2.2 Plot No. Component Variables Units Description 1 Pressurizer CI*T-604 psia System pressure l 2 RPV RPVPWR kW Core power 3 RPV TOIRPV, TD8RPV, 'F Core inlet / outlet temperature, ST08RPV saturation temperature 4 SG CI"T-201, CPT-204, psia Prunary and secondary pressures in SG CPT-301, CPT-302 5 D VI-l WWTDVIL1, Ibm /sec. Individual components and total flow in WWTIRWII, D VI-l WOUTACCl, WWTIRWI3 6 DVI-2 WWTDVIL2, Ibm /sec. Individual components and total flow in W%TIRWI2, D VI-2 WOUTACC2, WWTIRWI4 7 CMT AMCMTIB, Ibm Fluid mass in CMTs (excludes balance AMCMT2B lines) 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT MIWDVIL1, Ibm Integrated mass out of ChfTs MIWDVIL2 10 CMT WWTDVIL1, Ibm /sec. Flow out of Chfrs WWTVIL2 11 Chfr WWTCLBLI, Ibm /sec. Flow into CMTs WW'fCLBL2 12 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 13 CMT UCMT1, UCMT2 Btu Fluid energy in CMTs 14 IRWST IRWST Ibm Mass of fluid in IRWST 15 IRWST CLDP-701 in. Collapsed liquid level in IRWST 16 IRWST WWIIRWII, Ibm /sec. Flow from IRWST to DVI lines WATIRW12 17 IRWST IRWSTOR Ibm /sec. Overflow from IRWST to sump 18 IRWST ADS 13TMR lbm/sec. Total ADS flow into IRWST 19 IRWST ADS 13TIR, MIIRWII, Ibm Integrated mass out of IRWST MIIRW12, MilRWIO 20 IRWST UIRWST Btu Fluid energy in IRWST 21 PRHR CLDP-802 in. Collapsed liquid level in PRHR HX O: mAap6002344w-52. nan.lb-too395 5.2.2-10 REVISION: 1 1
TABLE 5.2.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION S.2.2 Plot No. Component Variables Units I)escription 22 PRHR WWOTPRHR lbm/sec. Measured outlet flow from PRHR tube 23 Accumulator AMACCl, AMACC2 lbm Mass of fluid in accumulators 24 Accumulator CLDP-401, CLDP-402 in. Collapsed liquid level in accumulators 25 Accumulator WOUTACCl, Ibm /sec. Flow from accumulators l WOUTACC2 s 26 Accumulator MOUTACCl, Ibm Integrated mass out of accumulators MOUTACC2 27 Accumulator UACCl, UACC2 Btu Fluid energy in accumulators 28 Pnmary sump AMPSMP lbm Pnmary sump lluid mass ' 29 Pnmary sump CLDP-901 in. Pnmary sump level 30 Pnmary sump UPSMP Btu Primary sump lluid energy 31 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG primary side MSSGOP1, MSSGOP2 inlet / outlet plena 32 SG MSSGHT1, MSSGHT2, Ibm Mass of fluid in SG primary side hot l MSSGCTI, MSSGCT2 and cold tubes i 33 SG/PRHR CFT-201, CIT-301, psia & SGI pressure and PRHR integrated beat QPRHRI Btu output 34 Pressurizer PZM lbm Fluid mass in pressurizer 35 Pressurizer CLDP-601 in. Collapsed liquid level in pressurizer 36 Pressurizer UPZ Btu Fluid energy in pressurizer 37 Surge line PLM lbm Fluid mass in surge line 38 Surge line CLDP-602 in. Collapsed liquid level in surge line 39 Surge line UPSL Btu Fluid energy in surge line 40 RPV MWRPV lbm Total fluid mass in reactor vessel 41 RPV DCM lbm Fluid mass in downcomer 42 RPV LDP01DC in. Collapsed liquid level in downcomer compared to various reference elevations 43 RPV MWOIRPV lbm Fluid mass in lower plenum j 44 RPV MWO3RPV lbm Fluid mass in core region 45 RPV LDP03RPV in. Collapsed liquid level in core 46 RPV RPVAVDF2 Core exit void fraction 47 RPV RPVAQOU2 Core exit quality 48 RPV MWO6RPV lbm Fluid mass in the upper plenum l 49 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 50 RPV MWO8RPV lbm Fluid mass in the upper head , 51 RPV LDP08RPV in. Collapsed liquid level in the upper head v , m:Wm.52..on:ib-ioo395 5.2.2-11 REVISION: 1
TABLE 5.2.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.2.2 Plot No. Component Variables Units Description 52 RPV URPV Btu Total fluid energy in reactor vessel 53 RPV RPVXE, RPVASL2 ft Level of Tsat line 54 RPV RPVPab, RPVAPab2, kW Heated rod power above and below RPVPWR Tsat level and total 55 RPV RPVRXV,RPVASOU2 lbm/sec. Core steam generation rate 56 RPV RPVALIN2 lbm/sec. Calculated core flow 57 RPV IITMXRPV, S'II)8RPV 'F Maximum clad temperature and saturation temperature 58 Hot leg MWHL1, MWHL2 lbm Water mass in bot legs 59 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs 60 Cold leg CLIWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 61 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 62 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4s, and break 63 ADS and break BRKTIVF, AD13TIVF, Ibm Totalintegrated vapor flow for ADS AD41TIVF, AD42TIVF and break 64 ADS and break BRK11LF, ADl3TILF, Ibm Total integrated liquid flow for ADS AD41TILF, AD42TILF and break 65 ADS and break ADS 13SVR, Ibm /sec. Vapor flow out ADS 1-3 and ADS-4 ADS 41SVR, ADS 42SVR 66 ADS and break ADS 13SLR, Ibm /sec. Liquid flow out ADS 1-3 and ADS-4 ADS 41SLR, ADS 42SLR 67 ADS and break BRKSSVR Ibm /sec. Vapor flow out of break 68 ADS and break BRKSSLR Ibm /sec. Liquid flow out of break 69 ADS and break BRKSPEI, ADS 13EI, Btu Integrated fluid energy for ADS 1-3, ADS 41EI, ADS 42El ADS-4, and break 70 Mass balance TOTMASS lbm Total system mass inventory 71 Mass balance PRIMMASS, Ibm Measured primary system inventory and PRIMASS2 value from mass balance 72 Mass balance MERROR lbm Mass balance error 73 Mass balance MIN, MOUT lbm Integrated mass flow hi and out of SRCMASS primary system and source mass 74 Energy balance Various Btu Components of energy balance O mMp60m2%-52.aon:is tom 95 5.2.2-12 REVISION: 1
s. 1. 1 l' 1 4 i 1 4
);
1 1 4' i' a 1 4 i
- THE FIGURES LISTED IN TABLE 5.2.21 j
4 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT i t I ( l d- - t l t a 1 1 1 I l8 l i 1 l i i ) 4 i i 4 i ! } i ( l
]
EW344w 52.non:Ib-100995 5.2.2-13 REVISION: 1
- . . . . ~...---- _.-.- _.-.-. -._.-..-.._ - -.- _. -
Y
'l l
f 5.2.3 Long-Term Transient - i l De long-term transient covered the transition from IRWST to sump injection and provided information on the LTC response of AP600. For the 2-in. cold-leg break, Test SB18, the DAS did not record all the instnamentation signals beyond [ ]'6# seconds; therefore, the sump injection phase
- cannot be analyzed beyond this point. In this case, the long-term transient covered the time frame
- from [ ]'6# seconds. By this time, the initial stages of flow from the primary sump had !
J l begun via the check valves around the main injection valves, but the main sump injection valves had j not opened. The behavior of the test facility during this period of the transient is discussed in this j subsection using the plot package detailed in Table 5.2.31. These results concentrate on the l components of the primary system that remain active during the LTC phase, that is the RPV, the hot i legs, ADS-4, the sumps, and the IRWST. l 1
- 5.2.3.1 Maintenance of Core Cooling i
Reactor Pressure Vessel and Downcomer Mass Distribution i For the long-term transient, the passive core cooling systems must supply sufficient flow to prevent any overheating of the heater rods. At [ ]'6* seconds, the decay heat simulation of the heated rods reduced the power from [ ]'6# kW (Figure 5.2.3-1). As shown in
""'~'*'"*'*""""'""'""""*"'*"'"'**"*'"'""""'"*'*'"'"'
CJ core flow was maintained throughout the long-term transient. De mass of water in the RPV is shown in Figure 5.2.3-25. After an initial decline, the reactor vessel j water mass settled at an average value of around [ ]'6# lbm where it remained until the end of the l test. From [ l'6# seconds, oscillations in vessel inventory can be observed. Rese oscillations are evident in measurements throughout the primary system and are discussed further in Subsection 6.1.3. The core water mass and collapsed liquid level are shown in Figures 5.2.3-28 and 5.2.3-29. During the transient, the core remained nearly water-solid with only a low level of boiling (Figure 5.2.3-36). The level at which the core fluid reached saturation temperature was around [ ]'6# in, for the period of the transient (Figure 5.2.3-34). Figure 5.2.3-29 shows that at the end of the transient, the effect of the hot water arriving from the sump was detected as a decreased core collapsed liquid level. The collapsed liquid level in the upper plenum region is shown in Figure 5.2.3-32. During the period before sump injection began, the collapsed liquid level in the upper plenum remained at the top of the hot legs. Following the start of the hot water influx from the sumps, the level dropped to the middle of the hot legs. O ms.puerm.52..on:ib. loo 395 5.2.3-1 REVISION: 1 l
The mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.2.3-26 and l 5.2.3-27. Before sump injection began, the collapsed liquid level in the downcomer was at the level of the center of the cold legs. By the end of the transient, the effect of the initial stages of sump injection reduced the downcomer collapsed liquid level to the hot-leg elevation. I Loop Mass Distribution As discussed, the LTC phase of the transient shows that there is a low level of boiling in the upper l regions of the core. However, due to a failure of the DAS, analysis of the long-term transient is not detailed enough to track the behavior of this steam. Mass Ejected from the Primary System i i The integrated mass flow out of the primary system via the ADS and the break is shown in Figure 5.2.3-43. By the end of the transient, [ ]'6'lbm of water flowed out of the primary f system. During the LTC phase of the transient, the only significant outflow was through the ADS-4 valves with a small apparent flow through the break. The apparent break flow did not represent a flow out of the primary system, but indicated continued interaction between the break separator and the sump. The most marked manifestation of this interaction is at [ ]'6" seconds when the primary sump began to overflow into the secondary sump. This caused oscillating flow indications .in the liquid flow out of the break separator. This is confirmed by Figures 5.2.3-44 to 5.2.3-45, which
)
show the flows through the ADS and the break. During the sump injection phase of the transient, outflow was in the form of liquid out of the ADS-4 valves. Water flowed through each of these at an average rate of [ ]'6# lbm/sec. Although no steam flow was recorded by the vortex meters to show steam escaping from the system via ADS-4, as discussed in Subsection 6.2.2, steam most likely left the system by this route at a very low flow rate. At approximately [ ]'b' seconds into the transient, the level in the primary sump (Figure 5.2.314) reached the point at which overflow to the secondary sump occurred. At that time, l there was [ ]'6"lbm of water collected in the primary sump. From the beginning of primary sump overflow to the end of the analysis, [ ]'6* lbni of water was transferred to the secondary sump. Mass Ir.jected to the Primary System The total DVI line flow, CMT flow, and IRWST flows are shown in Figures 5.2.3-6 and 5.2.3-7, and the flow from the primary sump is shown in Figure 5.2.3-19. From around [ ]'6' I seconds, there was a contribution to the DVI flow from the CMTs as they finished their post-refill l draindown. l l 9 maap600c3m.52.noo:isioo395 5.2.3-2 REVISION: 1 l l
i During the presump injection phase of the transient, the IRWST flow proceeded at a gradually reduced rate, with the effect of the primary system oscillations superimposed. At [ ] seconds, flow
~ from the primary sump began through the check valves around the main injection valves. At the end of the analysis, the level in the IRWST fell to [ ~ ]** in., which was above the level at which the primary sump injection valves opened.
Mass Balance t Figure 5.2.3-46 shows the variation in the total system mass inventory during the entirety of Test SB18. Following the short-term transient, total inventory increased by [ ]** lbm to approximately [ ]** IM above the initial value at the time sump injection around the main valve started. The initiation of sump injection led to [ ] lbm increase in inventory. He mass balance calculation described in Subsection 5.2.2.1 was performed for the entirety of the transient, and the results are presented in Figures 5.2.3-47 to 5.2.3-49. From [ l'6' seconds, the error remained between 2 [ ]'6'lbm. Following the overflow from the primary to secondary sump, the error increased due to the interaction between the break separator and sumps. 5.2.3.2 Energy Balance 1 (
' ~
Figure 5.2.3-50 shows all the energy components in the heat balance for the system during the LK , phase. The LK phase for this test was abbreviated because the data acquisition system stopped recording data prematurely. During this phase, the heater rod bundle power was the dominant heat input to the system. De SG heat transfer ended during the LOCA phase and did not contribute to the overall energy balance during the LK phase. Thus, for the LTC phase, the active components in the overall energy balance were the rod bundle power, :he change in the fluid internal energy, the change { in the metal internal energy, and ambient losses, and the steam exhausted from the control volume, he fluid energy in the control volume increased steadily until sump injection began (approximately [ ]** seconds). At that time, the fluid throughout the system approached saturated conditions, and the rate of increase was lower. At the same time, the metal mass temperatures increased as the primary system temperature increased. Also, the ambient losses increased slightly as the control volume temperatures increased. He steam component of the ADS-4 flow left the control volume after the actuation of ADS-4. For times subsequent to the LOCA phase, this quantity was essentially zero, as measured by the vapor flow meters. h is concluded that the vapor flow meters in the ADS-4 separators did not measure the relatively small steam flow rates accurately Further discussion of this steam flow is provided in the mass and energy balance in Sections 6.2.2 and 6.2.3. O m%um3m-52.n=Itvioo395 5.2.3-3 REVislON: 1
l TABLE 5.2.31 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.2 3 9iI 1 LONG-TERM TRANSIFET Plot No. Component Variables Units Description 1 RPV RPVPWR kW Core power 1 2 Pnmary sump TSMPI), TSMP12 'F Sump injection line temperatures l 3 DVI TDVILI, TDVIL2 'F DVI line temperatures ! 4 RPV T01DC,11)2DC, T03DC, 'F Water and saturation temperatures in STOIDC downcomer 5 RPV 1Y)!RPV, T08RPV, 'F Core inlet / outlet temperature, ST08RPV saturation temperature 6 DVI-l WWlDVIL1, Ibm /sec. Individual components and total flow W%TIRWII, in DVI-I W%TIRWI3 7 DVI-2 WWTDVIL2, Ibm /sec. Individual components and total flow { WW'11RW12, in DVI-2 WWTIRWI4 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs j 9 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 10 IRWST IRWST Ibm Mass of fluid in IRWST I1 IRWST CLDP-701 in. Collapsed liquid level in IRWST 12 IRWST UIRWST Btu Fluid energy in IRWST 13 Pnmary sump AMPSMP lbm Primary sump fluid mass I 14 Pnmary sump CLDP-901 in. Pnmary sump level l 15 Pnmary sump UPSMP Btu Pnmary sump fluid energy 16 Secondary sump AMSSMP lbm Secondary sump fluid mass 17 Secondary sump CLDP-902 in. Secondary sump level 18 Secondary sump USSMP Btu Secondary sump fluid energy l 19 Pnmary sump WSTSMPET, WWTSMPIT Ibm /sec. Pnmary sump steam and liquid { injection rate ! l 20 Primary sump MISMPII, MISMPI2, Ibm Integrated primary sump and IRWST MISMPIT, MIIRWT flows l 21 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG side inlet / outlet MSSGOPI, MSSGOP2 plena 22 Surge line PLM lbm Fluid mass in surge line 23 Surge line CLDP-602 in. Collapsed liquid level in surge line 24 Surge line UPSL Btu Fluid energy in surge line 25 RPV MWRPV lbm Total fluid mass in reactor vessel O. 1 mW344w-52.noo:Istoo395 5.2.3-4 REVISION: } l l
\
l l
l l l
/m TABLE 5.2.31 (Continued)
OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.2J LONG TERM TRANSIENT ' Plot No. Component Variables Units Description 26 RPV DCM lbm Fluid mass in downcomer j 27 RPV LDP01DC in. Collapsed liquid level in downcomer compared to various reference l clevations 28 RPV MWO3RPV lbm Fluid mass in core region l 29 RPV LDP03RPV in. Collapsed liquid level in core 30 RPV RPVAVDF2 Core exit void fraction 31 RPV RPVAQOU2 Core exit quality 32 RPV LDP06RPV in. Collapsed liquid level in the upper plenum . 33 RPV MWO8RPV lbm Fluid mass in the upper bead , l 34 PSV RPVASL2 ft. Level of Tsat line 35 RPV RPVAPab2, RPVPWR kW Heated rod power above and below Tsat level and total 36 RPV RPVASOU2 lbm/sec. Core steam generation rate 37 RPV RPVALIN2 lbm/sec. Calculated core flow 38 RPV HTMXRPV, 'F Maximum clad temperature, saturation (% 4 ST08RPV temperature and delta
/
'~'
39 Hot leg MWHL1, MWHL2 lbm Water mass in bot legs 40 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs 41 Cold leg CLlWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 42 Cold leg CL1VMS, CL2VMS, Ibm Vapor mass in cold legs ! CL3VMS, CL4VMS
; 43 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS 4, and break 44 ADS and break ADS 13TLR, ADS 41TLR, Ibm /sec. Liquid flow out ADS 1-3 and ADS-4 ADS 42TLR 45 ADS and break BRKSTLR lbmhec. Liquid flow and total flow out of break 46 Mass balance TOTMASS lbm Total system mass inventory 47 Mass balance PRIMMASS, PRIMASS2 lbm Measured pnmary system inventory and valve from mass balances 4
48 Mass balance MERROR Ibm Mass balance error 49 Mass bahmce MIN, MOUT SRCMASS lbm Integrated mass flow in and out of primary system and source mass 50 Energy balance Various Blu Component of energy balance 51 ADS-4 ADS 41TLR, ADS 42TLR lbm/sec. Oscillations in ADS-4 liquid flow
,o 52 Surge line CLDP-602 in. Oscillations in surgeline level e s
'v/ 53 RPV Cf'T-107 psia Oscillations in upper bead pressure mAap600(2344w.52.non:Itw100395 REVISION: 1 5.2.3-5
1 i i TABLE 5.2.34 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.2.3 9j ; LONG TERM TRANSIENT l 1 Plot No. Component Variables Units a .escription 54 RPV CLDP-il3 in. Oscillations in upper plenum level ! 55 RPV LDP03RPV in. Oscillations in core level j 56 RPV LDP01DC in. Oscillations in downccmer level l l i l l 1 I Ol ' O m:W3m-52.non:t b-too395 5.2.3-6 REVISION: 1
_ _ . _ . . _ _ _ _ . _ . . . _ _ _ _ . _ . . - _ __ __ _ - _ . . . . - . ~ . _ _ . - - - _ _ . _ _ . . _ - - - - - - . _ _ _ _ _ . 2 e s THE FIGURES LISTED IN TABLE 5.23-1 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT u 1 A O i i I J 1 1- i i J i I 1 l mSp60023w-52. on:1 Moo 395 5.23-7 REVISION: }
.. _. _ _ _ . . _ ~
l 5.3 Analysis of Matrix Test SB06 Matrix Test SB06 (OSU Test U0006) simulated a 4-in. cold-leg break SBLOCA with LTC and without the operation of the nonsafety-related systems. The break was located on the bottom of CL-3 and except for the break size, this test was similar to SB01, including the simulated failure of one of the ADS-4 lines. The analysis of Matrix Test SB06 is divided into three subsections as follows:
- Facility performance is discussed in Subsection 5.3.1, which provides a brief outline of the response of the test facility.
=
The short term transient for SB06 encompassed the start of the simulation up to [ ]'6# seconds. This period includes blowdown, natural circulation, ADS, and initial IRWST stages of the transient.
. The analysis of the long-term transient for SB06 encompassed the time frame from 1500 seconds to the end of the test. This phase of the transient includes the IRWST injection and covers the transition to sump injection. The long-term transient actually started at IRWST i injection, which is discussed as part of the short-term transient and so, the discussion of the long-term transient provided here begins at [ ]'6' seconds. At [ ] seconds, CMT-1 began to refill and CMT-2 followed [ ] seconds later. CMT refill phenomena are discussed further in Subsection 6.1.1.
The discussion of the short and long-term phase of the transient focuses on important thermal-hydraulic phenomena identified in the PIRT (Table 1.3-1). Key indicators of the quality of the analysis on which this discussion is based are the mass and energy balance results. These are discussed in detail in Subsections 6.2.2 and 6.2.3. /~N O maap600.23h53.noa:lb.ioo395 5.3-1 REVISION: 1
l i l tO V 5.3.1 Facility Performance i ne performance of the OSU test facility during Matrix Test SB06 in reference to the five transient phases is outlined in the following: l l
- Blowdown
. Natural circulation
. ADS
. IRWST injection I
. Sump injection I 1
l De overall performance of the facility during the transient is shown in Figures 5.3.1-1 to 5.3.1-4. l Figure 5.3.1-1 shows the pressurizer pressure throughout the test with various phases and operating I components. He time scale was reduced for clarity since there were only small changes in system pressure during the long-term phase of the transient. Figure 5.3.1-2 shows the total DVI line flow and its composition from the various sources at each time in the transient. Figure 5.3.1-3 shows the calculated core steam generation rate throughout the test and Figure 5.3.1-4 the variation in average measured core outlet temperature and peak clad temperature relative to the core outlet saturation j temperature. l
;]V Figures 5.3.1-1 and 5.3.1-2 show that throughout the transient there was a continuous flow of cooling water to the core from the passive safety-related systems. Once initiated, the ADS rapidly I depressurized the primary system and thus enhanced CMT and accumulator injection flow rates.
Ultimately, the ADS-4 valves reduced the system pressure sufficiently to allow gravity-driven IRWST l injection to begin. Operation of the passive injection systems overlapped so that as one source of water drained the next was available to continue the cooling process. He level of steam generation in the core and the response of the average measured core outlet fluid temperatures and maximum clad j temperatures is shown in Figures 5.3.13 and 5.3.1-4. Rese figures show that the cooling flow l prevented excessive core heating, and the core remained covered. he core remained subcooled for large periods of the transient and when steam production occurred, the rate of generation remained well below the rate at which water is delivered to the core. l 5.3.1.1 Blowdown Phase ne blowdown phase began at time zero when the break was initiated and continued until the primary system pressure was in equilibrium with the secondary-side pressure at about [ l'A' seconds. During this phase of the transient, cooling flow was provided from the two CMTs, which remained in the recirculation mode and heat was removed from the primary system via the SG. The pressurizer and surge line completely drained at [ ]'A' and [ ]*** seconds, respectively.
\
maapsoo 344w-53.nonmoo395 5.3.1-1 REVISION: 1
5.3.1.2 Natural Circulation Phase in this LOCA simulation, after a brief period of stability, the single and two-phase natural circulation phase was marked by a gradual reduction in system pressure rather than by the more stable pressure observed in SB01. During this phase of the transient, the SG tubes drained by about [ ]'** seconds and at this time, heat removal from the primary system continued via the PRHR. In response to voiding in CL-3, CMT-1 transitioned to draindown mode at [ ]'6# seconds, after the end of the blowdown phase, and the falling CMT level reached the ADS low-level setpoint at [ ]'6# seconds. At [ ]'6# seconds, both accumulators began to inject at [ ]'6# lbm/sec. without affecting the CMT outflow. De rate of accumulator injection gradually rose to [ ]'** lbm/sec. over the next [ ]'6# seconds. He natural circulation phase of the transient continued to [ ]'*# seconds when the ADS-1 valve opened. He minimum RPV inventory of [ ]'** lbm was observed at about [ ]'** seconds. 5.3.1.3 Automatic Depressurization System Phase ADS-i actuation was followed by ADS-2 and ADS-3 [ ]'** and [ ]'** seconds later. Coinciding with the initiation of the ADS, a more rapid phase of accumulator injection began. De influx of cold water combined with increased venting via the ADS led to a rapid depressurization of the primary system. Actuation of ADS-4 at [ ]'6# seconds completed depressurization to a level which allowed IRWST injection at [ ]'6* seconds via DVI-l and [ ]'6" seconds via DVI-2. During the more rapid phase of accumulator injection, increased flow path resistance reduced flow out of the CMTs. As the accumulators drained, CMT flow resumed. De accumulators were fully drained [ ]'6# seconds before IRWST injection began. CMT flow became more sporadic [ ]'6# seconds after the start of IRWST injection, but the CMTs never fully drained during the short-term transient. Actuation of ADS-2 rapidly refilled the pressurizer as water and steam flowed out of the ADS. The pressurizer gradually drained by [ j'*# seconds. 5.3.1.4 In-Containment Refueling Water Storage Tank Injection IRWST injection signaled the transition from the short- to long-term phase of the transient. He initial phase of IRWST injection involved an increase in flow through the two DVI lines, which was followed by a gradual reduction as the driving head between the IRWST tank and the RCS fell due to the reduced IRWST water level. Once maximum flow was established, the influx of water from the IRWST was enough to keep the core subcooled until [ ]'6# seconds. Steam was subsequently generated in the core for the remainder of the transient. The IRWST injection phase between [ ]'A' seconds, was marked by oscillations in pressure and level throughout the primary system. These oscillations were also observed in the ADS-4 liquid flow rates. O maap60as2%-53.non:1b-too395 5.3.1-2 REVISION: 1
.~ . . - ~ . . . . . . - - . - _ . . . - . . . . - ~ - - . ~ _ - . - - . . - . . . . . . . . - . . . ~ -. . - . ~ .
1 4 ?, t 4 i 4 . ' 5.3.1.5 Sump injection i 5 i Flow from the primary sump began at [ ]'6' seconds, via the check valves around the main sump : j' injection valves, when the relative sump and IRWST driving heads allowed this. 'Ihe test was terminated before the IRWST level fell to (- ]'6# in. and thus, the main sump injection valves were 4- , ! not actuated. ' a i i i t 1 i l
- i 4
.i !O d i 5 1
- 1
- 1 i i l' i i i i a maap60&23h53.noa: b-loo 395 5.3.1-3 REVISION: 1 i'
d TABLE 5.3.1 1 O OSU TEST ANALYSIS PLOT PACKAGE FOR SUBSECTION 5.3.1 Plot No. Component Variables Units Description 1 Pressurizer CI'T-604 psia System pressure and event history 2 Water WWTDVII+WWTDVI2, Ibm /sec. Total of CMT, accumulator, IRWST, ; injection WOUTACCl+WOUTACC2, and sump injection flows l WWTIRW11+WWTIRW12, ) WWTSMPIT l 1 , 3 Reactor RPVASOU2 lbm/sec. Steam generation in reactor vessel vessel l l l 4 Reactor T08RPV, HBiXRPV, TS AT 'F Reactor vessel outlet temperature, vessel maximum clad temperature and fuel exit saturation temperature i 9 O m:vtxA2344w-53.non:Ib-100395 5.3.1 -4 REVISION: 1
-_ _m__._ _ _ __..___. ._____ _ _._ ._. - _ . - _ . . _ . _ . _ _ _ . _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _
M I l-t i I-e i i i , i i l l l t THE FIGURES LISTED IN TABLE 5.3.1 1 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT l l~ j 9 l l l i m:W344w.53.noo:lt>100395 5.3.1-5 REVISION: 1
4 4 5.3.2 Short-Term Transient i For the 4-in. cold-leg break LOCA simulation, Matrix Test SB06, the short-term transient encompassed the time frame up to [ ]'6d seconds. As can be seen from Figures .5.3.1-1, this period included the full depressurization of the facility through all four stages of the ADS together , with CMT and accumulator injection plus the initial stages of IRWSTinjection. The variation in mass, energy, pressure, and temperature throughout this stage of the tmnsient are illustrated in the plot ; package outlined in Table 5.3.2-1. The plots concentrate on the primary system including the accumulators, CMTs, IRWST, primary sump and flows from the primary system via the ADS, break and IRWST overflow. There were two principal parameters of interest for the short-term transient:
. Adequate flow from the passive systems to the reactor vessel must be maintained.
Adequate flow into the core must be maintained to ensure that decay heat was removed from the simulated fuel rods, without a temperature excursion. These parameters are addressed in the following discus.sion. 5.3.2.1 Maintenance of Core Cooling Mass Injected to the Primary System Figures 5.3.2-6 and 5.3.2-7 show the combined effect of the injection flows for the short-term phase of the transient. Separate plots of the individual contributions to the total flow can be located by consulting the plot package index given in Table 5.3.2-1. Figures 5.3.2-5 and 5.3.2-6 show how the CMTs, accumulators and IRWST combined to supply a continuous flow of cooling water to the core. During the first [ ]'6# seconds, cooling flow was provided by the CMTs and then was supplemented by gradually increasing flow from both accumulators. The flow from the CMTs began at an initial value of [ ]'6# Ibm /sec., and increased to over [ ]'6' lbm/sec. when the CMTs transitioned to draindown at [ ]'6" seconds. Following the start of accumulator injection, the CMT flow gradually reduced as the driving head fell in response to the CMT water heat-up and draindown until [ ]'6* seconds, when depressurization following ADS-1 initiation generated more rapid accumulator injection. The rapid accumulator flow temporarily stopped CMT flow, but led to an overall increase in flow to the core to a peak value of [ ]'6# lbm/sec. Following the end of accumulator injection, the CMTs again provided cooling flow until the flow from the IRWST caused CMT draindown to end. Since IRWST injection began before the CMTs had fully drained, there was no period of the short-term transient when the passive safety-related systems failed to provide flow to the RPV. p'y/ maap60ausm-53..oo: b.ioo395 5.3.21 REVISION: 1
Reactor Pressure Vessel and Downcomer Behavior The effect of water flow on the average measured core inlet / outlet temperatures and peak heater rod temperatures during the shon-term phase of the transient can be seen in Figures 5.3.2-3 and 5.3.2-57. He core outlet temperature reached the saturation point at approximately [ ]'6" seconds. Except for the period of rapid accumulator injection, the core exit temperature then remained at the saturation level for about [ ]'6# seconds, when the influx of water from the IRWST became sufficient to again subcool the core. He core then remained subcooled until approximately [ ]'6# seconds. Figure 5.3.2-57 shows that there were no significant excursions in heated rod temperatures throughout the short-term transient; therefore, sufficient core inventory and flow was maintained through this phase of the transient to remove the simulated decay heat generation. For significant ponions of the transient, a two-phase mixture was present in the core and upper plenum regions, with core boiling kept at a low level. The following discussion tracks the variation in water level and mass throughout ! the RPV and downcomer. De mass and level for the core region are shown in Figures 5.3.244 and 5.3.245. The collapsed liquid level in the core indicates that the heater rods remained covered with a single or two-phase mixture throughout the short-term transient. The minimum core inventory of [ ]*b'lbm occurred at [ ]'6# seconds into the transient, before the initial accumulator injection was established. As shown in Figure 5.3.245, during this phase of the transient, the collapsed liquid level dropped to [ ]" in. below the top of the heated rod length. The average void fraction of the core two-phase mixture may be estimated by dividing the measured core collapsed liquid level by the [ ]'6* in heated rod length. In this test, the minimum collapsed liquid level corresponded to a core void fraction of [ ].'6" l The collapsed liquid level in the upper plenum region covered by LDP-113, and the associated fluid mass are shown in Figures 5.3.249 and 5.3.248. It can be seen that during the period before accumulator i'1jection, the upper plenum within the span of LDP-ll3 fully drained. The start of rapid accumulator injection caused a refill to the elevation of the cold legs. Following the end of accumulator injection, the region of the upper plenum spanned by the LDP cell fully drained by [ ] seconds and remained so until IRWST injection supplied sufficient inventory to initiate a refill. Figures 5.3.2-50 and 5.3.2-51 show that the upper head drained very rapidly during SB06, but maintained a collapsed liquid level until after the upper plenum drained. The upper head refilled slightly during accumulator injection, and again when IRWST injection was fully established, but otherwise remained drained during the shon-term transient. The mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.3.241 and 5.3.242. The downcomer collapsed liquid level fell to the bottom of the cold-leg piping during the first [ ]'6# seconds, where it remained until IRWST injection was fully established. At the time of rapid accumulator injection, there was evidence of splashing between the core and downcomer. maap600s2ws3.non:1b-too395 5.3.2-2 REVISION: 1
O () IRWST injection maintained the collapsed liquid level at the center of the cold legs. As in SB21, the downcomer level showed irregular oscillations. For SB06, these began at [ J'*' seconds,and i continued until [ ]'** seconds. As in SB21, these may be indications of continued flow through l the SG tubes. There is also evidence of splashing between the core and downcomer at this time, as well as between [ ]'** seconds. l 5.3.2.2 Energy Transport from the Primary System i Following the break, energy was deposited in the primary system fluid by the heater rods to simulate l decay heat and the primary system metal as it cools down. Some fluid energy was lost to ambient and out of the break. Energy must be removed from the primary system to prevent excessive fluid and heater rod temperature excursions. In the AP600 plant, heat removal is designed to be achieved by a combination of the SGs and the PRIIR plus the ADS. Steam Generator and Passive Residual lleat Removal lleat Transfer During normal operation, most of the primary system heat was removed via the SGs; however, once the RCPs tripped, the reduced system flow caused a reduction in primary- to secondary-side heat transfer. The SGs were only available as heat sinks until the primary system pressure dropped to that
]
of the secondary side, then the two sides were in thermal equilibrium. The PRilR is designed to i remove heat from the primary system once the safety signal opens the isolation valve. The PRIIR will V continue to remove energy after the SGs are thermally isolated until ADS actuates. Once the ADS is actuated, it becomes the predominant path for the removal of energy from the primary system. Figure 5.3.2-33 shows the SG primary- and secondary-side pressure together with the PRIIR integrated heat transfer as represented by the IRWST fluid energy after allowing for the contribution from ADS 1-3 inflow. The SGs were potential sinks for primary system heat while the primary-side pressure was J above that of the secondary side, that is before [ ]*** seconds. PRIIR heat removal began l [ ]'*' seconds into the test and the PRIIR was responsible for all the IRWST heat-up until ADS-1 activation. Following the actuation cf the ADS, PRIIR heat transfer continued for a further l [ j'** seconds. This was a distinct departure from the behavior observed in the smaller break cases. During the active phase, the PRIIR transferred heat to the IRWST at an average rate of [ ]'** Btu /sec. Energy Transport via the Break and Automatic Depressurization System The mass flow rate from the primary system via the break is shown in Figures 5.3.2-67 and 5.3.2-68. During the first [ ]'*' seconds following the break, nearly [ J'A'lbm of water appeared to flow out of the primary system via the break (Figure 5.3.2-62). As shown in Figure 5.3.2 71, there was no evidence for such a loss of primary system inventory from the mass balance. The apparent mass flow n resulted from the indicated increase in break separator inventory caused by a rise in pressure above the V) ( ma.wxnww-53.noo:n,ioo395 5.3.2-3 REVIslON: 1
l fluid in response to the opening of the break. The first [ ]*6* seconds of Figures 5.3.2-67 and 5.3.2-68 should therefore be ignored as far as mass loss from the break is concerned. Following initiation of ADS 1-3, flow through the break continued and was augmented by steam and liquid flow through the ADS 1-3 valves. Between [ ]'6# and [ ]'6# seconds, ADS 1-3 caused the l system to depressurize rapidly and at [ ]'6" seconds, ADS-4 was initiated and the primary system continued to depressurize to containment pressure. . 1 I Beyond [ ]'6" seconds, there was continued flow through the break as cold-leg refill was occurring. ! Flow through the ADS continued at a declining rate until [ ]'6" seconds when the flow through the ADS stages 1-3 terminated and was replaced by flow through the lower resistance ADS-4 paths for the rest of the short-term transient. I Integrated mass flow from the primary system via the ADS and the break is shown in Figure 5.3.2-62 and the corresponding integrated energy flow is shown in Figure 5.3.2-69. The total system inventory plot given in Figure 5.3.2-70 indicates that during the short-term transient up to [ ]'6# Ibm of inventory was lost as steam through the break. This is consistent with the quantity of steam measured 1 flowing through the break (Figure 5.3.2-63). O I i 1 1 I l l 9 l maap60&2344 53.noo:Ib.ioo395 5.3.2-4 REVIs!ON: 1 I
~ . - - . - .- - _ . - - . - . . --. . - . . - - . -- .
l I i TABLE S.3.21 l OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SURSECTION 5.3.2 ) Plot No. Component Variables Units Description i Pressurizer CIT-foi psia System pressure 2 RPV RPVPWR kW Core power ! 3 RV TOIRPV, T08RPV, 'F Core inlet / outlet temperature, l ST08RPV saturation temperature 4 SG CIT-201, CIT-204, psia Pnmary and secondary pressures in SG CPT-301, CIT-302 i 5 DVI-1 WWTDVIL1, Ibm /sec. Individual components and total flow in WWTIRWII, DVI l WOUTACCl, WWTIRW13 6 D VI.2 WWTDVIL2, Ibm /sec. Individual components and total flow in WWURWI2, DVI-2 WOUTACC2, WWTIRWI4 7 CMT AMCMTIB, Ibm Fluid mass in CMTs (excludes balance AMCMT2B lines) 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT MIWDVIL1, Ibm Integrated mass out of CMTs ! MIwDVIL2 10 CMT WW'IDVIL1, Ibm /sec. Flow out of CMTs WWIDVIL2 11 CMT WOUTCLBI, Ibm /sec. Flow into CMTs WOUICLB2 12 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 13 CMT UCMT1, UCMT2 Btu Fluid energy in CMTs 14 IRWST IRWST Ibm Mass of fluid in IRWST 15 IRWST CLDP-701 in. Collapsed liquid level in IRWST 16 IRWST WWTIRWII, Ibm /sec. Flow from IRWST to DVI lines WWTIRWI2 17 IRWST IRWSTOR lbm/sec. Overflow from IRWST to sump 18 IRWST ADS 13TMR lbm/sec. Total ADS flow into IRWST 19 IRWST ADS 13TIR, MIIRWII, Ibm Integrated mass out of IRWST MIIRW12, MIIRWIO 20 IRWST UIRWST Btu Fluid energy in IRWST 21 PRHR CLDP-802 in. Collansed liquid level in PRHR HX O ma.p6co23m-53.noosioo395 5.3.2-5 REVISION: 1
1
\
TABLE 5.3.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.3.2 Plot No. Component Variables Units Description 22 PRHR WWOTPRHR lbm/sec. Measured outlet flow from PRHR tube 23 Accumulator AMACCl, AMACC2 lbm Mass of fluid in accumulators 24 Accumulator CLDP-401, CLDP-402 in. Collapsed liquid level in accumulators 25 Accumulator WOUTACCl, Ibm /sec. Flow from accumulators WOUTACC2 26 Accurnulator MOUTACCI, Ibm Integrated mass out of accumulators MOUTACC2 l 27 Accumulator UACCI, UACC2 Btu Fluid energy in accumulators 28 Pnmary sump AMPSMP lbm Prunary sump fluid mass 29 Pnmary sump CLDP-901 in. Pnmary sump level 30 Pnmary sump UPSMP Btu Prunary sump fluid energy 31 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG pnmary side MSSGOPl. MSSGOP2 inlet / outlet plena 32 SG MSSGHTI, MSSGHT2. Ibm Mass of fluid in SG primary side hot MSSGCT1, MSSGCT2 and cold tubes 1 33 SG/PRHR CFT-201, CPT-301, psia & SGI pressure and PRHR integrated heat l QPRHRI Btu output 34 Pressurizer PZM lbm Fluid mass in pressurizer 35 Pressurizer CLDP-601 in. Collapsed liquid level in gressurizer 36 Pressurizer UPZ Blu Fluid energy in pressurizer 37 Surge line PLM lom Fluid mass in surge line 38 Surge line CIDP-602 in. Collapsed liquid level in surge line 39 Surge line UPSL Btu Fluid energy in surge line 40 RPV MWRPV lbm Total fluid mass in reactor vessel 41 RPV DCM lbm Fluid mass in downcomer 42 RPV LDP01DC in. Collapsed liquid level in dowmcomer compared to various reference elevations 43 RPV MWOIRPV lbm Fluid mass in lower pienum 44 RPV MWO3RPV lbm Fluid mass in core region 45 RPV LDP03RPV in. Collapsed liquid level in core 46 RPV RPVAVDF2 Core exit void fraction 47 RPV RPVAQOU2 Core exit quality 48 RPV MWO6RPV lbm Fluid mass in the upper plenum 49 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 50 RPV MWO8RPV lbm Fluid mass in the upper head 51 RPV LDP08RPV in. Collapsed liquid level in the upper head mnap60m:344.-53.noo: b.ioo395 5.3.2-6 REVISION: 1
U TABLE 5.3.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.3.2 Plot No. Component Variables Units Description 52 RPV URPV Btu Total fluid energy in reactor vessel 53 RPV RPVXE, RPVASL2 in. Level of Tsat line 54 RPV RPVPab, RPVAPab2, kW Heated rod power above and below RPVPWR Tsat level and total 55 RPV RPYRXV,RPVASOU2 lbm/sec. Core steam generation rate 56 RPV RPVALIN2 lbm/sec. Calculated core flow 57 RPV HTMXRPV, ST08RPV 'F Maximum clad temperature and saturation temperature 58 Hot leg MWHL1, MWHL2 lbm Water mass in bot legs 59 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs 60 Cold leg CLlWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 61 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 62 ADS and break BRK311x, ADS 12TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4s, and break 63 ADS and break BRKTIVF, ADl3TIVF, Ibm Total integrated vapor flow for ADS AD41TIVF, AIM 2TIVF O> 64 ADS and break BRKTILF, AD13TILF, Ibm and break Total integrated liquid flow for ADS AD41TILF, AIM 2TILF and break 65 ADS and break ADS 13SVR, Ibm /sec. Vapor flow out ADS 1-3 and ADS-4 ADS 41SVR, ADS 42SVR 66 ADS and break ADS 13SLR, Ibm /sec. Liquid flow out ADS 1-3 and ADS-4 ADS 41SLR, ADS 42SLR 67 ADS and break BRKSSVR Ibm /sec. Vapor flow out of break 68 ADS and break BRKSSLR lbm/sec. Liquid flow out of break 69 ADS and break BRKSPEI, ADS 13EI, Btu Integrated fluid energy for ADS 1-3, ADS 41EI, ADS 42EI ADS-1, and break 70 Mass balance TOTMASS lbm Total system mass inventory 71 Mass balance PRIMMASS, Ibm Measured prunary system inventory and PRIMASS2 value from mass balance 72 Mass balance MERROR Ibm Mass balance error 73 Mass balance MIN, MOUT lbm Integrated mass flow in and out of SRCMASS primary system and source mass 74 Energy balance Various Btu Components of energy balance v m:W3h53.noo:sts too395 5.3.2-7 REVISION: 1
- . =
. . ~... ~ . . . . - . . . . . , . . - . . ~ _ . - .- ... _ .. .. . .- .. . . . - - . - - . . . - . - - - - - - ~ ~ . . . . . . - . , ~ . . . . - . . .
) i f O I TIIE FIGURES LISTED IN TABLE 53.2-1 ARE NOT INCLUDED IN TIIIS NONPROPRIETARY DOCUMENT O O "hA344w-53.non:Ib-100395 5.3.2-8 REVISION: 1
1 5.3.3 Long Term Transient The long-term transient started with initiation of IRWST injection, covers the transition from IRWST to sump injection, and provides information on the long-term cooling response of the AP600 plant. For the large cold-leg break, Matrix Test SB06, the long-term transient analyzed runs from [ ]*" seconds to the end of the test at near [ l'* seconds. The behavior of the test facility during this period of the transient is discussed in this subsection using the plot package detailed in Table 5.3.3-1. This analysis concentrates on the components of the primary system that remained active during the long-term cooling phase, that is, the RPV, the hot legs, ADS-4, the sumps, and the IRWST. l During the long-term transient, the main thermal-hydraulic phenomena of interest were:
. Maintenance of core cooliti and removal of energy from the primary system.
l Level oscillations (from [ ]** seconds, there were system-wide level and pressure oscillations, which are discussed further in Subsection 6.1.3). 5.3.3.1 Maintenance of Core Cooling O (/ Mass Injected into Primary System , j I Total DVI line flow, CMT flow and IRWST flows are shown in Figures 5.3.3-6 and 5.3.3-7 and the flow from the primary sump is shown in Figure 5.3.3-19. From around [ ]'* seconds, there was a contribution to the DVI flow from the CMTs as the CMTs reached post-refill draindown. ) i During the IRWST injection phase of the transient, IRWST flow proceeded at a gradually declining rate with the effect of the primary system oscillathns superimposed. At [ j'" seconds, flow from the pumary sump began through check valves around the main injection valves. This caused the IRWST flow to cease. The test was not continued to the point where the IRWST level fell to the low-low setpoint and tims, the main sump injection valves were not actuated. Reactor Pressure Vessel and Downcomer Response The effect of the water inflow on the average measured downcomer fluid temperatures, core inlet and core outlet temperatures, and peak clad temperatures during the long-term phase of the transient is shown in Figures 5.3.3-4,5.3.3-5, and 5.3.3-38. Figure 5.3.3-4 shows that there is a general increase in average downcomer fluid temperatures during the long-term transient. By the end of the test, this average temperature reached an equilibrium [ ]'* 'F below saturation. Figure 5.3.3-5 shows that the core average outlet fluid temperature remained at saturation during the long-term transient. p Figures 5.3.3-34 to 5.3.3-36 show that the DVI line flow method discussed in Section 4.11 indicates () that a small level of boiling was maintained throughout the long-term transient. Nevertheless, the level mhp6cm3uw 53.oon:itsioo395 5.3.3- 1 REVIs!ON: 1
l 1 of boiling was small and the test results showed that the inflow from the IRWST was sufficient to maintain cooling. Figure 5.3.3-38 shows that throughout the long-term transient there were no significant excursions in heater rod temperatures; therefore, sufficient core inventory and flow was maintained through this phase of the transient to remove the decay heat generated. For significant portions of the trarsient, a l two-phase mixture was present in the co- * . cpper plenum regions. He following discussion tracks the variation in water level and mass ; , vot the RPV and downcomer. Mass and level for the core region are shown in Figures 5.3.3-28 and 5.3.3 29. The collapsed liquid level in the core indicated that the heater rods were always covered with a single- or two-phase mixture. During the initial core oscillations, the maximum core void fraction was [ ],"' and for much of the transient the average core void fraction was [ ] ."# During the long-term transient represented, the calculated steam generation rate maintained an average value of around [ ]** lbm/sec. (Figure 5.3.3-36). The collapsed liquid level in the upper plenum region covered by LDP-ll3, is shown in Figure 5.3.3-32. For the entire long-term transient, the average collapsed liquid level was at, or above, the hot-leg elevation. The effect of the oscillations during the long term transient were observed. De mass of water in the RPV is shown in Figure 5.3.3-25. For the entire long-term transient, the RPV water mass remained at an average value of ( )"# lbm, which is [ ]"# percent of the initial vessel water inventory. From ! ]'6# seconds, oscillations in vessel inventory were observed. Figures 5.3.3-51 to 5.3.3-56 illustrate these oscillations using plots on a restricted time frame from [ ]'6' seconds. These oscillations are observed in primary system measurements from the upper plenum to the ADS-4 flows. The oscillations occurred with a period of about [ ]"' seconds. The oscillations in the ADS flow lagged behind those in the upper head pressure. These oscillations and possible mechanisms for their production are discussed further in Subsection 6.1.3. The mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.3.3-26 and 5.3.3-27. The collapsed liquid level remained above the cold legs until the end of the test. Until [ ]'6" seconds, there were irregular oscillations which may be related to continued steam flow and core to downcomer splashing. 5.3.3.2 Energy Transport from the Primary System During the long-term transient, energy continued to be deposited in the primary system from the heater rods, metal and fluid flowing from the primary sump. The SGs and PRHR remained inactive throughout this phase of the transient, and the primary path for energy out of the primary system was via the ADS-4 valves. maap600cm-51ao :ib.ioo395 5.3.3-2 REVislON: 1
I Integrated mass flow from the primary system via the ADS and the break is shown in Figua 5.3.3-43. This figure shows that during the LTC phase of the transient, the only significant energy outflows are through the ADS-4 valves and, until [ ]'6* seconds, the break. This is confirmed by Figures 5.3.3-44 to 5.3.3-45, which show flow through the ADS and break. After [ ]'6* seconds, there was reverse flow through the break as is indicated by the reducing integrated flow shown in Figure 5.3.3-43. Figure 5.3.3-36 shows the calculated steam generation rate as determined by the DVI line flow method During the long-term transient, steam was being generated at an average rate of [ ]'6* lbm/sec., although the steam vortex meters indicate little or no flow out of the ADS-4 valves. However, examination of the fluid thermocouples on the outlet of both the ADS-4 valves indicated that they reached saturation temperature during much of the transient after [ J'6* seconds. Furthermore, it can be concluded from the discussion in Subsection 6.1.3 that it was not possible for all the steam generated in the core to flow from the upper head to the downcomer via the bypass holes. It can therefore be concluded that steam was leaving the primary system via the ADS-4 valves although the steam vortex meters did not measure any steam flow. Figure 5.3.3-50 shows all the components contributing to the system energy balance. Further discussion of steam loss from the primary system is provided in the mass and energy balance discussions of Section 6.2. O mAap60m23uw-53.noa:16 ioo395 5.3.3-3 REVislON: 1 l
TABLE 5.3 31 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSELTION 5.3.3 LONG TERM TRANSIFRI' Plot No. Component Variables Units Description 1 RPV RPVPWR kW Core power 2 Pnmary sump TSMPII, TSMPI2 *F Sump injection line temperatures 3 DVI TDVILI, TDVIL2 'F DVI line temperatures 4 RPV T01DC, TI)2DC, T03DC, 'F Water and saturation temperatures in S'11)lDC downcomer 5 RPV TOIRPV, TI)8RPV, 'F Core inlet / outlet temperature, STT)8RPV saturation temperature 6 DVl-1 WWTDVIL1, Ibm /sec. Individual components and total flow WWTIRWII, in DVI-l WWTIRW13 7 DVI-2 WWIDVIL2, Ibm /sec. Individual components and total flow WWTIRW12, in DVI-2 WWTIRW14 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT CLDP-509, CLDPSIO in. Level CL-CMT balance lines 10 IRWST IRWST lbm Mass of fluid in IRWST I1 IRWST CLDP-701 in. Collapsed liquid level in IRWST 12 IRWST UIRWST Btu Fluid energy in IRWST 13 Pnmary sump AMPSMP lbm Pnmary sump fluid mass 14 Pnmary sump CLDP-901 in. Pnmary sump level 15 Prunary sump UPSMP Btu Pnmary sump fluid energy 16 Secondary sump AMSSMP lbm Secondary sump fluid mass 17 Secondary sump CLDP-902 in. Secondary sump level 18 Secondary sump USSMP Btu Secondary sump fluid energy 19 Pnmary sump WSTSMPET, WWTSMPIT lbm/sec. Pnmary sump steam and liquid injection rate 20 Pnmary sump MISMPII, MISMPI2, Ibm lategrated primary sump and IRWST MISMPIT, MIIRWT flows 21 SG MsSGIP1, MSSGIP2, Ibm Mass of fluid in SG side inlet / outlet MSSGOP1, MSSGOP2 plena 22 Surge line PLM lbm Fluid mass in surge line 23 Surge line CLDP-602 in. Collapsed liquid level in surge line 24 Surge line UPSL Btu Fluid energy in surge line 25 RPV MWRPV lbm Total fluid mass in reactor vessel O mAap60m2m53.noa:Ib.100395 5.3.3-4 REVISION: 1
TABLE 5.3.31 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.3.3 LONG TERM TRANSIENT Plot No. Component Variables Units Description 26 RPV DCM lbm Fluid mass in downcomer 27 RPV LDPOIDC in. Collapsed liquid level in downcomer compared to various irference elevations 28 RPV MWO3RPV lbm Fluid mass in core region 29 RPV LDP03RPV in. Collapsed liquid level in core 30 RPV RPVAVDF2 Core exit void fraction 31 RPV RPVAQOU2 Core exit quality 32 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 33 RPV MWO8RPV lbm Fluid mass in the upper head . 34 RPV RPVASL2 in. Level of Tsat line 35 RPV RPVAPab2, RPVPWR kW Heated rod power above and below Tsat level and total 36 RPV RPVASOU2 lbm/sec. Core steam generation rate 37 RPV RPVALIN2 lbm/sec. Calculated core flow 38 RPV HTMXRPV, 'F Maximum clad temperature, saturation ST08RPV temperature and delta 39 Hot leg MWHL1, MWHL2 lbm Water mass in bot legs 40 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs 41 Cold leg CLIWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 42 Cold leg _ CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 43 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 13, ADS 41TIR, ADS 42TIR ADS 4, and break 44 ADS and break - ADS 13TLR, ADS 41TLR, Ibm /sec. Liquid flow out ADS 1-3 and ADS.4 ADS 42TLR 45 ADS and break BRKSTLR lbm/sec. Liquid flow and total flow out of break 46 Mass balance TOTMASS lbm Total system mass inventory 47 Mass balance PRIMMASS, PRIMASS2 lbm Measured primary system inventory and valve from mass balances 48 Mass balance MERROR lbm Mass balance error 49 Mass balance MIN, MOUT SRCMASS lbm Integrated mass flow in and out of primary system and source mass 50 Energy balance Various Btu Component of energy balance 51 ADS 4 ADS 41TLR, ADS 42TLR lbm/sec. Oscillations in ADS-4 liquid flow 52 Surge line CLDP-602 in. Oscillations in surge line level mAarue23m-53.noo:Ib-too395 5.3.3 5 REVISION: 1
TABLE 5.3.31 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.3.3 LONG TERM TRANSIENT Plot No. Component Variables Units Description 53 RPV CI"T107 psia Oscillations in upper bead pressure 54 RPV CLDP-il3 in. Oscillations in upper plenum level 55 RPV LDP03RPV in. Oscillations in core level 56 RPV LDP01DC in. Oscillations in downcomer level O O m:'ep6002344w-53.non:lb-100%s 5.3.3-6 REVISION: 1
t l O _ THE FIGURES LISTED IN TABLE 533-1 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT O O m:\ap60m2344w-53.non:1b-100995 5.3.3 7 REVISION: 1
. _ - . _ _ _ _ _ . _ . - _ . . _ _ - . . . . _ . _ . _ _ _ _ . . . ~ _ - _ _ _ _ _ . _ _ . _ ~ _ . _ _ . _ . _ _ . _ . _ _ . _ . _ _
i i a f 5.4 Analysis of Matrix Test SB09 Matrix Test SB09 (OSU Test U0009) simulated a 2-in. LOCA in the CL-3 to CMT-1 balance line with LTC and without the operation of the nonsafety-related systems. The break was located on the horizontal section of the balance line before the vertical rise to CMT-1. Except for the break location, this test was similar to SB01, including the simulated failure of one of the ADS-4 lines. Changes to ; the OSU facility since the performance of SB01 are noted in the Final Data Report.m ; The analysis of Matrix Test SB09 is divided into three subsections as follows:
- Facility performance is discussed in Subsection 5.4.L lt provides a brief outline of the .
response of the test facility; further details are available in the Final Data Report.*
- The short-term transient for SB09 encompassed the start of the simulation up to [ J'6#
seconds. This period included the blowdown, natural circulation, ADS, and initial IRWST stages of the transient.
. The analysis of the long-term transient for SB09 encompassed the time frame from [ J'6" seconds to the end of the test. This phase of the transient includes the IRWST injection and covers the transition to sump injection. The long-term transient actually started at IRWST injection, which is discussed as part of the short-term transient. Between the end of the short-term transient and [ J'6# seconds, the system remained relatively inactive with the exception of the CMT-2 refill. At [ J'6# seconds, CMT-2 began to refill; CMT-1 [
]'6# during the SB09 transient. CMT refill phenomena are discussed further in Subsection 6.1.1, and the discussion of the long-term transient provided here begins at [ ]'6# seconds.
The discussion of the short- and long-term phase of the transient focuses on important thermal-hydraulic phenomena identified in the PIRT (Table 1.3-1). Key indicators of the quality of the analysis on which this discussion is based are the mass and energy balance results. These are discussed in detail in Subsections 6.2.2 and 6.2.3. maap600c344w.54.non:Ib 100395 5.4-1 REVISION: 1
. - - _ ~ . - . . _ . . . - . - - - . . - . - . - . - . - - . . - _ - - ...-.- .-.
5.4.1 Facility Performance ' De performance of the OSU test facility during Matrix Test SB09 in reference to the five transient phases is outlined in the following:
. Blowdown
,. Natural circulation
. ADS
. 1RWST injection
. Sump injection l The overall performance of the facility during the transient is shown in Figures 5.4.1-1 to 5.4.1-4.
! Figure 5.4.1-1 shows the pressurizer pressure throughout the test with various phases and operating , 1 components delineated on the figure. De time scale was reduced for clarity since there were only , l small changes in system pressure during the long-term phase of the transient. Figure 5.4.1-2 shows
. the total DVI line flow and its composition from the various sources at each time in the transient.
Figure 5.4.1-3 shows the calculated core steam generation rate throughout the test. Figure 5.4.1-4 shows the variation in average measured core outlet temperature and peak clad temperature relative to the core outlet saturation temperature. l Figures 5.4.1-1 and 5.4.1-2 show that there was a continuous flow of water to the core from the
\
passive safety-related systems throughout the transient. Once initiated, the ADS lines rapidly l depressurized the primary system, enhancing the CMT and accumulator injection flow rates. Ultimately, the ADS-4 valves reduced the system pressure sufficiently to start gravity-driven IRWST , ! injection. The passive injection systems operation overlapped so that as one source of water drained j the next was available to continue the cooling process. De level of steam generation in the core and l the response of the average measured core outlet fluid temperatures and maximum clad temperatures are shown in Figures 5.4.1-3 and 5.4.1-4. These figures show that the cooling flow prevented core heatup, and the core remained covered. The core remained subcooled for large periods of the transient. When steam was produced, the rate of generation remained well below the rate at which water was delivered to the core. 5.4.1.1 Blowdown Phase i l The blowdown phase began at time zero when the break was initiated and continued until the primary l circuit pressure is in equilibrium with the secondary-side pressure at around [ ]'*d seconds. During l this phase of the transient, cooling flow was provided from the intact CMT, while the CMT with a broken balance line injected very little mass. CMT-2 remained in the recirculation mo& "ntil almost the end of this phase, and heat was removed from the primary circuit via the SGs. Th: pressurizer j and surge line completely drained at [ ]'*# and [ ] seconds, respectively. t m:\ap600\2M4w-54.non:Ib-100395 5.4.1-1 REVISION: 1
5.4.1.2 Natural Circulation Phase - In this LOCA simulation, the single- and two-phase natural circulation phase initially continued the gradual reduction in system pressure characteristic of blowdown; later on in this phase of the transient, the rate of depressurization increased significantly once all the SG tubes had drained at about [ ]'6# seconds. The tubes in SG-2 in the nonbalance line loop completed draining almost [ ]'6# seconds later than those in SG-1. After [ ]'6# seconds, heat removal from the primary circuit continued via the PRHR and the break. In response to the enhanced depressurization rate, CMT-2 transitioned into a rapid draindown at [ ]'6# seconds, and the falling CMT level reached the ADS low-level setpoint so that the ADS-1 valve began to open at [ ]'6# seconds. By [ ]'6# seconds, both accumulators began to inject. 5.4.13 Automatic Depressurization System Phase ADS-1 actuation was followed by ADS-2 and ADS-3 [ ]'6# and [ ]'b' seconds later. An increased accumulator injection rate was observed with initiation of ADS-2. He influx of cold water combined with increased venting via the ADS led to an even more rapid depressurization of the primary system. Actuation of ADS-4 at [ ]'6# seconds completed the depressurization to the extent that the IRWST began injecting at [ ]'6# seconds via DVI-2 and [ ]'6# seconds via DVI-i. During the rapid accumulator injection, increased flow path resistance reduced flow out of CMT-2 and stopped flow out of CMT-1. CMT flow resumed at rates approaching [ ]'6#1bm/sec. as the accumulators drained. Because the CMT-1 balance line contained the break, CMT-1 injection was delayed and overlapped with IRWST-2 injection for more than [ ]'6# seconds. CMT-1 flow continued concurrently with in [ ]'6# seconds of IRWST-1 injection. He minimum RPV mass inventory of [ ]'6# lbm occurred just before IRWST injection began. Actuation of ADS-1 and ADS-2 rapidly refilled the pressurizer as water and steam flowed out of the ADS. De pressurizer gradually drained by [ ]'6# seconds. 5.4.1.4 In Containment Refueling Water Storage Tank Injectior, IRWST injection signals the transition from the short- to long term phase of the transient. Initially, IRWST injection was delivered solely through the IRWST-2 DVI line with flow gradually increasing as the driving head between the IRWST and the RCS increased. The pressure differential increased because RCS pressure decreased as the core steam generation decreased from [ ]'6# lbm/sec. at IRWST-2 initiation to zero at [ ]'6# seconds. De maximum IRWST flow was established shortly thereafter, then it gradually decreased with the decrease in pressure differential as the IRWST continued to drain. The influx of water from the IRWST was enough to keep the core subcooled until [ ]'6# seconds. Steam was subsequently generated in the core for the remainder of the transient. Following the restan of core steam generation, IRWST injection between [ ]' 6# seconds was marked by oscillations in pressure and level throughout the primary system. These oscillations were also observed in the ADS-4 liquid flow rates. mw60cc344+54.non:tb-ioo395 5.4.1-2 REVISION: 1 I
- _ _ _ - . _ _ . _ _ . _ _ . ._. . . _ _ _ _ _ . . _ . _ _ - . .__ _ _.. _ . _-... _ _._ _ . __ _ ____=_._._ D h 5.4.1.5 Sump Injection Sump flow began at [ ]** seconds through the check valves around the main sump injection valve when the driving head from the sump was sufficient for flow to initiate in DVI-2. Sump injection through DVI l began to enter the IRWST once the main sump injection valves opened; at that time DVI 2 exhibited a corresponding increase in flow out of the IRWST, which meant that there was no increase in IRWST inventory. Flow through the main sump injection valves began when those valves opened at about [ ]*' seconds. i l l i I l l 1 l O V m:up6cos2344 54.=:isioo395 5.4.1-3 REVISION: 1
)
TABLE 5.4.1 1 OSU TEST ANALYSIS PLOT PACKAGE FOR SUBSECTION 5.4.1 Plot No. Component Variables Units Description 1 Pressurizer CPT-604 psia System pressure and event history 2 Water WWTDVil+WWTDV12, Ibm /sec. Total of CMT, accumulator, IRWST, injection WOUTACCl+WOUTACC2, and sump injection flows WWTIRWII+WWTIRW12, WWTSMPIT 3 Reactor RPVASOU2 lbm/sec. Steam generation in reactor vessel vessel 4 Reactor T08RPV, HTMXRPV, TSAT *F Reactor vessel outlet temperature, vessel maximum clad temperature and fuel exit saturation temperature 6 I l 9 mspun2344w-54.noa:tb.noo395 5.4.1-4 REVISION: 1
O -
.THE FIGURES LISTED IN TABLE 5.4.11 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT O
O m:W344w-54.aon:1b-190395 3,4,1 5 REVISION: 1
i--.-._.-.-. i i 1 5.4.2 Short Term Transient - For the 2 in, balance line break LOCA simulatian, Matrix Test SB09, the short-term transient l encompassed the time frame up to 3000 seconds. As shown in Figure 5.4.1-1, this period included { full depressurization of the facility through all four stages of the ADS, together with CMT and accumulator injection plus the initial stages of IRWST injection. He variation in mass, energy, ! pressure, and temperature throughout this stage of the transient are illustrated in the plot package i outlined in Table 5.4.2-1. The plots concentrate on the primary system, including the accumulators, CMTs, IRWST, primary sump and flows from the primary system via the ADS, break, and IRWST ! overflow. There were two principal parameters to be examined for the short-term transient:
. Adequate flow from the passive systems to the reactor vessel must be maintained.
Adequate flow into the core must be maintained to ensure that decay heat was removed from the simulated fuel rods, without a temperature excursion. These parameters are addressed in the following discussion. 5.4.2.1 Maintenance of Core Cooling Mass Injected to the Primary System Figures 5.4.2-6 and 5.4.2-7 show the combined effect of the injection flows for the short-term phase of the transient. Separate plots of the individual contributions to the total flow can be located by consulting the plot package index given in Table 5.4.2-1. Figures 5.4.2-5 and 5.4.2-6 show how the CMTs, accumulators, and IRWST supplied a continuous flow of water to the core. During the first [ l'6" seconds, cooling flow was provided primarily by CMT-2, since CMT-1 flow was severely limited by the break in its balance line. By [ ]'6' seconds, CMT flow was supplemented by flow from both accumulators. The rate of flow from the CMTs was reduced or stopped by accumulator injection. Accumulator flow produced maximum DVI injection rates during the entire transient with values of [ ]'6' lbm/sec. and above in each DVI line. Following the end of accumulator injection, the CMTs again provided cooling flow. Because CMT-1 was still almost full at the end of accumulator injection, it provided greater flow than CMT 2 through the remainder of the CMT drain period. While CMT-2 emptied [ ]'6' seconds before the flow from IRWST-2 began, CMT-1 was still injecting at a rate of [ J'6" lbm/sec. so continuous injection was maintained. IRWST-1 injection began [ ]'6# seconds before the CMT-1 draindown completed. Since continuous IRWST injection through both DVI lines began before CMT-2 had fully drained, O there was no period of the short-term transient when the passive safety systems failed to provide flow to the RPV, 1 maap600(23h5toon:tb. loo 395 5.4.2 1 REVISION: 1
Reactor Pressure Vessel and Downcomer Behavior ne effect of water flow on the average measured core inlet / outlet temperatures and peak clad temperatures during the short-term phase of the transient is shown in Figures 5.4.2-3 and 5.4.2-57. The core outlet temperature first reached the saturation point at [ ]** seconds. The core outlet fluid temperature became subcooled again returning to the saturation level for the period between [
]** seconds, after which, the influx of water from the accumulators kept the core subcooled until
[ ]'6# secs. At [ ]'6# seconds, the influx of water from the IRWST was sufficient to subcool the core again. De core then remained subcooled until the end of the short-term transient. Figure 5.4.2-57 shows that there were no significant excursions in heated rod temperatures throughout the short-term transient, therefore, sufficient core inventory and flow were maintained through this phase of the transient to remove the decay heat generated. For significant portions of the transient, a two-phase mixture was present in the core and upper plenum regions, with core boiling kept at a low level. The following discussion tracks the variation in water level and mass throughout the reactor vessel and downcomer. The mass and level for the core region are shown in Figures 5.4.2-44 and 5.4.2-45. The collapsed liquid level in the core indicates that the heated rods remained covered with a single- or two-phase mixture throughout the short-term transient. The minimum core inventory of [ ]'6# lbm occurred at about [ ]** seconds into the transient before the initial accumulator injection was fully established. As shown in Figure 5.4.2-45, the collapsed liquid level dropped to [ ]*6* in. below the top of the heated rod length during this phase of the transient. The average void fraction of the core two-phase mixture may be estimated by dividing the measured core collapsed liquid level by the [ ]** in. heated rod length. In this test, the minimum collapsed liquid level corresponded to a core void fraction of [ ].** By the end of the short-term transient, the effect of IRWST injection ended core boiling (Figure 5.4.2-55), and the core was again water-solid. I The collapsed liquid level in the upper plenum region and the associated fluid mass are shown in Figures 5.4.2-49 and 5.4.2-48. Figures 5.4.2 50 and 5.4.2-51 show that the upper head had only ) partially drained while accumulator injection and ADS-1 actuation occurred; then it rapidly refilled by about [ ]*6# in. during Matrix Test SB09. The upper head resumed draining and refilled later in the short-term transient during maximum IRWST injection. l l The mass of fitdd and collapsed liquid level in the RPV downcomer are shown in Figures 5.4.2-41 and 5.4.2-42. The downcomer collapsed liquid level fell to the bottom of the cold-leg piping during the first[ ]** seconds. IRWST injection maintained the collapsed liquid level within the cold-leg pipe perimeter after [ ]** seconds in the transient. O maap6am2344w-54.non:ib-too395 5.4.2-2 REVISION: 1
.. - ~ - . ~ . - - - . . - - - - . - . . - . ~ . - - - - _ - . . - - - - - . - . - . . . - - -
d ) 5.4.2.2 Energy Transport from the Primary System ' L !. Following the break, energy was deposited in the primary circuit fluid by the heater rods to simulate { decay heat and by the primary circuit metal as it cooled down. Some fluid energy was lost to the i atmosphere and out of the break. Excess energy must be removed from the primary system to prevent , j excessive fluid and heater rod temperature excursions. The AP600 is designed to remove heat by a j combination of the SGs and the PRHR plus the ADS.
- f. Steam Generator and Passive Residual Heat Removal Heat Transfer i
! During normal operation, most of the primary system heat was removed via the SGs; however, once the RCPs tripped, the reduced system flow decreased primary-to-secondary side heat transfer. De SGs were only available as heat sinks until the primary system pressure dropped to that of the secondary side; afterward, the two sides were then in thermal equilibrium. The PRHR is designed to remove heat from the primary system, once the safety signal opens the isolation valve. He PRHR continued to remove energy after the SGs were thermally isolated until ADS actuated. Once ADS actuated, ADS 1-3 became the predominant path for the removal of energy from the primary circuit.- Figure 5.4.2-33 shows the SG pressure equalization together with the PRHR integrated heat transfer as represented by the IRWST fluid energy after allowing for the contribution from ADS 1-3 inflow. As shown, heat was transferred to the secondary side of the SGs for only the first [ j'6# seconds. 5 PRHR heat removal began [ ]'6# seconds into the test, and the PRHR was responsible for all the IRWST heat-up until ADS-1 activation, after which the PRHR heat transfer was significantly reduced. During the natural circulation phase, the PRHR transferred heat to the IRWST at an average rate of [ ]'6' Btu /sec. Energy Transport via the Break and Automatic Depressurization System 1 The mass flow rate from the primary system via the break is shown in Figures 5.4.2-67 and 5.4.2-68. j During the first [ ]'6" seconds, break flow rose to a maximum value of over [ ]'6" lbm/sec. of water, then diminished as system pressure fell. At [ ]'6# seconds, mass appeared to flow out of the j primary circuit via the break (Figure 5.4.2-62) at an increased rate. Figure 5.4.2-71 shows the same trend. Apparently, this mass flow increase resulted from a decrease in the enthalpy of the fluid exiting the break. Following initiation of ADS 13, flow through the treak dropped greatly as steam and liquid flow. through the ADS 13 valves became the primary mass release path. Between [ ]'6#and [ ]'6# seconds, ADS 1-3 caused the system to depressurize rapidly at [ ]'6# seconds, ADS-4 was initiated and the primary system continued its depressurization to the BAMS header pressure. Beyond [ ' ]'6# seconds, there war continued flow through the break as cold leg refill was occurring due to accumulator injection. Liquid flow through the break in the balance line continued at a mw60m2344w.54.noo:ib.ioo395 5.4.2-3 REVISION: 1
l l l declining rate until [ ]'6' seconds, the flow through the break was then minimal for the rest of the short-term transient. Integrated mass flow from the primary system via the ADS and the break is shown in Figure 5.4.2-62. The corresponding integrated energy flow is shown in Figure 5.4.2-69. 'Ihe total system inventory plot given in Figure 5.4.2-70 indicates that only [ ]'6* lbm ofinventory left the system during the short-term transient. Components of the energy balance are shown in Figure 5.4.2-74. O l 9 mAapfM2344w-54. con:lb-100395 5,4,24 REVISION: 1
TABLE 5.4.21 - OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.4.2 Plot No. Component Variables Units Description 1 Pressurizer CPT-6N psia System pressure 2 RPV RPVPWR kW Core power 3 RPV TOIRPV, TD8RPV, 'F Core inlet / outlet temperature, ST08RPV saturation temperature 4 SG CPT-201, CPT-2N, psia Pnmary and secondary pressures in SG Cfrr-301, CI'T-302 5 DVI l WWTDVIL1, Ibm /sec. Individual components and total flow in WWTIRWil, DVI l WOUTACCl, WWTIRWI3 6 DVI 2 WWIDVIL2, Ibm /sec. Individual components and total flow in WWTIRWI2, DVI-2 WOUTACC2, WWTIRW14 7 CMT AMCMTIB, Ibm Fluid mass in CMTs (excludes balance AMCMT2B lines) 8 CMT CLDP-502, CLDP-507 . in. Collapsed liquid level in CMTs 9 CMT MIWDVIL1, Ibm Integrated mass out of CMTs O 10 CMT MIWDVIL2 WWrDVIL1, WWTDVIL2 Ibm /sec. Flow out of CMTs 11 CMT WOUTCLBI, Ibm /sec. Flow into CMTs WOUTCLB2 12 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 13 CMT. UCMTI, UCMT2 Btu Fluid energy in Chfrs 14 IRWST IRWST Ibm Mass of fluid in IRWST 15 IRWST CLDP-701 in. Collapsed liquid level in IRWST 16 IRWST WWITRWII, Ibm /sec. Flow from IRWST to DVI lines WWTIRWI2 . 17 IRWST IRWSTOR lbm/sec. Overflow from IRWST to sump 18 IRWST ADS 13TMR lbm/sec. Total ADS flow into IRWST 19 IRWST ADS 13TIR, MIIRWil, Ibm Integrated mass out of IRWST MIIRWI2, MIIRWlO 20 IRWST UIRWST Btu Fluid energy in IRWST 21 PRHR CLDP-802 in. Collapsed liquid level in PRHR HX O maap600\2344w 54.non:1b-loo 395 5.4.2-5 REVISION: 1 1
TABLE 5.4.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.4.2 Plot No. Component Variables Units Description 22 PRHR WWOTPRHR Ibm /sec. Measured outlet now from PRHR tube 23 Accumulator AMACC1, AMACC2 lbm Mass of fluid in accumulators 24 Accumulator CLDP-401, CLDP-402 in. Collapsed liquid level in accumulators 25 Accumulator WOUTACCl, Ibm /sec. Flow from accumulators WOUTACC2 26 Accumulator MOUTACC1, Ibm Integrated mass out of accumulators MOUTACC2 27 Accumulator UACCl, UACC2 Btu Fluid energy in accumulators 28 Pnmary sump AMPSMP lbm Pnmary sump fluid mass 29 Primary sump CLDP-901 in. Pnmary sump level 30 Pnmary sump UPSMP Blu Pnmary sump fluid energy 31 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG primary side MSSGOP1, MSSGOP2 inlet / outlet plena 32 SG MSSGHT1, MSSGHT2, Ibm Mass of fluid in SG primary side bot MSSGCrl, MSSGCr2 and cold tubes 33 SG/PRHR Cirr.201, CI'r-301, psia & SGI pressure and PRHR integrated beat QPRHR1 Btu output i 34 1 Pressurizer PZM lbm Fluid mass in pressurizer ' 35 Pressurizer CLDP 601 in. Collapsed liquid level in pressurizer 36 Pressurizer UPZ Btu Fluid energy in pressurizer 37 Surge line PLM lbm Fluid mass in surge line 38 Surge line CLDP-602 in. Collapsed liquid level in surge line 39 Surge line UPSL Btu Fluid energy in surge line 40 RPV MWRPV lbm Total fluid mass in reactor vessel 41 RPV DCM lbm Fluid mass in downcomer 42 RPV LDP0lDC in. Collapsed liquid level in downcomer compared to various reference elevations 43 RPV MWOIRPV lbm Fluid mass in lower plenum 44 RPV MWO3RPV lbm Fluid mass in core region 45 RPV LDP03RPV in. Collapsed liquid level in core 46 RPV RPVAVDF2 Core exit void fraction 47 RPV RPVAQOU2 Core exit quality 48 RPV MWO6RPV lbm Fluid mass in the upper plenum 49 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 50 RPV MWO8RPV lbm Fluid mass in the upper head 51 RPV LDP08RPV in. Collapsed liquid level in the upper bead maap6co2344.-54.noa:Ib-ioo395 5.4.2-6 REVISION: 1
- - . _ - -. _ - - .~_.. -. -. - .-. . . -... _-.. - .- _ _ - . ._ - . - - - . -
O TABLE 5.4.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.4.2 Plot No. Component - Variables Units Description 52 RPV URPV Btu Total fluid energy in reactor vessel 53 RPV RPVXE, RPVASL2 in. Level of Tsat line 54 RPV RPVPab, RPVAPab2, kW Heated rod power above and below RPVPWR Tsat level and total 55 RPV RPVRXV,RPVASOU2 lbm/sec. Core steam generation rate 1 56 RPV RPVALIN2 lbm/sec. Calculated core tiow 57 RPV HTMXRPV, ST08RPV 'F Maximum clad temperature and saturation temperature 58 Hot leg MWHL1, MWHL2 lbm Water mass in hot legs 59 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs 60 Cold leg CLlWMS, CL2WMS, Ibm Water mass in cold legs ) CL3WMS, CL4WMS 61 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs , CL3VMS, CL4VMS ' 62 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4s, and break 63 ADS and break BRKTIVF, AD13TIVF, Ibm Totalintegrated vapor flow for ADS AD41TIVF, AD42TIVF and break 64 ADS and break BRKTILF, ADl3TILF, Ibm Totalintegrated liquid flow for ADS AD41TILF, AD42TILF and break 65 ADS and break ADS 13SVR, Ibm /sec. Vapor flow out ADS 13 and ADS-4 ADS 41SVR, ADS 42SVR 66 ADS and break ADS 13SLR, ihm/sec. Liquid flov out ADS l-3 and ADS-4 ADS 41SLR, ADS 42SLR 67 ADS and break BRKSSVR lbm/sec. Vag x flow out of break 68 ADS and break BRKSSLR lbm/sec. Liquid flow out of break 69 ADS and break BRKSPEl ADS 13El, Btu Integrated fluid energy for ADS 13, ADS 41El, ADS 42El ADS-4, and break 70 Mass balance TOTMASS lbm Total system mass inventory 71 Mass balance PRIMMASS, Ibm Measured primary system inventory and PRIMASS2 value from mass balance l 72 Mass balance MERROR Ibm Mass balance error 73 Mass balance MIN, MOUT lbm Integrated mass flow in and out of J SRCMASS primary system and source mass { 74 Energy balance Various Blu Components of energy balance O mMpue2ws4.noo: b-too395 5.4.2-7 REVISION: 1
O TIIE FIGURES LISTED IN TABLE 5.4.2-1 ARE NOT INCLUDED IN Tills NONPROPRIETARY DOCUMENT l l l l 9' O m:Vau344w 54.noa:luco395 5.4.2-8 REVISION: 1
V 5.4.3 Long-Term Transient he long-term transient started with initiation of IRWST injection, covered the transition from IRWST to sump injection, and provided information on the LTC response of the AP600. For the 2-in. cold-leg balance line break, Matrix Test SB09, the long-term transient analyzed begins at [ ]*** seconds and extends to the end of the test at [ ]*** seconds. 'Ihe behavior of the test facility during this period of the transient is discussed in this subsection using the plot package detailed in Table 5.4.3-1.
'Ihis analysis concentrates on the components of the primary system that remained active during the L'IC phase, that is, the RPV, the hot legs, ADS-4, the sumps, and the IRWST.
Hermal-hydraulic phenomena of interest for the long-term transient are:
- Maintenance of core cooling and removal of energy from the primary system.
Level oscillations (from [ l'** seconds. he were system wide level and pressure oscillations, which are discussed further in Subsection 6.1.3). 5.4.3.1 Maintenance of Core Cooling Mass Injected into Primary System (q '\.) Total DVI line flow, CMT flow, and IRWST flows are shown in Figures 5.4.3-6 and 5.4.3-7. Flow from the primary sump is shown in Figure 5.4.3-19. From around [ ]'** seconds, there was a contribution to the DVI flow from the CMT-2 as the previously refilled CMT-2 drained. CMT-1 did not refill during test SB09. During thc pre-sump injection phase of the transient, IRWST flow proceeded at a gradually decreasing rate with the effect of the primary system oscillations superimposed. At [ ]*** seconds, flow from the primary sump began through the main injection valves, which opened as the IRWST has reached the low-low level set-point. This resulted in a reversal of flow through the IRWST injection line-1 almost equal to the IRWST flow into DVI-2. De net result was that the IRWST level decreased less than [ ]'A' in, between the inception of sump injection flow through the primary injection valves and the end of the transient. He initial sump injection through the check valves around the main injection valves at [ ]'** seconds decreased the IRWST rate by about [ ]*A' percent in each of the DVI lines. Reactor Pressure Vessel and Downcomer Response The effect of water inflow on the average measured downcomer fluid temperatures, core inlet and core outlet temperatures, and heater rod temperatures during the long-term phase of the transient is shown p in Figures 5.4.3-4,5.4.3 5, and 5.4.3 38. Figure 5.4.3-4 shows that there was a general increase in 'Q average downcomer fluid temperatures during the long-term transient. By the end of the test, this maap6ammS4.non:ib ioo395 5.4.3-1 REVISION: 1
l average temperature reached a value about [ ]'6" *F below saturation. Figure 5.4.3-5 shows that the core exit temperature remained at or near saturation for the majority of the long-term transient after [ ]'6" secs. Figures 5.4.3-34 to 5.4.3-36 show that the DVI line flow method described in Section 4.11 indicates that a small level of boiling was maintained after [ ]'6# seconds into the l transient. Nevertheless, the level of boiling was small, and the test results showed that the inflow 1 from the IRWST and sumps was sufficient to maintain cooling. 1 Figure 5.4.3-38 shows that there were no significant excursions in heated rod temperatures throughout ) the long-term transient therefore, sufficient core inventory and flow was maintained through this phase l of the transient to remove the decay heat generated. For significant portions of the transient, a two- j phase mixture was present in the core and upper plenum regions, i ne following discussion tracks tne variation in water level and mass throughout the reactor vessel and downcomer. The mass and level for the core region are shown in Figures 5.4.3-28 and 5.4.3-29. The collapsed liquid level in the core indicated that the heated rods, were always covered with 9 single- or two-phase mixture. During the sump injection stage of the transient (beyond [ ]'6# seconds), the collapsed liquid level remained just below the top of the heated rods, and the core void fraction was [ ].'6# The reduction in the core collapsed liquid level following the start of sump injection produced no marked impact on core cooling; in this test the sump water was relatively cold (Figures 5.4.3-4 and 5.4.3-5). During sump injection the calculated steam generation rate was at a maximum of about [ ]'6* lbm/sec. (Figure 5.4.3-36). The collapsed liquid level in the upper plenum region is shown in Figure 5.4.3-32; the level remained between the hot leg and DVI line elevations throughout the transient. Figure 5.4.3-33 shows the mass of water in the upper head, which remained below [ ]'b' lbm from the inception of sump injection until the end of the test. The mass of water in the RPV is shown in Figure 5.4.3-25. For the sump injection portion of the long-term transient, the reactor vessel water mass reached an equilibrium value of about [ ]'6# lbm, which is [ ]'6# percent of the initial vessel water inventory. From [ ]'6' seconds, oscillations in vessel inventory were observed. Figures 5.4.3-51 to 5.4.3-56 illustrate these oscillations using plots on a restricted time frame from [ ]'6' seconds. These oscillations are observed in primary system measurements from the upper plenum to the ADS-4 flows. He SB09 oscillations have a less uniform period during this time interval than do the oscillations observed in other tests. The oscillations in ADS flow lagged behind those in the upper head pressure by around [ ]'6' seconds. These oscillations and possible mechanisms for their production are discussed further in Subsection 6.1.3. The mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.4.3-26 and 5.4.3-27. The collapsed liquid level remained above the cold leg midplane until [ ] seconds when it started to fa'l to an elevation below the center of the hot legs. This was close to the time that CMT-2 completed draindown and corresponds to the time the cold legs began to drain m:\ap600044w-54.noo:Ib-100395 5.4.3-2 REVISION: 1
- _ . - _ . - .- - - . -. - . - - . . . - .. .- - ..- ~.. - . . - _ - . - _ . _ _ . _ .
l l l l f (Figure 5.4.3-41) and the ADS-4 valve liquid flow increased (Figure 5.4.3-44)~ Bere was no effect on downcomer level resulting from the start ofinjection through the primary sump valves. 5.4.3.2 Energy Transport from the Primary System
- During the long-term transient, energy continued to be deposited in the primary system from the heated rods, metal, and fluid flowing from the primary sump. He SGs and PRHR remained inactive throughout this phase of the transient, and the principal path for energy out of the primary system was via the ADS-4 valves.
Integrated mass flow from the primary system via the ADS and the break is shown in Figure 5.4.3-43. During the L'Ir phase of the transient, the only significant outflow was through the ADS-4 valves. After [ ]'6# seconds, there was some reverse flow through the break as is indicated by the reduction in the integrated flow shown in Figure 5.4.3-43; the break flow integral returned to its pre- [ ]'b' second value prior to sump injection. - During the sump injection phase of the transient, outflow from the ADS-4 valves was liquid. By the end of the test, liquid flowed out through these ! valves at a combined average rate of [ ]'6* lbm/sec. Figure 5.4.3-36 shows the calculated steam generation rate as determined by the DVI line flow . i method. During the sump injection phase of the transient, steam was generated at up to [ ]'6' lbm/sec., although the steam vortex meters indicate little or no flow out of the ADS-4 valves. However, there are two indications that steam is leaving the primary system by this route. l Figure 5A3-46 shows total measured system fluid inventory. During this phase of the transient after the start of primary sump injection (from [ ]'6# seconds, that is, l -when core steam generation was most significant), the total system inventory fell by about [ ]'6# lbm. This amount corresponds to a steam flow rate of [ ]'6* lbm/sec., which would not have been detected by the vortex meters. i j
- Examination of the fluid thermocouples on the outlet of the ADS-4 valves indicates that l temperatures remained at or above saturation temperature following the start of sump injection.
l i It was not possible for all the steam generated in the core to flow from the upper head to the downcomer via the bypass holes (Subsection 6.1.3). Therefore, steam was leaving the primary system via ADS-4. Figure 5.4.3-50 shows all the components to the system energy balance. Further discussion of steam loss from the primary circuit is provided in the mass and energy balance discussions of Section 6.2. i I
- o i
mAap6002344w-54.noo:lb-100395 REVIs!ON: 1 5.4.3-3 l _ . , -_ ___ __ . , ._
TABLE 5.4.31 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.4.3 LONG-TERM TRANSIENT Plot No. Component Variables Units Description 1 RPV RPVPWR W Core power 2 Pnmary sump TSMPil. TSMPI2 F Sump injection line temperatures 3 DVI TDVIL1, TDVIL2 'F DVI line temperatures 4 RPV TOIDC, T02DC, T03DC, 'F Water and saturation temperatures in STOIDC downcomer 5 RPV TOIRPV,TU8RPV, 'F Core inlet / outlet temperature, ST08RPV saturation temperature 6 D Vl-1 WWIDVIL1, Ibm /sec. Individual components and total flow WWTIRWII, in DVI-l WWTIRW13 7 DVl-2 WWTDVIL2, Ibm /sec. Individual components and total flow WWTIRW12, in DV12 WWIIRWI4 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 10 IRWST 1RWST Ibm Mass of fluid in IRWST 11 1RWST CLDP-701 in. Collapsed liquid level in IRWST < 12 IRWST UIRWST Btu Fluid energy in IRWST 13 Pnmary sump AMPSMP lbm Primary sump fluid mass l 14 Ptimary sump CLDP-901 in. Pnmary sump level 15 Pnmary sump UPSMP Btu Pnmary sump fluid energy j 16 Secondary sump AMSSMP lbm Secondary sump fluid mass l 17 Secondary sump CLDP-902 in. Secondary sump level l 18 Secondary sump USSMP Btu Secondary sump fluid energy l 19 Pnmary sump WSTSMPET, WWTSMPIT Ibm /sec. Pnmary sump steam and liquid I injection rate 20 Pnmary sump MISMPII, MISMPI2, Ibm Integrated primary sump and IRWST MISMPIT, MIIRWT flows 21 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG side inlet / outlet MSSGOPI, MSSGOP2 plena 22 Surge line PLM lbm Fluid mass in surge line 23 Surge line CLDP-602 in. Collapsed liquid level in surge line 24 Surge line UPSL Btu Fluid energy in surge line 25 RPV MWRPV lbm Total fluid mass in reactor vessel O m:\ap600chstaan.Ib-too395 5.4.3-4 REVISION: 1
p TABLE 5.4.31 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.4.3 LONG TERM TRANSIENT Plot No. Component Variables Units Description 26 RPV DCM lbm Fluid mass in downcomer 27 RPV LDP0lDC in. Collapsed liquid level in dowmcomer compared to various reference elevations 28 RPV MWO3RPV lbm Fluid mass in core region 29 RPV LDP03RPV in. Collapsed liquid level in core 30 RPV RPVAVDF2 Core exit void fraction 31 RPV RPVAQOU2 Core exit quality 32 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 33 RPV MWO8RPV lbm Fluid mass in the upper head 34 RPV RPVASL2 in. Level of Tsat line 35 RPV RPVAPab2, RPVPWR kW Heated rod power above and below Tsat level and total 36 RPV RPVASOU2 lbm/sec. Core steam generation rate 37 RPV RPVALIN2 lbm/sec. Calculated core flow 38 RPV HTMXRPV, 'F Maximum clad temperature, saturation (n) ST08RPV temperature and delta 39 Hot leg MWHL1, MWHL2 lbm Water mass in hot legs 40 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs 41 Cold leg CL1WMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 42 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 43 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4, and break 44 ADS and break ADS 13TLR, ADS 41TLR, Ibm /sec. Liquid flow out ADS 1-3 and ADS-4 ADS 42TLR 45 ADS and break BRKSTLR lbm/sec. Liquid flow and total flow out of break 46 Mass balance TOTMASS lbm Total system mass inventory 47 Mass balance PRIMMASS, PRIMASS2 lbm Measured pnmary system inventory and valve from mass balances 48 Mass balance MERROR lbm Mass balance error 49 Mass balance MIN, MOUT SRCMASS lbm Integrated mass flow in and out of pnmary system and source mass 50 Energy balance Various Btu Component of energy balance 51 ADS-4 ADS 41TLR, ADS 42TLR Ibm /sec. Oscillations in ADS-4 liquid flow 52 Surge line CLDP-602 in. Oscillations in surgeline level mAapum344w.54. con:lt> 100395 5.4.3-5 REVISION: I
l l l l l TABLE 5.4.31 (Continued) @\ ! OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.4.3 l LONG TERM TRANSIENT Plot No. Component Variables Units Description 53 RPV CPT-107 psia Oscillations in upper head pressure 54 RPV CLDP-ll3 in. Oscillations in upper plenum level 55 RPV LDP03RPV in. Oscillations in core level 56 RPV LDP01DC in. Oscillauons in downcomer level l 9 9 maap60&2344w.54.non:ib.ioo395 5.4.3-6 REVISION: 1
4 3 4 4 i. 4-3- i j i-i 4 i i i I THE FIGURES LISTED IN TABLE 5.4.31 ! ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT i i i i i i l ( 4 m:w344 -54..on:ib-too995 5.4.3-7 REVISION: 1
.-.,w-- , -
. . - - - , , , , - , , _ ...--, ,,-,w,,., e.a., .w..n,mwwre=-.m-e-m... v .y ,m.v . vywy ---,vr--
l l l m f 5.5 Analysis of Matrix Test SB10 - J Matrix Test SB10 (OSU Test UO110) simulated the double-ended rupture of the CL-3 to CMT-1 I l balance line with long-term cooling (LTC) and without the operation of the nonsafety-related systems. The break was located on the horizontal section of the balance line before the vertical rise to CMT 1. Except for the break size, this test was similar to SB09, including the simulated failure of one of the ADS-4 lines. Changes to the OSU facility since the performance of SB01 are noted in the Final Data Report.* The analysis of Matrix Test SB10 is divided into three subsections as follows: Facility performance is discussed in Subsection 5.5.1. It provides a brief outline of the response of the test facility; further details are available in the Final Data Report.* The short-term transient for SB10 encompassed the start of the simulation up to [ ]** seconds. This period includes blowdown, natural circulation ADS and initial IRWST stages of the transient. The analysis of the long-term transient SB10, encompassed the time frame from [ ]** seconds to the end of the test. This phase of the transient includes IRWST injection j (
\
and covered the transition to sump injection. The long-term transient actually starts at IRWST injection, which is discussed as part of the short-term transjent. Between the end of the short-term transient and [ ]** seconds, the system remained relatively inactive with the exception of the CMT-2 refill. At [ ]*6* seconds, CMT-2 began to refill; CMT-1 did not empty until [ ] seconds and [ ]** during the SB10 transient. CMT refill phenomena are discussed further in Subsection 6.1.1, and the discussion of the long-term transient provided in this section begins at [ ]** seconds. The discussion of the short and long-term phase of the transient focuses on important thermal-hydraulic phenomena identified in the PIRT (Table 1.3-1). Key indicators of the quality of the analysis on which this discussion is based are the mass and energy balance results. These are discussed in detail in Subsections 6.2.2 and 6.2.3. Q. m:Wm344.-55.oon: b.noo395 5.5-1 REVISION: 1
5.5.1 Facility Performance i V j The performance of the OSU test facility during Matrix Test SB10 in reference to the five transient phases is outlined in the following:
- Blowdown
- Natural circulation l
- ADS
- IRWST injection
* %mp injection The overall performance of the facility during the transient is shown in Figures 5.5.1-1 to 5.5.14.
Figure 5.5.1-1 shows the pressurizer pressure throughout the test with various phases and operating components delineated on the figures. De time scale was reduced for clarity since there were only small changes in system pressure during the long-term phase of the transient. Figure 5.5.1-2 shows the total DVI line flow and its composition from the various sources at each time in the transient. Figure 5.5.1-3 shows the calculated core steam generation rate throughout the test. Figure 5.5.1-4 shows the variation in average measured core outlet temperature and peak clad temperature relative to the core outlet saturation temperature. p Figures 5.5.1-1 and 5.5.1-2 show that there was a continuous flow of water to the core from the Q passive safety-related systems throughout the transient. Once initiated, the ADS rapidly depressurized the primary system and thus enhanced the CMT and accumulator injection flow rates. Ultimately, the ADS-4 valves reduced the system pressure sufficiently to allow gravity-driven IRWST injection to commence. The passive injection systems operation overlapped so that as one source of water drained the next was available to continue the cooling process. The level of steam generation in the core and the response of the average measured core outlet fluid temperatures and maximum clad temperatures is shown in Figures 5.5.1-3 and 5.5.1-4. These figures show that the cooling flow prevented core heatup, and the core remained covered. The core remained subcooled for large periods of the transient, and when steam was produced the rate of generation remained below the rate at which water was delivered to the core. 5.5.1.1 Blowdown Phase The blowdown phase began at time zero when the break was initiated and continued until the primary circuit pressure was in equilibrium with the secondary-side pressure at about [ ]'6' seconds. During this phase of the transient, cooling flow was provided from the intact CMT, while the CMT with a broken balance line injected no mass. CMT-2 remained in the recirculation mode until the end of the blowdown phase, and heat was removed from the primary circuit via the steam generators (SGs). The pressurizer and surge line completely drained at [ ]'6' and [ ]'b' seconds respectively. O mwsoocwS5.non:ib.ioo395 5.5.1 1 REVISION: 1
5.5.1.2 Natural Circulation Phase In this LOCA simulation, single and two-phase natural circulation initially continued the gradual reduction in systerr pressure that is characteristic of blowdown: later in this phase of the transient, the rate of depressurization increased significantly once all the SG tubes had almost completely drained at about [ ]'** seconds. The tubes in SG-2 completed draining almost [ ]'** seconds earlier than those in SG-1 because primary circuit mass was drawn to loop-1, where the reistively large balance line break was located. After [ ]**# seconds, heat removal from the primary system continued via the PRHR and the break. In response to the double-ended balance line break, CMT-2 transitioned into a rapid draindown at [ ]'** seconds, and the falling CMT level reached the ADS low-level setpoint so that the ADS-1 valve began to open at [ ]'A' seconds. Both accumulators had begun to inject at [ ]'** seconds and delivered at a rate of about [ J'** lbm/sec. before ADS actuation due to the depressurization caused by the break. 5.5.1.3 Automatic Depressurization System Phase I ADS-1 actuation was followed by ADS-2 and ADS-3 [ ]'** and [ ]'** seconds later. Coinciding with initiation of ADS-2 an increased accumulator injection rate was observed. The influx of cold water combined with increased venting via the ADS led to an even more rapid depressurization of the primary circuit. A::tuation of ADS-4 at [ ]'6# seconds completed the depressurization and allowed the IRWST to begin injecting at [ J'** seconds via DVI-2 and [ ]'** seconds via DVI-1. During the rapid accumulator injection, increased flow path resistance reduced flow out of CMT-2, and actually stopped it after ADS actuation. No flow was observed out of CMT-1 until about [ ]'** seconds after IRWST-1 injection had begun. Since both the CMT and IRWST lines attached to DVI-1 were open to containment, and the IRWST had greater gravity head, earlier injection through IRWST 1 was expected for this double-ended balance line break. As the accumulators drained, CMT-2 flow resumed at rates approaching [ ]'*# lbm/sec. Because the CMT-1 balance line contained the break, CMT-1 injection was delayed until after IRWST-1 and IRWST-2 injection had been achieved for almost [ ]'** seconds. CMT-1 flow continued concurrently with IRWST 1 injection thereafter. The minimum RPV mass inventory of [ ]'** lbm occurred shortly after accumulator injection began. l Actuation of ADS-1 and ADS-2 rapidly refilled the pressurizer as water and steam flowed out of the l ADS. The pressurizer gradually drained by [ ]'** seconds. 5.5.1.4 In-Containment Refueling Water Storage Tank Injection IRWST injection was the transition from the short- to long-term phase of the transient. Initially, IRWST injection began through the IRWST-1 DVI line, and flow through IRWST-2 started a few seconds thereafter. Both of these flows exhibited a gradual increase as the driving head between the
'IRWST tank and the RCS increased. This pressure differential increased because RCS pressure ,
decreased as the core steam generation decreased from [ ]'** lbm/sec. at IRWST-2 initiation to zero ma.p600c3m.55..on:ib-ioo395 5.5.1-2 REVISION: 1
by [ ]'6' seconds. Maximum IRWST flow was established a short time later, and afterward, it gradually decreased with the decrease in pressure differential as the IRWST continued to drain. The influx of water from the IRWST was enough to keep the core subcooled until [ ]'6" seconds. Steam was subsequently generated in the core for the remainder of the transient. Oscillations in pressure and level were observed in the primary circuit between [ J'6' and [ ]'6# seconds. 5.5.I.5 Sump Injection Sump flow began at [ ]'6' seconds through the check valves around the DVI-2 main sump injection valve since the driving head from the sump was sufficient for flow to initiate. Sump injection flow in DVI-1 began a short time later. Sump injection through DVI l began to enter the IRWST once the main sump injection valves opened; at that time DVI-2 exhibited a corresponding increase in the flow out of the IRWST, which meant that there wss no increase in IRWST inventory. Flow through the main sump injection valves began when those valves opened at about [ ]'6" seconds. O O mAap600G344w 55.non:lb-100395 5.5.1-3 i!EVISION: 1
l l l TABLE 5.5.1 1 OSU TEST ANALYSIS PLOT PACKAGE FOR SUBSECTION 5.5.1 Plot No. Component Variables Units Description 1 Pressurizer CI'r-604 psia System pressure and event history 2 Water WWTDVII+WWTDVI2, Ibm /sec. Total of CMT, accumulator, IRWST, injection WOUTACCl+WOLTTACC2, and sump injection flows WWTIRWII+WWTIRW12, WWTSMPlT 3 Reactor RPVASOU2 lbm/sec. Steam generation in reactor vessel vessel 4 Reactor T08RPV, HTMXRPV, TS AT F Reactor vessel outlet temperature, vessel maximum clad temperature and fuel exit saturation temperature e i e m:\ap60cc3h55.aoo:1b-too395 5.5.1 4 REVISION: 1
l 1 1
- i j
THE FIGURES LISTED IN TABLE 5.5.11 ! ARF NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT 1 l i 1 l l
- 1 i
- i i 1 1 1 I I 1 I 1 i 1 l l 4 5 mwan2344w.55.noo: wim395 5.5.1-5 REVISION: 1 4
.--,-.e,-n,. - --, , ,, .a --,-r,-,., - - - - , , , , - - - , -
/
( 5.5.2 Short Term Transient . For the double-ended balance line break LOCA simulation, Matrix Test SB10, the short-term transient encompassed the time frame up to 2000 seconds. As shown in Figure 5.5.1-1, this period included the full depressurization of the facility through all four stages of the ADS, together with CMT and accumulator injection plus the initial stages of IRWST injection. He variation in mass, energy, pressure, and temperature throughout this stage of the transient are illustrated in the plot package
- outlined in Table 5.5.2-1. The plots concentrate on the primary system, including the accumulators, l CMTs, IRWST, primary sump and flow from the primary system via the ADS, break, and IRWST overflow, i
l There were two principal parameters to be examined for the short-term transient: i
= Adequate flow from the passive systems to the reactor vessel must be maintained.
- Adequate flow into the core must be maintained to ensure that decay heat was removed from the simulated fuel rods without a temperature excursion.
l l Rese parameters are addressed in the following discussion. , I IOl 5.5.2.1 Maintenance of Core Cooling Mass Injected to the Primary System 1 l l Figures 5.5.2-6 and 5.5.2-7 show the combined effect of the injection flow for the transient. Separate i plots of the individual contributions to the total flow can be located by consulting the plot package i index given in Table 5.5.21.
- Figures 5.5.2-5 and 5.5.2-6 show how the CMTs, accumulators, and IRWST supplied a continuous flow of water to the core. During the first [ ] seconds, cooling flow was provided primarily by CMT-2, since CMT-1 flow was negligible due to the break in its balance line. By [ ] seconds, j CMT flow was supplemented by flow from both accumulators. He rate of flow from Chfr-2 was stopped for awhile by the accumulator injection. Accumulator flow produced the maximum DVI injection rates during the entire transient, with values of [ ]'6" lbm/sec. and above in each DVI line.
l Following the end of accumulator injection, CMT-2 once again provided the cooling flow into the l RPV. IRWST-1 injection began [ ]'6' seconds before the CMT-2 draindown was completed. Since I continuous IRWST injection through both DVI lines began before CMT-2 had fully drained, there was no period of the short-term transient when the passive safety-related systems failed to provide flow to the RPV. i O mAap60cemw-55.non:1b-too395 5.5.2-1 REVISION: 1
Reactor Pressure Vessel and Downcomer Behavior . The effect of water flow on the average measured core inlet / outlet temperatures and peak clad temperatures during the short-term phase of the transient is shown in Figures 5.5.2-3 and 5.5.2-57. The core outlet temperature first reached the saturation point at [ ]*** seconds. The core outlet temperature then remained at saturation for almost the entire interval until about [ ]'6# seconds when the influx of water from the CMT and IRWST became sufficient to again subcool the core. The core then remained subcooled beyond the end of the short-term transient. Figure 5.5.3-57 shows that there were no significant excursions in heated rod temperatures throughout the short-term transient; therefore, sufficient core inventory and flow were maintained throughout this phase of the transient to remove the simulated decay heat generation. For significant portions of the transient, a two-phase mixture was present in the core and upper plenum regions, wi:h core bo!"ng kept at a low level. The following discussion tracks the variation in water level and mass throughout the reactor vessel and 1 downcomer. The mass and level for the core region are shown in Figures 5.5.3-44 and 5.5.3-45. The l collapsed liquid level in the core indicated that the heated rods remained covered with a single- or two-phase mixture throughout the short-term transient. The minimum core inventory of [ ]'6* h
]
occurred at about [ ]'6# seconds into the transient, before the initial accumulator injection was fully established. As seen in Figure 5.5.2-45, the collapsed liquid level dropped to [ ]'** in. below the top of the heated rod length during this phase of the transient. The average void fraction of the core two-l phase mixture may be estimated by dividing the measured core collapsed liquid level by the [ ]'6# in. I heated rod length. In this test, the minimum collapsed liquid level corresponds to a core void fraction of[ ].' 6
- By the end of the short-term transient, the effect of IRWST injection ended core boiling (Figure 5.5.2-55) and the core was again water-solid.
The collapsed liquid level in the upper plenum region and the associated fluid mass are shown in Figures 5.5.2-49 and 5.5.2-48; the level decreases, is replenished by accumulator injection, decreases again, and is replenished again by IRWST injection. In the [ l'** second time interval, liquid splashed from the upper plenum to the downcomer and back again twice; the downcomer level twice reached the upper head bypass plate during this rapid condensation event, as discussed in the Final Data Report.m F gures 5.5.2-50 and 5.5.2-51 show that the upper head, which had drained by the time of significant accumulator injection, refilled along with the upper plenum during the initial rapid condensation event of SB10. The upper head eventually drained, only to refill later during the short-term transient due to another rapid condensation event at [ ]'A' seconds in the downcomer. This refill occurred at the time of maximum combined CMT/IRWST injection in the downcomer. The mass of fluid and collapsed liquid level in the RPV downcomer is shown in Figures 5.5.2-41 and 5.5.2-42. The downcomer collapsed liquid level fell to the cold-leg elevation during the first [ ]*** seconds, and rapidly refilled to the bypass plate during the interaction with the upper plenum. The level decreased and then recovered in the [ ]'** second time interval as the upper head maap600(2344w-55.noo:ib-too395 5.5.2-2 REVISION: 1
. - - - - - . . - . . ~ . - - . - . - . . . . - - - - - . . - - - - . - _ ~ . . . . - - . . . - .
i l'
, partially filled and then drained. IRWST injection maintained the collapsed liquid level within the cold-leg pipe perimeter after [ ]'** seconds in the shon-term transient.
5.5.2.2 Energy Transport from the Primary System Following the break, energy was deposited in the primary circuit fluid by the heater rods to simulate decay heat and the primary circuit metal as it cools down. Some fluid energy was lost to the i atmosphere and out of the break. Excess energy must be removed from the primary system to prevent excessive fluid and heater rod temperature excursions. In the AP600, heat removal is designed to be achieved by a combination of the SGs and the PRHR plus the ADS. Steam Generator and Passive Residual Heat Removal Heat Transfer During normal operation, most of the primary system heat was removed via the SGs; however, once the RCPs tripped, the reduced system flow caused a reduction in primary- to secondary-side heat transfer. De SGs were only available as heat sinks until primary-side system pressure dropped to that , of the secondary side. The two sides of the SG were in thermal equilibrium for a time, then tim secondary side became a heat source to the primary system, and the SG tubes drained rapidly. The PRHR is designed to remove heat from the primary system once the cafety eignal opens the PRHR isolation valve. The PRHR will continue to remove energy via natural circulation after the SGs are thermally isolated until ADS actuates. Once the ADS is actuated, ADS 1-3 becomes the predominant s path for the removal of energy from the primary system. Figure 5.5.2-33 shows the SG pressure equalization together with the PRHR integrated heat transfer as represented by the IRWST fluid energy after allowing for the contribution from ADS 1-3 inflow. As shown, heat was transferred to the secondary side of the SGs for only the first [ ]'A' seconds of the test. PRHR heat removal began about [ ]'A' seconds into the test, and the PRHR was responsible for all the IRWST heat-up until ADS-1 activation, after which PRHR heat transfer was significantly reduced. During the natural circulation phase, the PRHR transferred heat to the IRWST at an average rate of about [ ]'** Btu /sec. l l Energy Transport via the Break and Automatic Depressurization System The mass flow rate from the primary system via the break is shown in Figures 5.5.2-67 and 5.5.2-68. Immediately, the break flow rose to a maximum value near [ ]'A' lbm/sec. of water, then diminished as system pressure fell. At [ l'A' seconds, mass flow out of the break (Figure 5.5.2-62) decreased as the SG draining ended. Figure 5.5.2-71 shows the equivalent trend. Then, following the initiation of ADS 1-3, flow through the break dropped somewhat as steam and liquid flow through the ADS 1-3
- valves began; ADS 1-3 then became the primary mass release path.
! Beyond[ ]'A' seconds, flow through the break continued as the downcomer level refilled to the cold-leg elevation; refill was occurring due to IRWST and CMT-2 injection. Liquid flow through the macoac3m.55.noa:ib.ioo395 5.5.2-3 REVISION: 1 I
break in the balance line continued at a declining rate until almost [ ]* seconds, afterward flow through the break was minimal for the rest of the short-term transient, i l Integrated mass flow from the primary system via the ADS and the break is shown in Figure 5.5.2-62. The corresponding integrated energy flow is shown in Figure 5.5.2-69. The total system inventory plot given in Figure 5.5.2-70 indicates that only [ ]'** lbm ofinventory left the system during the l short-term transient. Components of the energy balance are shown in Figure 5.5.2 '4. l O l l l l 1 l l I l 9 m:Wa344w-55.non: b-too395 5.5.2-4 REVISION: 1
. ~ - - ~ . - . . - . . - - - - - . . _ . - - - . - . - . - . . . . . . - . . - . . . . - - - .
i-a- j ? 'f
~
!\ TABLE 5.5.21 l l OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECI' ION 5.5.2
- Plot No. Component - Variables Units Description l 1 Pressurizer CirT-604 psia System pressure 2 RPV RPVPWR kW Core power l
l 3 RPV TOIRPV, T08RPV, 'F Core inlet / outlet temperature, l l ST08RPV saturation temperature ! l , 4 SG CI'T-201, CI'T-204, psia Pnmary and secondary pressures in SG !' CPT 301, CI"T-302 { ! 5 DVI-I WWTDVIL1, Ibm /sec. Individual components and total flow in
- WWTIRWil, DVI-l WOUTACCl, l l
- I WWTIRWI3 6 DVI-2 WWTDVIL2, Ibm /sec. Individual components and total flow in WWTIRW12, DVI2 l WOUTACC2, l WWTIRWI4 7 CMT AMCMTIB, Ibm Fluid mass in CMTs (excludes balance AMCMT2B lines) )
8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT MIWDVIL1, Ibm Integrated mass out of CMTs I MIWDVIL2 10 CMT WWTDVIL1, Ibm /sec. Flow out of CMTs . WWIDVIL2 11 CMT WOUTCLBl. Ibm /sec. Flow into CMTs WOUTELB2 2 12 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 13 CMT UCMT1, UCMT2 Btu Fluid energy in CMTs 14 IRWST IRWST lbm Mass of fluid in IRWST 15 IRWST CLDP-701 in. Collapsed liquid level in IRWST 16 IRWST WWTIRWII, Ibm /sec. Flow from IRWST to DVI lines WWI1RWI2 17 IRWST 1RWSTOR lbm/sec. Overflow from IRWST to sump 18 IRWST ADS 13TMR Ibm /sec. Total ADS flow into IRWST 19 IRWST ADSl3TIR, MIIRWII, Ibm Integrated mass out ofIRWST MIIRWI2, MIIRWIO 20 IRWST UIRWST Btu Fluid energy in IRWST 21 PRHR CLDP-802 in. Collapsed liquid level in PRHR HX 1 m:W344. 55.non:Ib-100395 5.5.2-5 REVISION: 1
TABLE 5.5.21 (Continued) ! OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.5.2 Plot No. Component Variables Units Description 22 PRHR WWOTPRHR Ibm /sec. Measured outlet flow from PRHR tube 23 Accumulator AMACCl, AMACC2 lbm Mass of fluid in accumulators 24 Accumulator CLDP-401, CLDP-402 in. Collapsed liquid level in accumulators 25 Accumulator WOUTACC1, Ibm /sec. Flow from accumulators WOlJTACC2 26 Accumulator MOUTACCl, Ibm Integrated mass out of accumuiators MOUTACC2 27 Accumulator UACCl, UACC2 Btu Fluid energy in accumulators 28 Pnmary sump AMPSMP lbm Primary sump fluid mas 29 Pnmary sump CLDP-901 in. Prunary sump level 30 Prunary sump UPSMP Blu Primary sump fluid energy 31 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG primary side j MSSGOPI, MSSGOP2 inlet / outlet plena 32 SG MSSGHT1, MSSGHT2, Ibm Mass of fluid in SG pnmary side hot , MSSGCTI, MSSGCT2 and cold tubes 1 33 SG/PRHR CPT-201, CPT-301, psia & SGI pressure and PRHR integrated beat l QPRHR1 Btu output 34 Pressurizer PZM lbm Fluid mass in pressurizer j 35 Pressurizer CLDP401 in. Collapsed liquid level in pressurizer 36 Pressurizer UPZ Btu Fluid energy in pressurizer 37 Surge line PLM lbm Fluid mass in surge line 38 Surge line CLDP-602 in. Collapsed liquid level in surge line 39 Surge line UPSL Btu Fluid energy in surge line 40 RPV MWRPV lbm Total fluid mass in reactor vessel 41 RPV DCM lbm Fluid mass in downcomer 42 RPV LDP01DC in. Collapsed liquid level in downcomer compared to various reference elevations 43 RPV MWO1RPV lbm Fluid mass in lower plenum 44 RPV MWO3RPV lbm Fluid mass in core region 45 RPV LDP03RPV in. Collapsed liquid level in core 46 RPV RPVAVDF2 Core exit void fraction 47 RPV RPVAQOU2 Core exit quality 48 RPV MWO6RPV lbm Fluid mass in the upper plenum 49 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 50 RPV MWO8RPV lbm Fluid mass in the upper head 51 RPV LDP08RPV in. Collapsed liquid level in the upper bead maap60m2h55.non: b-ioons 5.5.2-6 REVISION: 1
N. TABLE 5.5.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.5.2 Plot No. Component Variables Units Description 52 RPV URPV Btu Total fluid energy in reactor vessel l 53 RPV RPVXE, RPVASL2 in. Level of Tsat line l 54 RPV RPVPab, RPVAPab2, kW Heated rod power above and below l RPVPWR Tsat level and total 55 RPV RPVRXV,RPVASOU2 lbm/sec, Core steam generation rate 56 RPV RPVALIN2 lbm/sec, Calculated core flow 57 RPV HTMXRPV, ST08RPV 'F Maximum clad temperature and saturation temperature 58 Hot leg MWHL1, MWHL2 lbm Water mass in hot legs 59 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs 60 Cold leg CLIWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS l 61 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 62 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4s, and break 63 ADS and break BRKTIVF, ADl3TIVF, Ibm Total integrated vapor flow for ADS I AD41TIVF, AD42TIVF and break V 64 ADS and break BRKTILF, ADl3TILF, Ibm Total integrated liquid flow for ADS AIMITILF, AD42TILF and break l 65 ADS and break ADS 13SVR, ihm/sec. Vapor flow out ADS 1-3 and ADS-4 l ADS 41SVR, ADS 42SVR l 66 ADS and break ADS 13SLR, Ibm /sec. Liquid flow out ADS 1-3 and ADS-4 ADS 41SLR, ADS 42SLR 67 ADS and break BRKSSVR Ibm /sec. Vapor flow out of break 68 ADS and break BRKSSLR lbm/sec. Liquid flow out of break 69 ADS and break BRKSPEI, ADS 13EI, Btu Integrated fluid energy for ADS 1-3, ADS 41EI, ADS 42EI ADS-4, and break 70 Mass balance TOTMASS lbm Total system mass inventory 71 Mass balance PRIMMASS, Ibm Measured primary system inventory and PRIMASS2 value from mass balance 72 Mass balance MERROR lbm Mass balance error 73 Mass balance MIN, MOUT lbm Integrated mass flow in and out of SRCMASS primary system and source mass 74 Energy balance Various Btu Components of energy balance maap600c344.-55.noo:ib.too395 5.5.2-7 REVISION: 1
O THE FIGURES LISTED IN TABLE 5.5.21 ARE NOT INCLUDED IN TIIIS NONPROPRIETARY DOCUMENT O l l 9 m:W344w 55.aon:Ib t00395 5.5.2-8 REVISION: 1
1 i D 5.5.3 Long-Term Transient ] () 1
- The long-term transient started with initiation of IRWST injection, covered the transition from IRWST to sump injection, and provided information on the LTC response of the AP600. For the double-ended cold-leg balance line (CLBL) break, Matrix Test SB10, the long-term transient analyzed runs from
[ ]'6* seconds to the end of the test to about [ ]'*' seconds. The behavior of the test facility during this period of the transient is discussed in this subsection using the plot package detailed in Table 5.5.31. This analysis concentrates on the components of the primary system that remained active during the LTC phase, that is, the RPV, the hot legs, ADS-4, the sumps, and the IRWST. The main thermal-hydraulic phenomena of interest for the long-term transient are: Maintenance of core cooling and removal of energy from the primary system. Level oscillations (from [ ]'6' mod Are were system wide level and pressure oscillations, which are discussed further in Subsection 6.1.3) 5.5.3.1 Maintenance of Core Cooling Mass Injected into Primary System ( V Total DVI line flow, CMT flow and IRWST flow are shown in Figures 5.5.3-6 and 5.5.3-7, and flow from the primary sump is shown in Figure 5.5.3-19. As shown in this figure, from about [
]'*# seconds, there was a large contribution to the DVI flow from the CMT-2 as the previously refilled CMT-2 drained. CMT-1 did not empty completely during test SB10 until [ ]'*' seconds as a result of the break location. It provided a relatively small amount of flow between
[ ]"' seconds until it emptied. i Prior to sump injection, during the IRWST injection phase of the transient, IRWST flow proceeded at a gradually decreasing rate with the effect of the primary system oscillations and CMT-1 delively superimposed. At [ ]'*' seconds, flow from the primary sump began through the main injection valves which had opened once the IRWST reached the low-low level set-point. This resulted in a reversal of flow through the IRWST injection line-1 almost equal to the IRWST flow into DVI-2. The net result was that the IRWST level decreased less than [ J'*' in, between the inception of sump injection flow through the primary injection valves and the end of the transient. The initial sump injection through the check valves around the main injection valves at [ ]'6* seconds decreased the IRWST injection rate by about [ ]'** percent in each of the DVI lines. Reactor Pressure Vessel and Downcomer Response The effect of the water inflow on the average measured downcomer fluid temperatures, core inlet and O g core outlet temperatures, and heater rod temperatures during the long-term phase of the transient is msp60m23w55.noo:ib.ioo395 5.5.3-1 REVISION: 1
shown in Figures 5.5.3-4,5.5.3-5 and 5.5.3-38. Figure 5.5.3-4 shows that there was a general increase in average downcomer fluid temperatures during the long-term transient. By the end of the test, this average temperature reached a value about [ ]'b#'F below saturation. Figure 5.5.3-5 shows that the core exit temperature remained near saturation for the majority of the long-term transient. Figures 5.5.3-34 to 5.5.3-36 show the DVI line flow method described in Section 4.11 indicating that a small level of boiling was maintained after [ ]'*" seconds into the transient. Nevertheless, the level of boiling was small, and the test results showed that the inflow from the IRWST and sumps was sufficient to maintain cooling. Figure 5.5.3-38 shows that there were no significant excursions in heated rod temperatures throughout the long-term transient; therefore, sufficient core inventory and flow was maintained throughout this phase of the transient to remove the decay heat generated. For significant portions of the transient, a two-phase mixture was present in the core and upper plenum regions. He following discussion tracks the variation in water level and mass throughout the reactor vessel and downcomer. The mass and level for the core region are shown in Figures 5.5.3-28 and 5.5.3-29. The collapsed liquid level in the core indicated that the heated rods were always covered with a single- or two-phase mixture. During the sump injection stage of the transient (beyond [ ]'** seconds), the collapsed liquid level remained just below the top of the heated rods, and the core void fraction was I [ ) .'*# De reduction in the core collapsed liquid level following the start of sump injection produced no marked impact on core cooling; in this test the sump water was relatively warm (Figures 5.5.3-4 and 5.5.3-5). During sump injection, the calculated steam generation rate was at a maximum of about [ ] lbm/sec. (Figure 5.5.3-36). The collapsed liquid level in the upper plenum region is shown in Figure 5.5.3-32; the level remained l between the hot leg and DVI line elevations during the sump injection transient. Figure 5.5.3-33 shows the mass of water in the upper head, which drained shortly after the inception of sump injection I and thereafter remained nearly empty until the end of the test. He mass of water in the RPV is shown in Figure 5.5.3-25. For the sump injection portion of the long-term transient, the reactor vessel water mass reached an equilibrium value of about [ ]'** lbm, which was [ ] percent of the initial vessel water inventory. Between [ ]'** seconds, oscillations in vessel inventory were observed. Figures 5.5.3-51 to 5.5.3-56 illustrate these oscillations using plots on a restricted time frame from [ ]'*# seconds. These oscillations are observed in primary system measurements from the upper plenum to the ADS-4 flows. The osciliations in the ADS flow lagged behind those in the upper head pressure by around [ l**# seconds. These oscillations and possible mechanisms for their production are discussed further in Subsection 6.1.3. The history of fluid mass and collapsed liquid level in the RPV downcomer during the test are shown in Figures 5.5.3-26 and 5.5.3-27. The collapsed liquid level remained above the cold-leg midplane until [ ]*** seconds, and afterward the level was at or just below the mid-level cold-leg elevation maap60em.55.non:n>.tco395 5.5.3-2 REVISION: 1
for the remainder of the transient. The cold legs did not drain (Figure 5.5.3-41) in this transient. The double-ended balance line break at an elevation above the cold legs kept the liquid level high in the downcomer and into the cold legs throughout sump injection. Liquid continued to flow through the break for the remainder of the transient (Figure 5.5.3-45). There was no effect on downcomer level resulting from the start ofinjection through the primary sump valves. 5.5.3.2 Energy Transport from the Primary System During the long-term transient, energy continued to be deposited in the primary system from the heated rods, metal, and fluid flowing from the primary sump. The SGs and PRHR remained inactive throughout this phase of the transient, and the principal path for energy out of the primary system was via the ADS-4 valves.
- Integrated mass flow from the primary system via the ADS and the break is shown in Figure 5.5.3-43.
During the LTC phase of the transient, the most significant energy outflows are through ADS-4 valves. Because the ADS-4 lines were both connected to the ADS separator for test SB10, ADS 4A flow _ shown in Figure 5.5.3-43 represents the total flow for the ADS-4 paths. After [ ]** seconds, ~ flow through the break previously identified is reflected in the integrated flow shown in Figure 5.5.3-43. By the end of the test, water flowed out through the ADS-4 valves at a combined average rate of ( ]** lbm/sec. Figure 5.5.3-36 shows the calculated steam generation rate as determJned by the DVI line flow method. During the sump injection phase of the transient, steam was generated at about [ ]** lbm/sec., indicate little or no flow from the steam vortex meters out of the ADS-4 valves. Steam left the primary circuit by this route as shown by the following. Figure 5.5.3-46 shows total measured system fluid inventory. During this phase of the transient after the start of primary sump injection (from [ ]** seconds, that is, when core steam generation was most significant, the total system inventory fell by over [ ]** lbm. This amount corresponds to a steam flow rate of ( ]** lbm/sec., which would not have been detected by the vortex meters.
- Examination of the fluid thermocouples on the outlet of the ADS-4 valves indicates that temperatures remained at or above saturation temperature following the start of sump injection.
Furthermore, as discussed in Subsection 6.1.3, it was not possible for all the steam generated in the core to flow from the upper head to the downcomer via the bypass holes. It can therefore be concluded that steam was leaving the primary system via ADS-4. Figure 5.5.3-50 shows all the components to the system energy balance. Further discussion of steam loss from the primary circuit is provided in the mass and energy balance discussions of Section 6.2. O ma.p600s2ws5.noodb ioo395 5.5.3-3 REVISION: 1
l l i TABLE 5.531 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.53 LONG TERM TRANSIENT Plot No. Component Variables Units Description 1 RPV RPVPWR kW Core power 2 Pnmary sump TSMPII, TSMPI2 'F Sump injection line temperatures 3 DVI TDVIL1, TDVIL2 'F DVI ime temperatures 4 RPV T01DC, T02DC, T03DC, 'F Water and saturation temperatures in STDIDC downcomer 5 RPV T01RPV, TV8RPV, "F Core inlet / outlet temperature, ST08RPV saturation temperature 6 DVI-I WWTDVILI, Ibm /sec. Individual cornponents and total flow WWTIRWil, in DVI I WWTIRW13 7 DVI-2 WWTDVIL2, Ibm /sec. Individual components and total flow WWTIRWI2, in DVI-2 WWTIRW14 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT CLDP-509, CLDPSIO in. Level CL-CMT balance lines 10 IRWST 1RWST lbm Mass of fluid in IRWST I1 IRWST CLDP-701 in. Collapsed liquid level in IRWST 12 IRWST UIRWST Btu Fluid energy in IRWST 13 Pnmary sump AMPSMP lbm Pnmary sump fluid mass 14 Pnmary sump CLDP-901 in. Primary sump level 15 Pnmary sump UPSMP Btu Pnmary sump fluid energy 16 Secondary sump AMSSMP lbm Secondary sump fluid mass 17 Secondary sump CLDP-902 in. Secondary sump level 18 Secondary sump USSMP Btu Secondary sump fluid energy 19 Prunary sump WSTSMPET, WWTSMPIT Ibm /sec. Pnmary sump steam and liquid injection rate 20 Pnmary sump MISMPII, MISMPI2, Ibm Integrated primary sump and IRWST MISMPIT, MIIR%T flows 21 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG side inlet / outlet MSSriOP1, MSSGOP2 plena 22 Surge line PLM lbm Fluid mass in surge line 23 Surge line CLDP-602 in. Collapsed liquid level in surge line 24 Surge line UPSL Blu Fluid energy in surge line 25 RPV MWRPV lbm Total fluid mass in reactor vessel O mAap60&2344w-55.noo:Ib.100395 5.5.3-4 REVISION: 1 ~
n TABLE 5.5.3-1 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.5.3 LONG. TERM TRANSIENT Plot No. Component Variables Units Description 26 RPV DCM lbm Fluid mass in downcomer 27 RPV LDP01DC in. Collapsed liquid level in downcomer compared to various reference elevadons 28 RPV MWO3RPV lbm Fluid mass in core region 29 RPV LDP03RPV in. Collapsed liquid level in core 30 RPV RPVAVDF2 Core exit void fraction 31 RPV RPVAQOU2 Core exit quality 32 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 33 RPV MWO8RPV lbm Fluid mass in the upper head 34 RPV RPVASL2 in. Level of Tsat line 35 RPV RPVAPab2, RPVPWR kW Heated rod power above and below Tsat level and total 36 RPV RPVASOU2 lbm/sec. Core steam generadon rate 37 RPV RPVALIN2 lbm/sec. Calculated core flow (q; 38 RPV HTMXRPV, ST08RPV
*F Maximum clad temperature, saturation temperature and delta 39 Hot leg MWHL1, MWHL2 lbm Water mass in bot legs 40 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs 41 Cold leg CLlWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 42 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 43 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4, and break 44 ADS and break ADS 13TLR, ADS 41TLR, Ibm /sec. Liquid flow out ADS 1-3 and ADS-4 ADS 42TLR 45 ADS and break BRKSTLR Ibm /sec. Liquid flow and total flow out of break 46 Mass balance TOTMASS lbm Total system mass inventory 47 Mass balance PRIMMASS, PRIMASS2 lbm Measured primary system inventory and valve from mass balances 48 Mass balance MERROR lbm Mass balance error 49 Mass balance MIN, MOUT SRCMASS lbm Integrated mass flow in and out of primary system and source mass 50 Energy balance Various Btu Component of energy balance 51 ADS-4 ADS 4T h.R. ADS 42TLR lbm/sec. Oscilladons in ADS-4 liquid flo v p 52 Surge line CLDP-602 in. Oscilladons in surgeline level 53 RPV CI'F-107 psia Oscillations in upper bead pressure m:W344 -55.non:tb.ioo395 5.5.3-5 REVISION: 1
TABLE 5.5.3-1 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.5.3 LONG TERM TRANSIENT Plot No. Component Variables Units Description 54 RPV CLDP-ll3 in. Oscillations in upper plenum level 55 RPV LDP03RPV in. Ckcillations in core level 56 RPV LDP01DC in. Oscillations in downcomer level l l 9 maap6oac344w.55..on:ivioo395 5.5.3-6 REVISION: 1
i l l 4 / < l a l i; l i ! THE FIGURES LISTED IN TABLE 5 431 $ ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT
- l. l i
1 1 t
\
m:W344w.55.non:1b 101195 5.53-7 REVISION: 1
- . ~ . - - _ . . - . . . . . - . - . - - . - . . - . . - . - . - - - .. .- ~.. -.--..- _ . _ . - .
in S.6 Analysis of Matrix Test SB12 l-Matrix Test SB12 (OSU Test U0112) simulated a DEG DVI SBLOCA with LTC and without the operation of the nonsafety-related systems. Reactor-side break flow was piped to the break separator. All other connections to the break separator were isolated by using blind inserts in the piping source. I The CMT and accumulator-side break flow were aligned directly to the sump rather than a separator, l Before break initiation, the break separator and primary sump were isolated from the break sources by two break valves. The analysis of Matrix Test SB12 is divided into three subsections, as follows: i e Facility performance is discussed in Subsection 5.6.1. It provides a brief outline of the ] response of the test facility; further details are available in the Final Data Report.* The short-term transient for SB12 encompassed the start of the simulation up to [ l' 6 ' seconds. This period included blowdown, natural circulation. ADS, and initial IRWST stages ! i of the transient. j i ! The analysis of the long-term transient for SB12 encompassed the time frame from , i [ ]** seconds to the end of the test. This phase of the transient included IRWST injection I and covered the transition to sump injection. The long-term transient actually started at d IRWST injection, which is discussed as part of the short-term transient. Between the end of j the short-term transient and [ ]** seconds, the system remained relatively inactive. The discussion of the short- and long-term phase of the transient focuses on important thermal-hydraulic phenomena identified in the PIRT (Table 1.3-1). The mass and energy balance results are j key indicators of the quality of the analysis on which this discussion is based. These are discussed in I detail in Subsections 6.2.2 and 6.2.3. O v mAap600\2344w 56.non:1b.100395 5.6-1 REVISION: 1
% ?
,_ 5.6.1 Facility Performance - De performance of the OSU test facility during Matrix Test SB12 in reference to the five transient phases is outlined in the following:
. Blowdown
. Natural circulation
- ADS
. IRWST injection
. Sump injection .
De overall performance of the facility during the transient is shown in Figures 5.6.1-1 to 5.6.1-4. Figure 5.6.1-1 shows the pressurizer pressure throughout the test with various phases and operating . components delineated on the figure. The time scale was reduced for clarity since there were only small changes in system pressure during the long-term phase of the transient. Figure 5.6.1-2 shows the DVI-2 flow and its composition from the various sources at each time in the transient. Figure 5.6.1-3 shows the calculated core steam generation rate throughout the test, and Figure 5.6.1-4 shows the variation in average measured core outlet temperature and peak clad temperature relative to the core outlet saturation temperature. Figures 5.6.1 1 and 5.6.1-2 show that there was a continuous flow of cool water to the core from the 7
\
passive safety systems throughout the transient. Once initiated, the ADS depressurized the primary system, which enhanced the CMT-2 and ACC-2 injection flow rates. LDtimately, opening of the ADS-4 valves reduced the system pressure to start gravity-driven IRWST injection. De operation of the passive injection systems overlapped so that as one source of water drained, the next became operable to continue the cooling process. He level of steam generation in the core and the average measured core outlet fluid temperatures and maximum clad temperatures are shown in Figures 5.6.1-3 and 5.6.1-4. These figures show that there was sufficient cooling flow to prevent excessive core heating, and the core remained covered. De core remained subcooled for large periods of the transient and when steam production occurred, the rate of generation remained well below the rate at which water was delivered to the core. 5.6.1.1 Blowdown Phase The blowdown phase began at time zero when the break was initiated and continued until the primary system pressure was in equilibrium with the secondary-side pressure at about [ ]'6# seconds. Break flow from the CMT-1/ACC-1 side of the break started immediately when the break valves opened and was directed into the primary sump. Immediately following the opening of the break, the primary system pressure fell to the end of the blowdown phase. Both the liquid and steam flow from the break separator increased from the time of the break valve opening until about [ ]'6" seconds. I During this phase of the transient, cooling flow was provided from CMT-2, which remained in the recirculation mode, and heat was removed from the primary system via the SGs and break. He mMp60m2%-56.non:Ib-too395 5.6.1-1 REVISION: 1
l l l pressurizer and surge line completely drained at [ ]'6* M [ ]*6' mA msg @$ h We break flow created a rapid depressurization and level decrease in the RCS. 1 5.6.1.2 Natural Circulation Phase l In this LOCA simulation, the single- and two-phase natural circulation phase was marked by a slow reduction in system pressure, due to mass flow out the break. De liquid level in the downcomer fell below the elevation of the DVI line (break elevation), which resulted in the reduced break flow. After [ ]'6# seconds, the liquid and steam break flows decreased dramatically. During this phase of the transient, the SG tubes drained by about [ ]'6# seconds and at this time, heat removal from the primary system continued via the PRHR and the break. The steam in the SG tubes became superheated and remained so until the end of the transient. Due to the break, flow decreased in the RCS, the cold legs emptied, and the downcomer annulus liquid level was at the bottom of the DVI line. In response to voiding in CL-1, CMT-2 transitioned to draindown mode at [ ]'6# seconds. De falling CMT-1 level reached the ADS low-level setpoint at [ ]'6# seconds. The natural circulation phase of the transient continued to [ ]'6# seconds when the ADS-1 valve opened. 5.6.1.3 Automatic Depressurization System Phase ADS-1 actuation was followed by ADS-2 and ADS-3 actuation [ ]'6# and [ ]'6# seconds later, I respectively. After the initiation of ADS, ACC-2 injection began at [ ]'6" seconds. The influx of cold water, combined with increased venting via the break and ADS, led to a rapid depressurization of the primary system. The RCS inventory decreased and the minimum RPV inventory of [ ]'6' lbm was indicated at about [ l'** seconds. No temperature excursions were recorded as a result of the core reaching its minimum level. Actuation of ADS-2 caused a refill of the pressurizer as water and steam flowed out of the ADS. The pressurizer gradually drained by [ ]'6* seconds. [ ]'6# seconds after the CMT-1 low-level setpoint occurred, when the primary system pressure had reached the level of [ ]'6# psig, actuation l of ADS-4 at [ ]'6# seconds completed depressurization to a level that allowed IRWST injection at [ ]'6* seconds via DVI-2. IRWST injection continued for the remdader of the test. The rate of RCS depressurization changed very little as a result of ADS-4 actuation; the large break dominated the primary system depressurization. During the ACC-2 injection, increased resistance reduced flow out of CMT-2. As ACC-2 drained, l CMT-2 flow resumed. ACC-2 was completely drained at [ ]*6' seconds. The measured break flow significantly decreased due to the low liquid level in the downcomer, and the combined CMT-2 and ACC-2 injection was sufficient to start increasing the primary system inventory. mwp6coc3.ws6.noo:ib ioo395 5.6.1-2 REVISION: 1
5.6.1.4 In-Containment Refueling Water Storage Tank Injection - IRWST injection was the transition from the short- to long-term phase of the transient. De initial phase of IRWST injection involved an increase in flow through DVI-2 (until about [ J'6' seconds), which was followed by a gradual flow reduction after [ ]'6# seconds as the driving head between the IRWST and the RCS fell due to the reduced IRWST water level. The simultaneous occurrence of IRWST injection initiation and CMT-2 flow increase after ADS-4 actuation increased the inventory in the RCS and kept the bottom of the core subcooled. By the end of the test, about [ J percent of the core was subcooled. Steam generation rate in the core decreased throughout the SB12 short-term transient. The liquid levels in both the primary sump and the break separator continuously increased after about [ ] seconds, with the primary sump level about 5 in. higher than the level in the break separator. The primary sump overflowed into the secondary sump at [ ] seconds. When the water level in the primary sump increased above the level of the break separator penetration into the sump, the break separator loop seal flow reversed at about [ ] seconds. The break flow l decreased and reversed at about [ ]'6# seconds, when the water level in the break separator reached j the elevation of the break line penetration into the break separator. The flow through the ADS-4 i increased with the break flow reversal. 5.6.1.5 Sump Injection Injection from the primary sump via the check valves around the main sump injection valves began when the level in the IRWST was low enough to allow flow. This injection caused a reduction in the flow rate from the IRWST. When the IRWST level fell to [ ]'in., the main sump injection valves opened at [ j# seconds. The primary sump-2 injection flowed to the reactor about [ ]'6# seconds after the sump valves opened; however, reverse primary sump-l injection flow was observed when the sump valves opened. The primary sump-l injection line sustained backflow from the IRWST to the primary sump for the remainder of the test as the level in the tanks equalized. O b m:\ap60m2%.56.noo:tb.ioo395 5.6.1-3 REVISION: 1
i I l l TABLE 5.6.1 1 l OSU TEST ANALYSIS PLOT PACKAGE FOR SUBSECTION 5.6.1 Plot Component Variables Units Description No. l 1 Pressurizer Cirr-6(M psia System pressure and event history l 2 Water injection WWTDVII+WWTDV12, Ibm /sec. Total of CMT, accumulator, IRWST, l WOUTACCl+WOUTACC2, and sump injection flows WWTIRWII+WWTIRWI2 WWTSMPIT 3 Reactor vessel RPVASOU2 lbm/sec. Steam generation in reactor vessel 1 i 4 Reactor vessel T08RPV, HTMXRPV, TS AT 'F Reactor vessel outlet temperature, j maximum clad temperature and fuel l exit saturation temperature l l l l l 4 l l l l l l l l l l l 9 maar600c3m 56.non:ib.ioo395 5.6.1 -4 REVISION: 1
O - s THE FIGURES LISTED IN TABLE 5.6.1-1 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT O O mAap60m2ms6.noa:ib ioo395 5.6.1-5 REVISION: 1
. . - . . _ . . . - . - . . - . .~. - - - - -. - ._.. - . . _ . . - . - - - ~ . . . - . . - - - - - - . .
5.6.2 Short Term Transient - For the DEG break of a DVI line, Matrix Test SB12, the short-term transient encompassed the time frame up to [ - ]'A' seconds. As shown in Figure 5.6.2-1, this period included the full depressurization of the RPV through the break and all four stages of the ADS, together with CMT and accumulator injection plus the initial stages of IRWST injection. De variation in mass, energy, pressure, and temperature throughout this stage of the transient are illustrated in the plot package outlined in Table 5.6.2-1. He plots concentrate on the primary system,incluaing the accumulators, CMTs, IRWST, primary sump, and flow from the primary system via the ADS, break, and IRWST overflow. There were two principal parameters of interest for the short-term transient: Adequate flow must be maintained from the passive systems to the reactor vessel. 2 Adequate flow into the core must be maintained to ensure that decay heat was removed from the simulated fuel rods without a temperature excursion. Dese parameters are addressed in the following discussion. p 5.6.2.1 Maintenance of Core Cooling Mass Injected to the Primary System ; Figures 5.6.2-5 and 5.6.2-6 show the combined effect of the injection flow for the short-term phase of the transient. Separate plots of the individual contributions to the total flow can be located by consulting the plot package index given in Table 5.6.2-1. Note that the flow measurements for the outflow from CMTs were temporarily out-of-range for this test. i Figure 5.6.2-6 shows how the CMT-2, ACC-2, and IRWST supplied a continuous flow of cool water to the core. During the first [ ]'** seconds, cooling flow was provided by the CMT-2. De rate of flow from the CMT-2 increased after it trarsitioned to draindown mode at [ ]'A' seconds. ACC-2 injection initiated at [ ]'** seconds, caused a temporary reduction in CMT-2 flow, but led to an overall increase in flow to the core to a peak value of about [ ]'** lbm/sec. Following the end of accumulator injection, the CMT-2 again provided cooling flow in combination with the IRWST injection flow. Reactor Pressure Vessel and Downcomer Behavior The effect of the water flow on the average measured core inlet / outlet temperatures and heater rod temperatures during the short-term phase of the transient is shown in Figures 5.6.2-3 and 5.6.2-57. CMT-2 flow was not sufficient to keep the core subcooled; shortly after the break valves were opened mAap600cw56.noo:1b.ioo395 5.6.2- 1 REVISION: 1
(at about [ ] seconds), the decreasing saturation temperature reached the core outlet temperature. After this time, the core outlet temperature followed the saturation temperature until about [ ] seconds. When the collapsed level in the downcomer fell to the elevation of the DVI line at about [ ]'6' seconds, steam was most likely vented out through the break, even though none was detected by the break steam flow meter. At the same time, ADS-1 actuated, and shortly afterward, ACC-2 injection occurred. ADS-1 and ACC-2 resulted in rapid primary system depressurization. After [ J seconds, the influx of water from the IRWST and CMT-2 injection became sufficient to subcool the lower part of the core. Figure 5.6.2-57 shows that there were no significant excursions in heater rod temperatures throughout the short-term transient; therefore, sufficient core inventory and flow was maintained through this phase of the transient to remove the simulated decay heat generation. For significant portions of the transient, a two-phase mixture was present in the core and upper plenum regions, with core boiling kept at a low level. The following discussion tracks the variation in water level and mass throughout the RPV and downcomer. He mass and level for the core region are shown in Figures 5.6.2-44 and 5.6.2-45. The collapsed liquid level in the core indicates that the heater rods remained covered with a single- or two-phase mixture. The minimum core inventory of about [ ]'6# lbm occurred about [ ]'6' seconds into the transient. Figure 5.6.2-45 shows that the collapsed liquid level fell [ ]'6' in. below the top of the core during this phase of the transient. He average void fraction of the core two-phase mixture may be estimated by dividing the measured core collapsed liquid level by the [ ]'6' in. heated rod length. In this test, the minimum collapsed liquid level corresponded to a core void fraction of [ ] ."6# By j about [ J'6* seconds, the IRWST injection reduced core boiling (Figure 5.6.2-55) to a low level. l Re collapsed liquid level in the upper plenum region and the associated fluid mass are shown in , Figures 5.6.2-49 and 5.6.2-48. After the break, the region of the upper plenum spanned by the LDP l cell fully drained and remained so until IRWST injection supplied sufficient inventory to initiate a refill of the region after [ ]'6# seconds. The upper plenum collapsed level stayed below the hot-leg elevation by the end of the short-term transient. Figures 5.6.2-50 and 5.6.2-51 show that the upper head lost inventory later than the upper plenum. At about [ ]'6# seconds, the upper head was drained to the upper support plate elevation until approximately [ ] seconds. He reduction in upper plenum and upper head inventory contributed to the reduction in overali RPV inventory. The mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.6.2-41 and 5.6.2-42. The downcomer collapsed liquid level fell to the DVI line elevation at [ ]'6' seconds and to the top of the heater rods at [ ]'6* seconds. The minimum downcomer inventory of [ ]'6# lbm was indicated at [ ]*6# seconds when the collapsed liquid level dropped 13-in. below the top of the heater rods. ACC-2 injection and IRWST inflow partly restored water inventory in the downcomer, and the collapsed liquid level increased to the elevation of the DVI line at about [ ] seconds. ma,6coc344w 56.noo:it,100395 3,6,22 REVISION: I
l l (3 g 5.6.2.3 Energy Transport from the Primary System l Following the break, energy was deposited in the primary system fluid by the heater rods to simulate decay heat and the primary system metal as it cooled down. Some fluid energy was lost to ambient and out the break. Energy must be removed from the primary system to prevent excessive fluid and heater rod temperature excursions. De AP600 is designed to remove heat by a combination of the ! SGs, PRilR, and the ADS. l Steam Generator and Passive Residual Heat Removal Heat Transfer During normal operation, most of the primary system heat was removed via the SGs; however, once the coolant pumps tripped, the reduced system flow decreased primary-to-secondary-side heat transfer. The SGs were only available as heat sinks until the time when the primary system pressure dropped to that of the secondary side; afterwerd, the secondary side became a potendal heat source for the primary side. The PRHR is designed to remove heat from the primary system via natural circulation once the safety signal opens the PRHR isolation valve. The PRHR continue to remove energy after the SGs are thermally isolated until ADS actuates, creating a path for the removal of energy from the i primary system. Figures 5.6.2-33 and 5.6.2-4 show the SG primary and secondary-side pressure together with the PRHR integrated heat transfer, as represented by the IRWST fluid energy after allowing for the O) ( contribution from ADS 1-3 inflow. He SGs were a potential sink for primary system heat while the SG primary-side pressure was above that of the secondary side, that is, before [ ]"# seconds. PRIIR heat removal began [ ]'6* seconds into the test. The PRHR was responsible for all the IRWST heatup until ADS-1 actuation. The PRHR heat transfer reduced significantly after [ ]"# seconds. During the active phase, the PRHR transferred heat to the IRWST at an average rate of approximately
-[ ]*' Btu /sec., as calculated from the IRWST fluid energy increase.
Energy Transport via the Break and Automatic Depressurization System De mass flow rate from the primary system via the break is shown in Figures 5.6.2-67 and 5.6.2-68. As shown in these figures, liquid flow was detected by the flow measuring devices for the short-term transient. The liquid flow reached its maximum flow rate of [ ]"# lbm/sec. at about [ ]"' seconds and then rapidly decreased. When ADS-1 actuated, the primary system depressurized to around [ ]"' psia (Figure 5.6.2-1). After the initiation of ADS 1, liquid flow through the break ceased at about [ ]** seconds (most likely replaced by steam flow) due to the fluid level in the downcomer dropping below the elevation of the DVI line. The smaller liquid flow rate (about [ ]"# Ibm /sec.) resumed through the break at about [ ]"' seconds and continued until the end of the short-term phase. ADS 13 rapidly depres-surized the system and at [ ]"# seconds, ADS-4 initiated and the primary system continued to depressurize to BAMS header pressure. O N) mwwxum 56.nonnico395 5.6.2-3 REVISION: 1
By the end of the short-term transient ([ )*** seconds), there was flow out-of the two ADS-4 valves (most likely two-phase flow) and through the break (Figures 5.6.2-64,5.6.2-63). Integrated mass flow from the primary system via the ADS and the break is shown in Figure 5.6.2-62, and the corresponding integrated energy flow is shown in Figure 5.6.2-69. The total system inventory plot given in Figure 5.6.2-70 indicates that about [ ]'** lbm ofinventory was gained during the short-term transient. The energy balance is shown in Figure 5.6.2-74. O O l i m:\ap6000344w 56.non:Ib-100395 5.6.2-4 REVISION: 1 1 i
s.) TABLE 5.6.21
~
OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.6.2 Plot No. Component Variables Units Description 1 Pressurizer CI'T-6N psia System pressure 2 RPV RPVPWR kW Core power 3 RPV TOIRPV, 71)8RPV, *F Core inlet / outlet temperature, ST08RPV saturation temperature 4 Steam generator CPT-201, CI"T-2N, psia Pnmary and secondary pressures in SG CI"T-301, CIYT-302 5 DVI-l WWIDVIL1, Ibm /sec. Individual components and total flow in WWTIRWII, DV!-l WOlJTACCI, WWTIRWI3 6 DVI-2 WWTDVIL2, Ibm /sec. Individual components and total flow in WWTIRWI2, DVI-2 WOUTACC2, WWTIRWI4 7 CMT AMCMTIB, Ibm Fluid mass in CMTs (excludes balance AMCMT2B lines) 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT MIWDVIL1, Ibm Integrated mass out of CMTs p MIWDVIL2 V 10 CMT WWTDVILI, Ibm /sec. Flow out of CMTs WWTDVIL2 11 CMT WOUTCLBI, Ibm /sec. Flow into CMTs WOUTCLB2 12 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 13 CMT UCMT1, UCMT2 Btu Fluid energy in CMTs 14 IRWST IRWST Ibm Mass of fluid in IRWST 15 IRWST CLDP-701 in. Collapsed liquid level in IRWST 16 IRWST WWTIRWII, Ibm /sec. Flow from IRWST to DVI lines W%TIRWI2 17 IRWST IRWSTOR Ibm /sec. Overflow from IRWST to sump 18 IRWST ADS 13TMR Ibm /sec. Total ADS flow into IRWST 19 IRWST ADS 13TIR, MIIRWil, Ibm Integrated mass out of IRWST MIIRW12, MIIRWlO 20 IRWST UIRWST Btu Fluid energy in IRWST 21 PRHR CLDP-802 in. Collapsed liquid level in PRHR HX %./ mamp600s2mw-56.non:Is100395 5.6.2-5 REVISION: 1
l l TABLE 5.6.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.6.2 Plot No. Component Variables Urdts Description 22 PRHR WWOTPRHR lbm/sec. Measured outlet flow from PRHR tube 23 Accumulator AMACCl, AMACC2 lbm Mass of fluid in accumulators 24 Accumulator CLDP-401, CLDP-402 in. Collapsed liquid level in accumulators 25 Accumulator WOUTACCl, Ibm /sec. How from accumulators WOIJTACC2 26 Accumulator MOLTTACCl, Ibm Integrated mass out of accumulators MOUTACC2 27 Accumulator UACCl, UACC2 Btu Huid energy in accumulators 28 Pnmary sump AMPSMP lbm Pnmary sump fluid mass 29 Primary sump CLDP-901 in. Primary sump level 30 Pnmary sump UPSMP Btu Pnmary sump fluid energy 31 Steam generator MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG primary side MSSGOP1, MSSGOP2 inlet / outlet plena 32 Steam generator MSSGHTI, MSSGHT2, Ibm Mass of fluid in SG primary side bot MSSGCTI, MSSGCT2 and cold tubes 33 Steam Cf"T-201, CPT-301, psia & SGI pressure and PRHR integrated heat generator /PRHR QPRHRI Btu output 34 Pressurizer PZM lbm Fluid mass in pressurizer l 35 Pressurizer CLDP-601 in. Collapsed liquid level in pressurizer 36 Pressurizer UPZ Btu Fluid energy in pressurizer 37 Surge line PLM lbm Fluid mass in surge line 38 Surge line CLDP-602 in. Collapsed liquid level in surge line 39 Surge line UPSL Btu Fluid energy in surge line 40 RPV MWRPV lbm Tr,tal fluid mass in reactor vessel 41 RPV DCM lbm Fluid mass in downcomer 42 RPV LDP01DC in. Collapsed liquid level in downcomer compared to various reference elevations 43 RPV MWOIRPV lbm Fluid mass in lower plenum 44 RPV MWO3RPV lbm Fluid mass in core region 45 RPV LDP03RPV in. Collapsed liquid level in core 46 RPV RPVXVFO Core exit void fraction 47 RPV RPVXRQO Core exit quality 48 RPV MWO6RPV lbm Fluid mass in the upper plenum 49 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 50 RPV MWO8RPV lbm Fluid mass in the upper head 51 RPV LDP08RPV in. Collapsed liquid level in the upper head m:\ap6000344w 56.aoo:lb-100395 5.6.2-6 REVISION: 1
l V TABLE 5.6.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.6.2 Plot No. Component Variables Units Description 52 RPV URPV Btu Total fluid energy in reactor vessel 53 RPV RPVXE, RPVASL2 in. Level of Tsat line 54 RPV RPVPab, RPVAPbl, kW Heater rod power above and below Tsat RPVPWR level and total 55 RPV RPVRXV,RPVASOU2 lbm/sec. Core steam generation rate 56 RPV RPVALIN2 lbm/sec. Calculated core flow 57 RPV HTMXRPV, ST08RPV 'F Maximum clad temperature and saturation temperature 58 Hot leg MWHL1, MWHL2 lbm Water mass in bot legs 59 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs 60 Cold leg CLIWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 61 Cold leg CLIVMS, CL2VMS, ihm Vapor mass in cold legs CL3VMS, CL4VMS 62 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4s, and break 63 ADS and break BRKTIVF, ADl3TIVF, Ibm Totalintegrated vapor flow for ADS p AD41TIVF, AD42TIVF and break 64 ADS and break BRKTILF, AD13TILF, Ibm Total integrated liquid flow for ADS AD41TILF, AD42TILF and break 65 ADS and break ADS 13SVR, Ibm /sec, Vapor flow out ADS 13 and ADS-4 ADS 41SVR, ADS 42SVR 66 ADS and break ADS 13SLR, Ibm /sec. Liquid flow out ADS 1-3 and ADS 4 ADS 41SLR, ADS 42SLR 67 ADS and break BRKSSVR Ibm /sec. Vapor flow out of break 68 ADS and break BRKSSLR lbm/sec. Liquid flow out of break 69 ADS and break BRKSPEI, ADS 13El, Btu Integrated fluid energy for ADS 1-3, ADS 41El, ADS 42EI ADS-4, and break 70 Mass balance TOTMASS lbm Total system mass inventory 71 Mass balance PRIMMASS, Ibm Measured primary system inventory and PRIMASS2 value from mass balance 72 Mass balance MERROR lbm Mass balance error 73 Mass balance MIN, MOUT lbm Integrated mass flow in and out of SRCMASS primary system and source mass 74 Energy balance Various Btu Components of energy balance O maapm2344w 56.non:It>100395 5.6.2-7 REVISION: 1
O TIIE FIGURES LISTED IN TABLE 5.6.21 ARE NOT INCLUDED IN Tills NONPROPRIETARY DOCUnfENT 1 l l 9i r O m:W344w.56. con:1bl00395 5.6.2-8 REVISION: 1
. . . - - - _ . -. ~ . ---- . -.. . -. .
A ( 5.6.3 Long-Term Transient s The long-term transient started after initiation of IRWST injection, covered the transition from IRWST to sump injection, and provided information on the LTC response of the AP600 plant. For the DEG DVI break, Matrix Test SB12, the long-term transient analyzed extends from [ ]**# seconds to the end of the test around [ ]'6' seconds. The behavior of the test facility during this period of the transient is discussed in this subsection using the plot package detailed in Table 5.6.3-1. This analysis concentrates on the components of the primary system that remained active during the LTC phase, that is, the RPV, the hot legs, ADS-4, the sumps, and the IRWST. During the long-term transient, the main thermal-hydraulic phenomena of interest were the maintenance of core cooling and the removal of energy from the primary system. 5.6.3.1 Maintenance of Core Cooling Mass injected into Primary System Total DVI line flow, CMT flow, and IRWST flows are shown in Figures 5.6.3-6 and 5.6.3 7, and the flow from the primary sump is shown in Figure 5.6.3-19. Throughout the jong-term transient, there was no contribution to the DVI flow from the CMTs as the CMTs were drained. p 1 i h During the IRWST injection phase of the transient, IRWST-2 flow proceeded at a gradually declining l rate. At [ ]'6# seconds, flow from primary sump-2 began through the check valves, around the ! main injection valves, reducing IRWST-2 flow. The IRWST flow rate was maintained through DVI-2 decreasing from [ ]'6' lbm at [ ]'6' seconds to [ ]'6* lbm/sec, at the end of the transient. At j [ ]'6' seconds, the primary sump injection valves opened and primary sump injection flow (about ! [ ]'6' lbm/sec, started at [ ]'6' seconds, slightly decreasing the IRWST flow in line-2. The net ) result was that an injection flow rate of [ ]'6' lbm/sec. was maintained through DVI-2. Reactor Pressure Vessel and Downcomer Response The effect of the water inflow on the average measured downcomer fluid temperatures, core inlet, and core outlet temperatures, and heater rod temperatures during the long-term phase of the transient is shown in Figures 5.6.3-4,5.6.3-5, and 5.6.3-38. Figure 5.6.3-4 shows that there was an increase in average downcomer fluid temperatures during the long-term transient. By the end of the test, this average temperature was about [ ]'6' *F below the saturation temperature for the primary system. Figure 5.6.3-5 shor ' hat the core remained near saturation throughout the long-term transient. As l shown in Figures 5.6.3-34 to 5.6.3-36, the DVI line flow method described in Section 4.11 indicates that a small level of boiling was maintained throughout the long-term transient. Nevertheless, the level of boiling was small and showed that the inflow from IRWST-2 was sufficient to cool the core effectively, map 600s2344 56.noa:ib ioo395 5.6.3-1 REVISION: 1
Figure 5.6.3 38 shows that there were no significant excursions in heater rod temperatures throughout the long-term transient; therefore, sufficient core inventory and flow were maintained through this phase of the transient to remove decay heat. The following discussion tracks the variation in water level and mass throughout the RPV and downcomer. The mass and level for the core region are shown in Figures 5.6.3-28 and 5.6.3-29. The collapsed liquid level in the core indicated that the heater rods were always covered with a single- or two-phase mixture. During the long-term transient, the collapsed liquid level remained below the top of the core until it reached the top of the core at the end of the transient. The impact of hot water injected from IRWST-2 as well as reversed break flow on the system temperatures is shown in Figures 5.6.3-4 and 5.6.3-5 as a small increase in fluid temperature in the downcomer and at the core inlet. The hot water also led to an increase in the calculated steam generation rate as shown in Figure 5.6.3-36 and a corresponding fall in the level at which the core reached saturation temperature (Figure 5.6.3-34). The collapsed liquid level in the upper plenum region is shown in Figure 5.6.3-32. This figure indicates that the collapsed liquid level increased during this period, from the elevation of the DVI line at [ J seconds; after the break flow was reversed at about [ ]"" seconds, the collapsed liquid level gradually increased above the hot-leg elevation. The mass of water in the RPV is shown in Figure 5.6.3-25. After an initial decline, the RPV water mass settled at an average value of [ ] lbm until the break flow was reversed, after which time the RPV water inventory gradually increased to [ ] lbm ([ ] percent of the initial RPV liquid mass) by the end of the transient. The mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.6.3-26 and 5.6.3-27. The collapsed liquid level remained at the elevation of the DVI line until about [ ]'6# seconds, when due to the reversed break flow, the level raised above the hot-leg elevation. The sump injection had no noticeable effect on the levels in the RPV and downcomer. 5.6.3.2 Energy Transport from the Primary System During the long-term transient, energy continued to be deposited in the primary system from the heater rods and metal. The SGs remained inactive throughout this phase of the transient. The small PRIIR flow continued until about [ J seconds. The principal path for energy out of the primary system was via the ADS-4 valves. Integrated mass flow from the primary system via the ADS and the break is shown in Figure 5.6.3-43. During the LTC phase of the transient, the only significant energy outflow was through the ADS-4 valves. When the break flow was reversed after [ ] seconds, the outflow through ADS-4 increased to maintain the primary system inventory almost constant. This is confirmed by Figures 5.6.3-44 and 5.6.3-45, which show the flow through the ADS and the break. Throughout the mup600c3m.56.non:ib-too395 5.6.3-2 REVISION: 1
LTC, significant inflow was provided by IRWST-2 inje, tion. During the sump injection phase of the transient (about [ J'A'second duration), outflow through the ADS-4 valves continued. Figure 5.6.3 36 shows the calculated steam generation rate, as determined by the DVI line flow method discussed in Section 4.11. During this phase of the transient, steam generation decreased. , i Figure 5.6.3-50 shows all the components to the system energy balance. The calculated steam generation rate, mass flow, and energy balance are affected by the break flow after it was reversed. O l O i l m:W344w 56. con:Ib-100395 5.6.3-3 REVISION: 1 I l
TABLE 5.63-1 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.63 i LONG TERM TRANSIENT l Plot No. Component Variables Units Description 1 RPV RPVPWR kW Core power 2 Primary sump TSMPII, TSMP12 'F Sump injection line temperatures 3 DVI TDVIL1, TDVIL2 'F DVI line temperatures 4 RPV T01DC, T02DC, T03DC, 'F Water and saturation temperatures in ST01DC downcomer 5 RPV TOIRPV, 'II)8RPV, *F Core inlet / outlet temperature, ST08RPV saturation temperature 6 DVI-l WW'IDVIL1, Ibm /sec. Individual components and total flow WWTIRWII, in DVl-1 WWTIRWU 7 DVI-2 WWTDVIL2, Ibm /sec. Individual components and total flow WWTIRWI2, in DVI-2 WWTIRWI4 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 10 IRWST IRWST Ibm Mass of fluid in IRWST 11 IRWST CLDP-701 in. Collapsed liquid level in IRWST 12 IRWST UIRWST Btu Fluid energy in IRWST i3 Prunary sump AMPSMP lbm Pnmary sump fluid mass 14 Pnmary sump CLDP-901 in. Prunary sump level 15 Pnmary sump UPSMP Btu Prunary sump fluid energy 16 Secondary sump AMSSMP lbm Secondary sump fluid mass 17 Secondary sump CLDP-902 in. Secondary sump level 18 Secondary sump USSMP Btu Secondary sump fluid energy 19 Pnmary sump WSTSMPET, WWTSMPIT Ibm /sec. Pnmary sump steam and liquid injection rate 20 Pnmary sump MISMPII, MISMPI2, Ibm integrated primary sump and IRWST MISMPIT, MIIRWT flows 21 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG side inlet / outlet MSSGOPI, MSSGOP2 plena 22 Surge line PLM lbm Fluid mass in surge line 23 Surge line CLDP-602 in. Collapsed liquid level in surge line 24 Surge line UPSL Btu Fluid energy in surge line 25 RPV MWRPV lbm Total fluid mass in reactor vessel 9 m:\ap600(2344w-56.non:Ib-100395 5.6.3-4 REVISION: 1
TABLE 5.6.31 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.6.3 l LONG-TERM TRANSIENT ' Plot No. Component Variables Units Ikscription 26 RPV DCM lbm Fluid mass in downcomer 27 RPV LDP0lDC in. Collapsed liquid level in downcomer I compared to various reference elevations 28 RPV MWO3RPV lbm Fluid mass in core region 29 RPV LDP03RPV in. Collapsed liquid level in core 30 RPV RPVRXVFO Core exit void fraction 31 RPV RPVRXQO Core exit quality ! 32 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 33 RPV MWO8RPV lbm Fluid mass in the upper head 34 RPV RPVASL in. Level of Tsat line l 35 RPV RPVAPab, RPVPWR kW Heater rod power above and below Tsat level and total 36 RPV RPVASOUT lbm/sec. Core steam generation rate 37 RPV RPVTMRI lbm/sec. Calculated core flow ; 38 RPV HTMXRPV, 'F Maximum clad temperature, saturation ! [ S'II)8RPV temperature and delta l 39 Hot leg MWHL1, MWHL2 lbm Water mass in hot legs - 40 Hot leg MVHL1, MVHL2 lbm Vapor mass in hot legs 41 Cold leg CLlWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 42 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 43 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4, and break 44 ADS and break ADS 13TLR, ADS 41TLR, Ibm /sec. Liquid flow out ADS 13 and ADS-4 ADS 42TLR 45 ADS and break BRKSTLR Ibm /sec. Liquid flow and total flow out of break 46 Mass balance TOTMASS lbm Total system mass inventory 47 Mass balance PRIMMASS, PRIMASS2 lbm Measured primary system inventory and valve from mass balances 48 Mass balance MERROR lbm Mass balance error 49 Mass balance MIN, MOUT SRCMASS lbm Integrated mass flow in and out of primary system and source mass l 50 Energy balance Various Btu Component of energy balance maap60m23m-56.non:6 loo 395 5.6.3-5 REVISION: 1
O TIIE FIGURES LISTED IN TABLE 5.6.31 ARE NOT INCLUDED IN TIIIS NONPROPRIETARY DOCUMENT O O maaraxn3w56.aoo:isioo395 5.6.3-6 REVISION: 1
l l l 1 5.7 Analysis of Matrix Test SB13 ! l Matrix Test SB13 (OSU Test U0ll3) simulated a 2-in, break of DVI-1 SBLOCA with LTC and i without the operation of the nonsafety-related systems. To simulate a 2-in break of DVI-1, piping was installed from the DVI line to the break separator. De piping from the break separator connected to the DVI line via two break valves. One break valve isolated the break nozzle and separateir inlet from the common injection line of CMT-1 and ACC-1. De other break valve isolated the common break nozzle and the inlet to the break separator from the portion of the DVI line, which connected to the reactor. De analysis of Matrix Test SB13 is divided into three sections as follows: Facility performance is discussed in Subsection 5.7.1. It provides a brief outline of the response of the test facility; further details are available in the Final Data Report.m
=
ne short-term transient for SB13 encompassed the start of the simulation up to [ ]** seconds. His period included blowdown, ADS, and initial IRWST stages of the transient. The analysis of the long-term transient for SB13 encompassed the time frame from
]** seconds to the end of the test. This phase of the transient included IRWST O' [
injection and covered the transition to sump injection. The long-term transient actually started at IRWST injection, which is discussed as part of the short-term transient. Between the end of 1 the short-term transient and [ ]'6* seconds, the system remained relatively inactive with the l exception of the CMT and pressurizer refill. At [ ]** seconds, CMT-2 began to refill and l CMT-1 followed [ ]** seconds later. The CMT refill phenomena is discussed further in Subsection 6.1.1 and the discussion of the long-term transient provided here begins at I [ ]** seconds. l l The discussion of the short- and long-term phase of the transient focuses on important thermal-hydraulic phenomena identified in the PIRT (Table 1.31). The mass and energy balance results are key indicators of the quality of the analysis on which this discussion is based. These are discussed in detail in Subsections 6.2.2 and 6.2.3. l O 1 l m:\ap600\2344w-57.non:lb-100395 5.7 1 REVISION: 1
.~- . - .
i l 5.7.1 Facility Performance - The performance of the OSU test facility during Matrix Test SB13 in reference to the four transient phases is outlined in the following:
= Blowdown
- ADS
- IRWST injection
. Sump injection The overall performance of the facility during the transient is shown in Figures 5.7.1-1 to 5.7.1-4.
Figure 5.7.1-1 shows the pressurizer pressure throughout the test with various phases and operating components delineated on the figure. The time scale was reduced for clarity since there were only small changes in system pressure during the long-term phase of the transient. Figure 5.7.1-2 shows the total DVI line flow and its composition from the various sources at each time in the transient. Figure 5.7.1-3 shows the calculated core steam generation rate throughout the test, and Figure 5.7.1-4 shows the variation in average measured core outlet temperature and heater rod temperature relative to the core outlet saturation temperature. Figures 5.7.1-1 and 5.7.1-2 show that there was a continuous flow of water to the core from the ; p passive safety-related systems throughout the transient. Once initiated, the ADS lines rapidly l depressurized the primary system, which enhanced the CMT and accumulator injection flow rates. ) Ultimately, the opening of ADS-4 valves sufficiently reduced the system pressure to start gravity-driven IRWST injection. Operation of the passive injection systems overlapped so that as one source of water drained, the next became operable to continue the cooling process. The level of steam generation in the core and the response of the average measured core outlet fluid temperatures and maximum clad temperatures are shown in Figures 5.7.1-3 and 5.7.1-4. These figures show that there , was sufficient cooling flow to prevent excessive core heating, and the core remained covered. The core remained subcooled for large periods of the transient and when steam production occurred, the rate of generation remained well below the rate at which water was delivered to the core. 5.7.1.1 Blowdown Phase The blowdown phase began at time zero when the break was initiated and continued until the primary system pressure was in equilibrium with the secondary-side pressure at arourd [ ]'*" seconds. Immediately following the opening of the break, the primary system pressure fell gradually until about [ ]'6' seconds when core power started to reduce. During this phase of the transient, cooling flow was provided from CMT-2, which remained in the recirculation mode, and heat was removed from the primary system via the SGs. The pressurizer and surge line completely drained at [ ]'6*and [ ]" seconds, respectively. O m:wo2m -57.non:it,-ioo395 5.7.1-1 REVISION: 1
1 I When core power started to reduce at about [ ]'6# seconds and CMT-2 transitioned to the injection mode of operation (at [ ]'6# seconds) the primary system pressure nearly stabilized at about [ ]'6* psia until the end of the blowdown phase. During this time, the SG tubes drained by about [ ]'6# seconds and at this time, heat removal from the primary system continued via the PRiiR. l The steam in the SG tubes became superheated and remained so until the end of the transient. , 1 In response to voiding in CL-3, CMT-1 transitioned to draindown mode at [ ]'** seconds, and the falling CMT level reached the ADS low-level setpoint at [ ]'6# seconds. The blowdown phase of the transient continued to [ ]'6* seconds when the primary- and secondary-side pressures equalized and the ADS-1 valve opened. 5.7.1.2 Natural Circulation Phase here is no natural recirculation phase in Test SB13 since the primary and secondary-side pressures equalized at the time ADS-1 initiated. l 5.7.1.3 Automatic Depressurization System Phase l ADS-1 actuation was followed by ADS-2 and ADS-3 actuation [ ]'6# and [ ]# seconds later, . respectively. Accumulator injection began shortly after initiation of the ADS. The influx of cold I water combined with increased venting via the ADS led to a rapid depressurization of the primary i system. Actuation of ADS-4 at [ ]'6# seconds completed depressurization to a level that allowed IRWST injection at [ ]'*# seconds via DVI-1 and [ l'6* seconds via DVI-2. During accumulator injection, increased flow path resistance reduced flow out of the CMTs. As the accumulators drained, CMT flow resumed. The accumulators were fully drained at [ ]'6# seconds, before IRWST commenced. CMT-1 and CMT-2 fully drained at [ ]' 6
- and [ ],'** respectively.
The minimum RPV inventory of [ ]'*# lbm was observed at [ ]'*# seconds, shortly after ADS-3 actuation. The transfer from CMT-2 to IRWST-2 injection was indicated by a low RPV inventory of about [ j'6# lbm before IRWST-2 injection started. Actuation of ADS-1 rapidly refilled the pressurizer as water and steam flowed out of the ADS. The pressurizer gradually drained by [ ]'6# seconds. 5.7.1.4 In Containment Refueling Water Storage Tank Injection IRWST injection was the transition from the short- to long-term phase of the transient. The initial phase of IRWST-2 injection involved an increase in flow through DVI-2, which was followed by a gradual flow reduction as the driving head between the IRWST and the RCS fell due to the reduced IRWST water level. IRWST-2 injection flow started at [ ]'*# seconds, about at the same time the CMT-2 emptied. Between [ ]'*# and [ ]'*# seconds, the only cooling available to the core was via IRWST-2 injection. m%wxt.2344.-57.nco:ib-too395 5.7.1-2 REVIs!ON: 1
l O V When maximum IRWST-2 flow was established, the influx of water from the IRWST was sufficient to
]'6# seconds. The steam generation rate increased after keep the core subcooled from [
[ J'** seconds from minimum steam generation rate of about [ ]'*# lbm/sec. with a decreasing j IRWST-2 injection rate and stabilized at [ ]'"# lbm/sec. after [ ] seconds. After [ ]'** l seconds, temperature at the top of the core followed the saturation temperature for primary system pressure. After reflooding, the level of both CMTs stayed essentially constant for several thousand seconds. As the level in the IRWST decreased, the backpressure in each DVI line decreased and was low enough at approximately [ ]'*# seconds for CMT-1 to start injecting to the break. CMT-2 started injecting to I the vessel at about [ ]# seconds. The injection flow from the CMTs was oscillating between approximately [ ]'** seconds. At about [ ]'6# seconds flow from the break separator to the primary sump was reversed as indicated by flow meter FMM-905. Shortly after that the break flow was reversed as indicated by both decreasing integral break flow (see Figure 5.7.3-43) I and break liquid flow (Figure 5.7.3-45). I i i 5.7.1.5 Sump Injection l . Injection from the primary sump via the check valves around the main sump injection valves began at ! about [ ]'** seconds when the level in the IRWST was low enough to allow flow. 'Ihis caused a
]
, V] [ reduction in the flow rate from the IRWST-2. When the IRWST level fell to [ ]'6# in., the main sump injection valves opened and the sump injection flow rate increased. This increase occurred at i I [ ]'** seconds and the driving head from the sump was sufficient for flow from the sump to the ; IRWST on the DVI line. ! ., 1 i J m:\ap600\2344w 57.non:1b-100395 5.7.1 3 REVISION: 1
TABLE 5.7.1 1 OSU TEST ANALYSIS PLOT PACKAGE FOR SUBSECTION 5.7.1 Plot No. Component Variables Units Description i Pressurizer CPT-604 psia System pressure and event history 2 Water WWTDVil+WWTDV12, Ibm /sec. Total of CMT, accumulator, IRWST, injection WOUTACCl+WOUTACC2, and sump injection flows WWTIRWI1+WWTIRWI2, WWTSMPIT 3 Reactor RPVASOU2 lbm/sec. Steam generation in reactor vessel vessel 4 Reactor T08RPV, HTMXRPV, TS AT 'F Reactor vessel outlet temperature, vessel maximum clad temperature and fuel exit saturation temperature O O m:upm2344.-5tnoo:ib-1%395 5.7.1 4 REVISION: 1
. . _ . . . _ . . . _ . _ . _ . . . _ - __. .~.____ _ . _ ___. _ __ _ .- ___..._ _ - -
i 7 i THE FIGURES LISTED IN TABLE 5.7.1 1 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT - 1 ' l 1 l 1 i I 4 I l 1 -1 1. I J 4 j 4 J I m:\ap60m2344w-57.aos:Ib.100995 5.7.1-5 REVISION: 1
l 5.7.2 Short-Term Transient - For the 2-in. DVI line break, Matrix Test SB13, the short-term transient encompassed the time frame up to [ ]'6' seconds. As shown in Figure 5.7.1-1, this period included the full depressurization of l the facility through all four stages of the ADS, together with CMT and accumulator injection plus die l initial stages of IRWST injection. Variations in mass, energy, pressure, and temperature throughout this stage of the transient are illustrated in the plot package outlined in Table 5.7.21. The plots l concentrate on the primary system, including the accumulators, CMTs, IRWST, primary sump and i flow from the primary system via the ADS, break, and IRWST overflow.
]
There were two principal parameters of interest for the short-term transient:
- Adequate flow from the passive systems to the reactor vessel must be maintained.
Adequate flow into the core must be maintained to ensure that decay heat was removed from the simulated fuel rods, without a temperature excursion. These parameters are addressed in the following discussion. 5.7.2.1 Maintenance of Core Cooling d Mass Injected to the Primary System Figures 5.7.2-5 and 5.7.2-6 show the combined effect of the injection flows for the short-term phase of the transient. Separate plots of the individual contributions to the total flow can be located by consulting the plot package index given in Table 5.7.2-1. Figure 5.7.2-6 shows how the CMT-2, ACC-2 and IRWST 2 supply a continuous flow of cooling
- water to the core. During the first [ J# seconds, cooling flow was provided by CMT-2. The rate of flow from CMT-2 increased from zero to [ ]'6* lbm/sec. at [ ]'6' seconds after ADS-1 actuated and ACC-2 injection started. ACC-2 injection resulted in a decrease in the CMT-2 flow but led to an overall increase in flow to the core to a peak value of [ ]'6'lbm/sec. Following the end of accumulator injection, CMT-2 again provided cooling flow until it drained. The only period in which there was relatively little cooling flow was at about [ ]'6* seconds when the CMT-2 drained and IRWST injection started.
Reactor Pressure Vessel and Downcomer Behavior The effect of water flow on the average measured core inlet / outlet temperatures and heater rod temperatures during the shon-term phase of the transient is shown in Figures 5.7.2-3 and 5.7.2-57. The combined CMT-2 and ACC-2 flow was sufficient to keep the bottom of the core subcooled from L./ mvm2344.-57..on:ib-too395 5.7.2- 1 REVISION: 1
about [ l seconds. he core outlet temperature remained at the saturation level throughout the short-term transient until the influx of water from the IRWST-2 was sufficient to subcool the core. Figure 5.7.2-57 shows that there were no significant excursions in heater rod temperatures throughout the short-term transient; therefore, sufficient core inventory and flow was maintained through this phase of the transient to remove the simulated decay heat. For significant portions of the transient, a two-phan mixture was present in the core and upper plenum regions, with core boiling kept at a low level. He following discussion tracks the variation in water level and mass throughout the reactor vessel and downcomer. 1 De mass and level within the core region are shown in Figures 5.7.2-44 and 5.7.2-45. The collapsed liquid level in the core indicates that the heater rods remained covered with a single- or two-phase mixture. The minimum core inventory of [ ]lbm occurred at [ ]'6* seconds into the transient after ADS-3 actuated. Figure 5.7.2-45 shows that the collapsed liquid level dropped [ ]'6' in. below I the top of the heater rods during this phase of the transient. The average void fraction of the core I two-phase mixture may be estimated by dividing the measured core collapsed liquid level by the [ ]"' in. heated rod length. In this test, the minimum collapsed liquid level corresponded to a core I void fraction of [ ].**' After this time, core inventory increased due to injection from ACC-2 and CMT-2 and decreasing flow leaving the primary system through ADS 1-3 and the break. By the end of the short-term transient, IRWST-2 injection reduced core boiling (Figure 5.7.2-55) and the core was nearly water-solid. The collapsed liquid level in the upper plenum region and the associated fluid mass are shown in Figures 5.7.2-49 and 5.7.2-48. De collapsed liquid level in the upper plenum span of LDP-113 decreased immediately after the break and fell to a minimum level (a few inches) between ADS-4 actuation and IRWST-2 injection (before the end of the accumulator injection). Then, IRWST-2 injection supplied sufficient inventory to initiate a refill. He upper plenum collapsed liquid level i increased up to the elevation of the cold leg by the end of the short-term transient. l l Figures 5.7.2-50 and 5.7.2-51 show that the upper head also lost inventory after the break valves j opened and was almost drained by [ ]'6' seconds. The upper head was refilled (collapsed liquid l level at about 6 in.) at [ ] seconds when the pressurizer started to drain after refill and the IRWST-2 injection flow was close to maximum. The reduction in upper plenum and upper head inventory was responsible for the reduction in overall RPV Inventory until the IRWST injection started (Figure 5.7.2-40). The mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.7.2-41 and , 5.7.2-42. The downcomer remained water-solid until about [ ] seconds, when the collapsed l liquid level fell to the elevation of the cold legs, reaching the lowest level at about [ ]'6' seconds (5 in. below the elevation of the top of heater rods). Then, the collapsed level increased almost to the elevation of the cold legs by the end of the short-term transient. mAap6000344w-57.non:lb-100395 5.7.2-2 REVISION: 1
im 5.7.2.2 Energy Transport from the Primary System Following the break, the heater rods deposited energy in the primary system fluid to simulate decay heat and the primary system metal as it cooled down. Some fluid energy was lost to the ambient and out of the break. Energy must be removed from the primary system to prevent excessive fluid and heater rod temperature excursions. In the AP600 plant, heat removal is designed to be achieved by a combination of the SGs and the PRHR plus the ADS. Steam Generator and Passive Residual Heat Removal Heat Transfer During normal operation, most of the primary system heat was removed via the SGs; however, once the coolant pumps tripped, reduced system flow caused a reduction in primary-to secondary-side heat transfer. De SGs were only available as heat sinks until the primary system pressure dropped to that of the secondary side, the two sides were then in thermal equilibrium. After that, the secondary side became a potential heat source for the primary side. The PRHR is designed to remove heat from the primary system once the signal opens the isolation valve. The PRHR will continue to remove enerFy after the SGs are thermally isolated until ADS actuates. Once the ADS is actuated, ADS 1-3 becomes the predominant path for the removal of energy from the primary system. 3 (b Figure 5.7.2-33 shows the SG primary- and secondary-side pressurc (together with the PRHR integrated heat transfer as represented by the IRWST fluid energy after allowing for the contribution from ADS 1-3 inflow). PRHR heat removal began [ ]'6# seconds into the test and the PRHR was responsible for all the IRWST heat-up until ADS-1 activation; the PRHR heat transfer then was reduced. The PRHR outlet flow ceased at about [ l'6# seconds. During the active phase, the PRHR transferred heat to the IRWST at an average rate of [ ]'6# Btu /sec. Energy Transport via the Break and Automatic Depressurization System The mass flow rate from the primary system via the break is shown in Figures 5.7.2-67 and 5.7.2-68. The short-term transient liquid flow was calculated. During the first [ ]'6# seconds following the break, [ ]'6# lbm of fluid flowed out of the primary system via the break at an average rate of approximately [ ]'6# lbm/sec., consisting of liquid ([ ]'6# lbm) and steam ([ ]'6* lW Durig this period, the primary system depressurized to around [ ]'6# psia (Figure 5.2.21). With the initiation of ADS 1-3 and accumulator injection, steam flow through the break ceased and was replaced by steam and liquid flow through the ADS 1-3 valves. Between [ ]'6# and [ ]' 6# seconds, ADS 1-3 caused the system to rapidly depressurize to about [ ]'6# psia and at [ ]' 6
- g seconds, ADS-t was initiated and the primary system continued to depressurize to BAMS header Q pressure.
in:w60m344w-57.non:ib loo 395 5.7.2-3 REVIslON: I
The initiation of the ADS reduced the flow through the break. During ADS 1-3 depressurization, steam and liquid flow through the ADS 1-3 valves occurred at a rate of [ ]'** lbm/sec. ([ {
]'6# seconds). Flow through the ADS 1-3 continued at a declining rate until about [ ]'6" l seconds when the flow through ADS 1-3 terminated and was replaced by flow through the lower
)
resistance ADS-4 paths. By the end of the short-term transient ([ ]'** seconds), water was ficwing l out of the two ADS-4 valves at approximately [ ]'** lbm/sec. (Figure 5.7.2-64). Integrated mass flow from the primary system via the ADS and the break is shown in Figure 5.7.2-62, l and the corresponding integrated energy flow is shown in Figure 5.'i_ '. 'Ihe total system inventory plot given in Figure 5.7.2-70 indicates that up to [ ]'A'lbm ofinventory left the system during the short-term transient. Components of the energy balance are shown in Figure 5.7.2-74. O l l 1 l l l 9 m:\ap600\2344w 57.non:lb-100395 5.7.2-4 REVISION: 1
D TABLE 5.7.21 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.7.2 Plot No. Component Variables Units Description 1 Pressurizer CPT-6N psia System pressure 2 RPV RPVPWR kW Core power 3 RPV TOIRPV, 'IT)8RPV, 'F Core inlet / outlet temperature, ST08RPV saturation temperature 4 SG CPT 201, CPT-2N, psia Primary and secondary pressures in SG CPT-301, CPT-302 5 DVI l WWTDVIL1, Ibm /sec. Individual components and total flow in WWTIRWil, DV11 WOUTACCl, WWTIRW13 6 DVI-2 WWTDVIL2, Ibm /sec. Individual components and total flow in WWTIRW12, DVI-2 WOUTACC2, WWTIRWI4 7 CMT AMCMTIB, Ibm Fluid mass in CMTs (excludes balance AMCMT2B lines) 8 CMT CLDP 502, CLDP 507 in. Collapsed liquid level in CMTs 9 CMT MlWDVIL1, Ibm Integrated mass out of CMTs [~N MlWDVIL2 10 CMT WWTDVIL1, Ibm /sec. Flow out of CMTs WWTDVIL2 11 CMT WOUTCLBI, Ibm /sec. Flow into CMTs > WOUTCLB2 12 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 13 CMT UCMT1, UCMT2 Btu Fluid energy in CMTs 14 IRWST IRWST lbm Mass of fluid in IRWST 15 IRWST CLDP-701 in. Collapsed liquid level in IRWST 16 IRWST WWTIRWil, Ibm /sec. Flow from IRWST to DVI lines WWTIRWI2 17 1RWST IRWSTOR lbm/sec. Overflow from IRWST to sump 18 IRWST ADS 13TMR lbm/sec. Total ADS flow into IRWST 19 IRWST ADS 13TIR, MIIRWII, Ibm Integrated mass out of IRWST MIIRW12, MllRWIO 20 IRWST UIRWST Btu Fluid energy in IRWST 21 PRHR CLDP-802 in. Collapsed liquid level in PRHR llX O mAsp600G344w 57.noo:lb-100395 5.7.2-5 REVISION: 1
1 l l 1 TABLE 5.7.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.7.2 Plot No. Component Variables Units Description l 22 PRHR WWOTPRHR Ibm /sec. Measured outlet flow from PRIIR tube 23 Accumulator AMACCl, AMACC2 lbm Mass of fluid in accumulators 24 Accumulator CLDP-401, CLDP-402 in. Collapsed liquid level in accumulators 25 Accumulator WOUTACCl, Ibm /sec. Flow from accumulators WOUTACC2 26 Accumulator MOUTACCl, Ibm Integrated mass out of accumulato:s MOUTACC2 27 Accumulator UACCl, UACC2 Btu Fluid energy in accumulators 28 Pnmary sump AMPSMP lbm Primary sump fluid mass 29 Pnmary sump CLDP-901 in. Pnmary sump level 30 Pnmary sump UPSMP Btu Pnmary sump fluid energy 31 SG MSSGIPI, MSSGIP2, Ibm Mass of fluid in SG primary side MSSGOP1, MSSGOP2 inlet / outlet plena 32 SG MSSGHTI, MSSGHT2, Ibm Mass of fluid in SG primary side bot MSSGCrl, MSSGCT2 and cold tubes 33 SG/PRHR Cirr-201, CI'F-301, psia & SGI pressure and PRHR integrated heat QPRHRI Btu output 34 Pressunzer PZM lbm Fluid mass in pressurizer 35 Pressurizer CLDP-601 in. Collapsed liquid level in pressurizer 36 Pressurizer UPZ Btu Fluid energy in pressurizer 37 Surge line PLM lbm Fluid mass in surge line 38 Surge line CLDP-602 in. Collapsed liquid level in surge line 39 Surge line UPSL Btu Fluid energy in surge line 40 RPV MWRPV lbm Total fluid mass in reactor vessel 41 RPV DCM lbm Fluid mass in dowmcomer 42 RPV LDP01DC in. Collapsed liquid levelin downcomer compared to various reference elevations 43 RPV MW0lRPV lbm Fluid mass in lower plenum 44 RPV MWO3RPV lbm Fluid mass in core region 45 RPV LDP03RPV in. Collapsed liquid level in core 46 RPV RPVAVDF2 Core exit void fraction 47 RPV RPVAQOU2 Core exit quality 48 RPV MWO6RPV lbm Fluid mass in the upper plemun 49 RPV LDP06RPV in. Collapsed liquid level in the upper plenmn 50 RPV MWO8RPV lbm Fluid mass in the upper bead 51 RPV LDP08RPV in. Collapsed liquid level in the upper head m:\ap60m2344w.57.non:Ib-100395 5.7.2-6 REVISION: 1
.. .-. - =- - ._. . . .. - _ _ _ - _ _ -
1 I [~ '\ TABLE 5.7.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.7.2 Plot No. Component Variables Units Description 52 RPV URPV Btu Total fluid energy in reactor vessel , 53 RPV RPVXE, RPVASL2 in. Level of Tsat line 54 RPV RPVPab, RPVAPab2, kW Heater rod power above and below Tsat l RPVPWR level and total ) 55 RPV RPVRXV, RPVASOU2 lbm/sec. Core steam generation rate 56 RPV RPVALIN2 lbm/sec. Calculated core flow 57 RPV HTMXRPV, S71)8RPV 'F Maximum clad temperature and saturation temperature 58 Hot leg MWHL1, MWHL2 lbm Water mass in bot legs l 59 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs 60 Cold leg CLlWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 61 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 62 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4s, and break 63 ADS and break BRKTIVF, AD13TIVF, Ibm Totalintegrated valmr flow for ADS Og AIMITIVF, AD42TIVF and break i 64 ADS and break BRKTILF, AD13TILF, Ibm Totalintegrated liquid flow for ADS l
- AIMITILF, AIM 2TILF and break I 65 ADS and break ADS 13SVR, Ibm /sec. Vapor flow out ADS 1-3 and ADS-4 )
ADS 41SVR, l ADS 42SVR I 66 ADS and break ADS 13SLR, Ibm /sec. Liquid flow out ADS l-3 and ADS-4 ADS 41SLR, ADS 42SLR 67 ADS and break BRKSSVR lbm/sec. Vapor flow out of break 68 ADS and break BRKSSLR lbm/sec. Liquid flow out of break 69 ADS and break BRKSPEl, ADS 13El, Btu Integrated fluid energy for ADS 1-3, ADS 41EI, ADS 42EI ADS-4, and break 70 Mass balance TOTMASS lbm Total system mass inventory 71 Mass balance PRIMMASS, Ibm Measured primary system inventory and PRIMASS2 value from mass balance 72 Mass balance MERROR lbm Mass balance error 73 Mass balance MIN, MOUT lbm Integrated mass flow in and out of SRCMASS primary system and source mass 74 Energy balance Various Btu Components of energy balance U,O maap600\2344w 57.non:1b-100395 5.7.2-7 REVISION: 1
O THE FIGURES LISTED IN TABLE 5.7.2-1 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT O O mAap600\2344w 57.noo:Ib100995 5.7.2-8 REVISION: 1
l V.O 5.7.3 LongeTerm Transient The long-term transient staned with initiation ofIRWST injection, covered the transition from IRWST to sump injection, and provided information on the LTC response of the AP600 plant with sump injection. For the 2-in. DVI line break, Matrix Test SB13, the long-term transient phase began at [ ]** seconds awl continued to the end of the test at about [ ]'6 seconds. The behavior of the test facility during this period of the transient is discussed in this subsection using the plot package detailed in Table 5.7.3-1. This analysis concentrates on the components of the primary system that remained active during the LTC phase, that is, the RPV, the hot legs, ADS-4, the sumps, and the IRWST. During the long-term transient, the main thermal-hydraulic phenomena of interest were maintenance of core coolios and removal of energy from the primary system. 5.7.3.1 Maintenance of Core Cooling Mass Injected into Primary System Total DVI line flow, CMT flow, and IRWST flows are shown in Figures 5.7.3-6 and 5.7.3-7, and the flow from the primary sump is shown in Figure 5.7.3-19. From around [ ]'6# seconds, there was a contribution to the DVI flow from CMT-2 as CMT-2 reached post-refill draindown. Prior to sump injection, during the IRWST injection phase of the transient, IRWST flow proceeded at a gradually declining rate. At [ ]'6# seconds, flow from the primary sump began through the check valves around the main injection valves, resulting in further reduction in IRWST-2 flow. From [ ]'** seconds to the end of the transient, a near-steady flow rate of [ ]** lbm/sec. was maintained through DVI-2. At [ ]'6# seconds, the main sump injection valves opened, resulting in a reversal of flow through IRWST injection line-1 and an increase in IRWST flow in line-2. The net result was that the injection flow rate in DVI-2 gradually decreased from [ ]'6" lbm/sec. at [ ]'6# seconds to [ ]** lbm/sec. at the end of the transient. Reactor Pressure Vessel and Downcomer Response The effect of water inflow on the average measured downcomer fluid temperatures, core inlet and core outlet temperatures, and heater rod temperatures during the long-term phase of the transient is shown in Figures 5.7.3-4,5.7.3-5, and 5.7.3-38. Figure 5.7.3-4 shows that there is a general increase in average downcomer fluid temperatures during the long-term transient. By the end of the test, some temperature stratification was observed in the downcomer. The temperature in the bottom of the downcomer was [ ]6' *F below saturation. Figure 5.7.3-5 shows that the core remained saturated for the entire long-term transient. Figures 5.7.3-34 to 5.7.3-36, show that the DVI line flow method discussed in Section 4.11 indicates that a small level of boiling was maintained until the end of the mAap600\2344w-57.non:1b-100395 REVISION: 1 5.7.3-1
transient. Nevertheless, the level of boi!ing was small and showed that the inflow from the IRWST and sumps was sufficient to subcool the RPV. Figure 5.7.3-38 shows that there were no significant excursions in heater rod temperatures throughout the bng-term transient; therefore, sufficient core inventory and flow was maintained throughout this phase of the transient to remove the decay heat generated. He following discussion tracks the variation in water level and mass throughout the reactor vessel and downcomer. He mass and level for the core region are shown in Figures 5.7.3 28 and $.7.3-29. The collapsed liquid level in the core indicated that the heater rods were always covered with a single or two-phase mixture. During the later stages of the transient, the collapsed liquid level remained just below the top l of the heater rods, and the core void fraction was [ ].** The fallin core inventory was a result of l the influx of hot water from the primary sump as it flowed through the check valves. The impact of this hot water en the system temperatures is shown in Figures 5.7.3-4 and 5.7.3-5 as a sudden increase in fluid temperature in the downcomer and at the core inlet. The hot water also led to an increase in the calculated steam generation rate, as shown in Figure 5.7.3-36, and a corresponding fall in the level at which the core reached saturation temperature (Figure 5.7.3-34). The collapsed liquid level in the upper plenum region is shown in Figure 5.7.3 32. This figure indicates that during the period before sump injection began, the collapsed liquid level initially fell and then remained at the elevation of the hot legs. Following the influx of hot water from the sumps, the level dropped to the elevation of the top of the DVI injection lines where it remained for the rest of the transient. This level corresponds to a void fraction of [ ]** in the upper plenum. The mass of water in the RPV is shown in Figure 5.7.3-25. After an initial decline, the reactor vessel water mass settled at an average value of [ ]** lbm until sump injection started when it gradually fell to [ ]** lbm, which is [ ]** percent of the initial vessel water inventory, and remained at this level to the end of the transient. Oscillations in vessel inventory occurred as shown in Figures 5.7.3-51 through 5.7.3-56. Dese oscillations and possible mechanisms for their production are discussed in Subsection 6.1.3. The mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.7.3-26 and 5.7.3-27. The collapsed liquid level fell below the elevation of the cold legs after CMT-2 injection at about [ ]** seconds for the entire long-term transient. The sump injection had little effect on water level in the downcomer. 5.7.3.2 Energy Transport from the Primary System During the long-term transient, energy continued to be deposited in the primary system from the heater rods, metal, and fluid flowing from the primary sump. The SGs and PRHR remained inactive throughout this phase of the transient and the principal path for energy out of the primary system was via the ADS-4 valves. maap600(2344w.57.non:lb-100395 5.7.3-2 REVISION: 1
-. ._ -. =__ _. . _ . . - . . . - . . - . - - . - _ - - _ _ - . - - - _ - . _ _ _ _ _ integrated mass flow from of the primary system via the ADS and the break is shown in Figures 5.7.3-43. During the L'IC phase of the transient, the only significant outflow is through the ADS-4 valves. This is confirmed by Figures 5.7.3-44 and 5.7.3-45, which show flow through the ADS and the break. During the sump injection phase of the transient, outDow was indicated as liquid going through the ADS-4 valves. Water flowed through each of the valves at an average rate of [ ]"# lbm/sec. Figure 5.7.3-36 shows the calculated steam generation rate as determined by the DVI line flow method. During the sump injection phase of the transient, steam was generated at about [ ]"# lbm/sec., indicate little or no flow from the steam vortex meters out of the ADS 4 valves. Steam left the primary circuit by this route as shown by de following. Figure 5.7.3-46 shows total measured system fluid inventory. During this phase of the transient after the start of primary sump injection (from [ ]"# seconds, that is, when core steam generation was most sigolficant, the total system inventory fell by over [ ]"# lbm. This amount corresponds to a steam flow rate of [ ]"# lbm/sec., which would not have been detected by the vortex meters.
. Examination of the fluid thermocouples on the outlet of the ADS-4 valves indicates that temperatures remained at or above saturation temperature following the start of sump injection.
Furthermore, as discussed in Subsection 6.1.3, it was not possible for all the steam generated in the core to flow from the upper head to the downcomer via the bypass holes. It can therefore be concluded that steam was leaving the primary system via ADS 4. Figure 5.7.3-50 shows all the components to the system energy balance. Further discussion of steam loss from the primary circuit is provided in the mass and energy balance discussions of Section 6.2. f% C m:W344w 57.non:lb.100995 5.7.3-3 REVIs10N: 1
TABLE 5.7J-1 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.7.3 LONG-TERM TRANSIENT Plot No. Component Variables Units Description
! RPV RPVPWR kW Core power 2 Pnmary sump TSMPII, TSMPI2 'F Sump injection line temperatures 3 DVI TDVIL1, TDVIL2 'F DVIline temperatures 4 RPV T01DC, 'IT)2DC, T03DC, 'F Water and saturation temperatures in ST01DC downcomer 5 RPV TOIRPV, 'll)8RPV, "F Core inlet / outlet temperature, STT)8RPV saturation temperature 6 DVI-l WWTDVIL1, Ibm /sec. Individual components and total flow WWTIRWII, in DVI-l WWTIRWI3 7 DVI-2 WWIDVIL2, Ibm /sec. Individual components and total flow WWTIRWI2, in DVI-2 WWTIRWI4 8 CMT CLDP402, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 10 IRWST IRWST Ibm Mass of fluid in IRWST I1 IRWST CLDP-701 in. Collapsed liquid level in IRWST 12 IRWST UIRWST Btu Fluid energy in IRWST 13 Pnmary sump AMP 3MP lbm Pnmary sump fluid mass 14 Pnmary stunp CLDP-901 in. Pnmary sump level 15 Pnmary sump UPSMP Btu Pnmary sump fluid energy 16 Secondary sump AMSSMP lbm Secondary sump fluid mass 17 Secondary sump CLDP-902 in. Secondary sump level
{ 18 Secondary sump USSMP Btu Secondary sump fluid energy j 19 Pnmary sump WSTSMPET, WWTSMPIT Ibm /sec. Pnmary sump steam and liquid I injection rate ) 20 Pnmary sump MISMPII, MISMP12, Ibm Integrated primary sump and IRWST MISMPIT, MIIRWT flows 21 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG side inlet / outlet MSSGOP1, MSSGOP2 plena 22 Surge line PLM lbm Fluid mass in surge line 23 Surge line CLDP-602 in. Collapsed liquid level in surge line 24 Surge line UPSL Btu Fluid energy in surge line 25 RPV MWRPV lbm Total fluid mass in reactor vessel O ninap600\2344w-57.non:1b-100395 5.7.3-4 REVISION: 1
I l
~N (V TABLE 5.7.3-1 (Continued)
OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.7.3 LONG-TERM TRANSIENT Plot No. Component Varbbles Units Descriptien 26 RPV DCM. Ibm Fluid mass in downcomer 27 RPV LDP01DC in. Cellapsed liquid level in downcomer compared to various reference elevations 28 RPV MWO3RPV lbm Fluid mass in core region 29 RPV LDP03RPV in. Collapsed liquid level in core 30 RPV RPVAVDF2 Core exit void fraction 31 RPV RPVAQOU2 Core exit quality 32 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 33 RPV MWO8RPV lbm Fluid mass in the upper head 34 RPV RPVASL2 in. Level of Tsat line 35 RPV RPVAPab2, RPVPWR kW Heater rod power above and below Tsat level and total 36 RPV RPVASOU2 lbm/sec. Core steam generation rate 37 RPV RPVALIN2 lbm/sec. Calculated core flow 38 RPV HTMXRPV, 'F Maximum clad temperature, saturation t ST08RPV temperature and delta ( 39 Hot leg MWHL1, MWHL2 lbm Water mass in hot legs 40 Hot leg MVHL1, MVHL2 lbm Vapor mass in hot legs 41 Cold leg CLIWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 42 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 43 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS 4, and break 44 ADS and break ADS 13TLR, ADS 41TLR, Ibm /sec. Liquid flow out ADS 13 vnd ADS-4 ADS 42TLR 45 ADS and break BRKSTLR Ibm /sec. Liquid flow and total flow out of break 46 Mass balance TOTMASS lbm Total system mass inventory 47 Mass balance PRIMMASS, PRIMASS2 lbm Measured pnmary system inventory and valve from mass balances 48 Mass balance MERROR lbm Mass balance error 49 Mass balance MIN MOUT SRCMASS lbm Integrated mass flow in and out of primary system and source mass 50 Energy balance Various Btu Component of energy balance 51 ADS-4 ADS 41TLR, ADS 42TLR Ibm /sec. Oscillations in ADS-4 liquid flow p 52 Surge line CLDP-602 in. Oscillations in surgeline level V 53 RPV Cirr-107 psia Oscillations in upper head pressure maap600c344w-57.non:1b-Ioo395 5.7.3-5 REVISION: I ) 1
TABLE 5.7.3-1 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.7.3 LONG TERM TRANSIFET Plot No. Component Variables Units Description 54 RPV CLDP-ll3 in. Oscillations in upper plenum level 55 RPV LDP03RPV in. Oscillations in core level 56 RPV LDP01DC in. Oscillations in downcomer level O m:%taA2344w-57.noo:1bl00395 5.7.3-6 REVISION: 1
... ~ ..- .. . . _ .-....-.. .. . - -
.. - . - - . . . . . ~ . . . . . - . . . . . . . ~ . .- -- - - . . . - - . - - . - . .
I i-l THE FIGURES LISTED IN TABLE 5.7.31 ' ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT O O m:W344w-57.non:lb-100995 5.7.3-7 REVISION: 1
l l l 5.8 Analysis of Matrix Test SB14 - I Matrix Test SB14 (OSU Test U0014) simulated a no-break LOCA event with inadvertent actuation of l the ADS and progressing to LTC, without operation of the nonsafety-related systems. In this event, i ADS-1 opened to initiate the transient, while other safety-related systems activated in response to the - passive safeguards actuation signals received. The simulated single failure was one of the ADS-4 lines. 1 The analysis of Matrix Test SB14 is divided into three sections as follows:
- Facility performance is given in Subsection 5.8.1 and it provides a brief outline of the response of the test facility. Further details are available in the Final Data Report.*
- The short-term transient for SB14 encompassed the start of the simulation up to I
[ ]** seconds. Dis period includes the initial blowdown through the ADS and initial IRWST injection stages of the transient.
- The long-term transient for SB14 encompassed the time frame from [ ]** seconds to the end of the test. His phase of the transient included the IRWST injection and covered the transition to sump injection. The long-term transient actually started at IRWST injection, which is discussed as part of the short-term transient. Between the end of the short-term transient and [ ]** seconds, the system remained relatively inactive with the exception of ;
the CMT refill. At [ ]'6* seconds and [ ]'6* seconds, respectively, CMT-1 and CMT-2 l began to refill. The CMT-2 level instrument LDP-502 was unavailable during this test because the channel was used to initiate the transient. ADS actuation occurs when either CMT reaches the low-level setpoint. The operator induced a low-level signal from CMT-2 via LDP-502 to initiate SB14; the sum of CMT-2 level instruments LDP-504, LDP-506, and LDP-508 was therefore utilized to determine the CMT-2 level. Since CMT refill phenomena are discussed further in Section 6.1.1, the discussion of the long-term transient begins at [ ]'6* seconds. The discussion of the short- and long-term phases of the transient focuses on important thermal-hydraulic phenomena identified in the PIRT (Table 1.3-1). The mass and energy balance results are key indicators of the quality of the analysis on which this discussion is based. These are discussed in detail in Subsections 6.2.2 and 6.2.3. O mW3m.58.non:1b 100395 3,8.] REVISION: 1
O g 5.8.1 Facility Performance - The performance of the OSU test facility during Matrix Test SB14 is outlined in reference to the five transient phases as follows:
- Blowdown ,
- Natural circulation
- ADS
= IRWST injection
- Sump injection The overall performance of the facility during the transient is shown in Figures 5.8.1-1 to 5.8.1-4.
Figure 5.8.1-1 shows the pressurizer pressure throughout the test with various phases and operating components. The time scale was reduced for clarity since there were only small changes in system pressure during the long-term phase of the transient. Figure 5.8.1-2 shows the total DVI line flow and its composition from the various sources at each time in the transient. Figure 5.8.1-3 shows the calculated core steam generation rate throughout the test, and Figure 5.8.1-4 shows the variation in average measured core outlet temperature and peak clad temperature relative to the core outlet saturation temperature. Figures 5.8.1-1 and 5.8.1-2 show that there was a continuous flow of cooling water to the core from the passive safety-related systems throughout the transient. Once initiated, the ADS rapidly depressurized the primary system, and thereby enhanced CMT and accumulator injection flow rates. Ultimately, the ADS-4 vent path sufficiently reduced the system pres.sure to allow gravity-driven IRWST injection. The passive injection systems overlapped and as one source of water drained, the next became available to continue the cooling process. The level of steam generation in the core and the response of the average measured core outlet fluid temperatures and maximum clad - temperatures are shown in Figures 5.8.1-3 and 5.8.1-4. These figures show that the cooling flow was sufficient to prevent excessive core heating, and the core remained covered. The core remained subcooled for large periods of the transient and when steam production did occur, the rate of generation remained below the rate at which water was delivered to the core. 5.8.1.1 Blowdown Phase This event did not exhibit a blowdown phase as the break cases did; it began at time zero when ADS-1 was initiated, as though from a spurious actuation signal. The primary system pressure reached equilibrium with the secondary-side pressure at about [ ]'** seconds. Immediately after the ADS-1 valves opened, primary system pressure decreased as a vent path was established. During this phase, heat was removed from the primary system via the SGs until pressure equilibrium was reached with the secondary side. The pressurizer and surge line received flow from the primary loop to feed the open ADS valves throughout the initial phase of this event. I
\
m:\ap6000344w-58.non:lh-100395 5.8.1-1 REVISION: 1
l 5.8.1.2 Natural Circulation Phase - In this simulation, the single- and two-phase naturst circulation phase was insignificant since there was a large reduction in system pressure due to open AUS flow paths, compared to the more stable j pressure observed in SB01. During this phase of the transient, the SG tubes all drained around l [ ]'*' seconds. At that time, heat removal from the primary system continued via the ADS outflow and to a lesser extent, the PRHR. Steam in the SG tubes became superheated and remained so until l the end of the transient. In response to voiding in CL-3, CMT-1 transitioned to draindown mode l before [ ]'6* seconds elapsed. 5.8.1.3 Automatic Depressurization System Phase Actuation of ADS-1 was followed by ADS-2 and ADS-3 actuation [ ]'6
- and [ ]"' seconds later, respectively. Accumulator injection began after all paths of ADS 1-3 were open. The influx of cold water combined with venting via the ADS produced rapid depressurization of the primary system.
Actuation of ADS-4 at about [ ] seconds completed depressurization and resulted in IRWST injection at about [ ]'** seconds via DVI-1 and DVI-2. During accumulator injection, increased I flow path resistance reduced flow out of the CMTs. CMT flow resumed again as the accumulators drained. The accumulators had fully drained before [ ]'A' seconds of the transient clapsed, and the j CMT-1 and CMT-2 levels showed that the tanks had drained completely at about [ ]'6# seconds, l shortly after IRWST injection began. Minimum RPV mass inventory of [ ]'6# lbm occurred at l about [ ]'6# seconds into the transient. Actuation of ADS-1 caused a rapid refill of the pressurizer as water and steam flowed out of the ADS. The pressurizer gradually drained by [ ]*** seconds. 5.8.1.4 In Containment Refueling Water Storage Tank Injection The start of IRWST injection is the transition from the short- to long-term phase of the transient. The initial phase of IRWST injection involved an increase in flow through the two DVI lines, which was followed by a gradual reduction in flow as the driving head between the IRWST and the RCS fell with the reducing IRWST water level. Once maximum flow was established, the influx of water from the IRWST was sufficient to keep the core subcooled from [ ]'6# seconds to the end of the short-term transient. No RPV pressure / level oscillations were observed in the SB14 transient. 5.8.1.5 Sump Injection Injection from the primary sump via the check valves around the main sump injection valves began at [ ]'6# seconds when the level in the IRWST was low enough to allow that flow. This caused a reduction in the flow rate from the IRWST. When the IRWST level fell to [ ]'6' A h nn sump injection valves opened and the sump injection flow rate increased at [ ]'6' seconds via DVl-1. DVI-2 flow from the IRWST decreased to zero when the main sump valves opened; ma p600s2w58.noo:n-loo 395 5.8.1-2 REVISION: 1
7....__..._ i 1 1 1 l I l i thereafter, DVI-2 flow from the IRWST increased to approximately [ ] lbm/sec., the ! [ ]'6" second value.- This flow corresponded to the flow into the IRWST in DVI 1, which indicated that any additional decrease in IRWST inventory was small. I i a ) i i y a l i i l 1 I i 4 l 1 l I l i l 1 I I l l l m:W344w 58.non:1b-100395 5.8.1-3 REVISION: 1 I
TABLE 5.8.1 1 OSU TEST ANALYSIS PLOT PACKAGE FOR SUBSECTION 5.8.1 Plot No. Component Variables Units Description 1 Pressurizer CI'T-6N psia System pressure and event history 2 Water WWIDVll+WWIDV12, Ibm /sec. Total of CMT, accumulator, IRWST, injection WOUTACC1+WOUTACC2, and sump injection flows WWTIRWI1+WWTIRWI2, WWTSMPIT 3 Reactor RPVASOUT Ibm /sec. Steam generation in reactor vessel vessel 4 Reactor T08RPV, HTMXRPV, TS AT 'F Reactor vessel outlet temperature, vessel maximum clad temperature and fuel exit saturation temperature O O mhsmA2344w 58.noa:Istoo395 5.8.14 REVISION: 1
. - - - - - ~ - - - - . . . - . ~ . . . . ~ . , ~ . . . . . . . - . - - - . . --
. . - . . - . - - . - . . . . . . - - - . ~ . - - - .
4
!l.
~
l 1 1 THE FIGURES LISTED IN TABLE 5.8.1 1 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT l l l l 1 l l 1
)
i l l l I I
)
m:h, N344w-58. nan:Ib 100995 5.8.1-5 REVISION: 1 !
/
3 (V 5.8.2 Short-Term Transient For the inadvertent ADS actuation event, Matrix Test SB14, the short-term transient encompassed the time frame up to [ ]'6* seconds. As shown in Figure 5.8.1 1, this period included full depressurization of the facility through all four stages of the ADS, together with CMT and accumulator injection plus the initial stages of IRWST injection. The variation in mass, energy, pressure and temperature throughout this stage of the transient are illustrated in the plot package outlined in Table 5.8.2-1. The plots concentrate on the primary system, including the accumulators, CMTs, IRWST, the primary sump, and flows from the primary system via the ADS, break, and IRWST overflow. For the short-term transient there were two principal parameters to be examined: 1 l Adequate flow must be maintained from the passive systems to the reactor vessel. 1 Adequate flow into the core must be maintained to ensure decay heat removal from the i simulated fuel rods, without a temperature excursion. l l These parameters are addressed in the following discussion. l l ( 5.8.2.1 Maintenance of Core Cooling I l Mass injected into the Primary System l Figures 5.8.2-5 and 5.8.2-6 show the combined injection flows for the short-term phase of the transient. Separate plots of the individual contributions to the total flow are presented in the plot package index given in Table 5.82-1. It should be noted that for this test, the CMT-2 values for level, mass, and fluid energy are obtained by summing the three narrow range level instrument readings because the wide range instrument LDP-502 had been used to initiate the event. Figures 5.8.2-5 and 5.8.2-6 show how the CMTs, accumulators and IRWST supply a continuous flow of cool water to the core. During the first [ ]'** seconds, cooling flow was supplied by the CMTs. The rate of flow from the CMTs gradually increased to a steady value slightly above [ ]'6* lbm/sec. during the initial [ ]'#* seconds, then it fell in response to the start of accumulator injection, which temporarily shut off CMT flow. Following the end of accumulator injection, the CMTs again supplied cooling flow until they become fully drained. The only period in which there was minimal cooling flow was during [ ]'** seconds at the conclusion of CMT draining and the start of IRWST injection. Reactor Pressure Vessel and Downcomer Behavior q The effect of water flow on the average measured core inlet / outlet temperatures and peak clad Q temperatures during the short-term phase of the transient is shown in Figures 5.8.2-3 and 5.8.2-57. maap60ccmss.non:n.too395 5.8.2- 1 REVISION: 1
The core outlet temperature remained lower in the short-term transient than in the SBLOCA simulations because much of the initial hot RPV mass inventory was swept out of the open ADS valves quickly and replaced by cold passive safety system liquid. With no safety injection water lost to the break location, all passive safety-related system water was available to subcool the vessel. Figure 5.8.2-57 shows that there were no significant excursions in heated rod temperatures throughout the short-term transient; therefore, sufficient core inventory and flow were maintained throughout this phase of the transient to remove the decay heat generated. For significant portions of the transient, a two-phase mixture was present in the core and upper plenum regions due to the high rate of RPV depressurization; however, the amount of core boiling was kept at a low level. The following discussion explains the variation in water level and mass throughout the reactor vessel and downcomer. The mass and level for the core region are shown in Figures 5.8.2-44 and 5.8.245. The collapsed liquid level in the core indicated that the core voided rapidiy when the ADS 1-3 valve opening sequence was in progress. With this two-phase mixture, the minimum core inventory of [ ]** lbm occurred at [ ]** seconds into the transient, before accumulator injection was more than one-third complete. As shown in Figure 5.8.2-45, the collapsed liquid level dropped to almost the midpoint of the heated rod length during this phase of the transient. The average void fraction of the core two-phase mixture may be estimated by dividing the measured core collapsed liquid level by the [ ]** in. heated rod length, in this test, the minimum collapsed liquid level corresponds to a core void fraction that approached [ ).** By the end of the short-term transient, the effect of further accumulator, CMT, and IRWST injection ended all core boiling (see Figure 5.8.2-55), and the core was again water solid. The collapsed liquid level in the upper plenum region and the associated fluid mass are shown in Figures 5.8.2-49 and 5.8.2-48. During the period before accumulator injection, the upper plenum level decreased below the [ ]*'in. elevation and out of the LDP range. The start of accumulator injection provided the liquid necessary to quickly raise the collapsed upper plenum liquid level back to the elevation of the hot legs and above. Following the end of the accumulator injection, the upper plenum level decreased once again, but remained within the region spanned by the LDP until IRWST injection supplied inventory to initiate a refill. The upper plenum contained a two-phase mixture level at the minimum mass inventory point at the time of accumulator injection since the core outlet quality never exceeded a value of [ ].** The upper plenum was again water-solid by the end of the short-term transient. Figures 5.8.2-50 and 5.8.2-51 show that the upper head lost inventory rapidly during the initial [ ]** seconds of the transient, partially refilled, then drained completely after accumulator injection ceased. This refilling behavior is the opposite of what is typical of a continuous strong mass flow through the open ADS 1-3 valves. The influx of cooler accumulator water produced a significant amount of condensation in the RPV downcomer, which drew liquid from the upper plenum into the upper head region. Once the accumulator was empty, injection flow decreased since it was provided maap60cc344w-58.noa:ib-too395 REVISION: 1 5.8.2-2
.. . - - . - - . - - . - - n- - - . - - .-. . - - - . - . - . -
i e i-L l l [, i solely by CMT draindown. Downcomer condensation was no longer adequate to hold liquid any l longer in the RPV upper head, which in turn drained its water into the RPV upper plenum. When the 4 RPV mass increased due to continuous IRWST injection, upper head inventory was replenished again by downcomer condensation as occurred during accumulator injection (Figure 5.8.2-40) even as the core inventory increased (Figure 5.8.2-44). 4 Re mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.8.2-41 and j 5.8.2-42. He downcomer level fell rapidly upon ADS-1 actuation to a minimum value, then j recovered after accumulator injection.- De downcomer level only fell to the [ J'6' in. elevation at ! [ ]'6* seconds, implying that the upper plenum minimum level equaled about [ ] in., further
- confirming that the core remained covered during this portion of the test. Later in the transient, the downcomer refilled to the hot-leg and cold-leg piping elevations during IRWST injection.
5.8.2.2 Energy Transport from the Primary System ] 1 j Following ADS initiation, energy was deposited in the primary system fluid by the heater rods { simulating decay heat and by the primary system metal as it cooled down. Since fluid energy was lost j to the ambient, energy must be removed from the primary system to prevent excessive fluid and heater
- rod temperature excursions. De AP600 is designed to remove heat by a combination of the SGs, the PRHR, and the ADS.
1 i ( Steam Generator and Passive Residual Heat Removal Heat Transfer i During normal operation, most of the primary system heat was removed via the SGs; however, once l the coolant pumps tripped, reduced system flow reduced primary- to secondary-side heat transfer. The SGs were only avaliable as heat sinks until the primary system pressure dropped to that of the secondary side. He PRHR is designed to remove heat from the primary system once the safety signal opens the PRHR isolation valve. 'In test SB14 the PRHR continued to remove energy after the SGs were thermally isolated. Due to actuation of the ADS at time zero in this event, the ADS valve ! venting quickly became the preferred path for removal of energy from the primary system. l Figure 5.8.2-33 shows the SG pressure equalization together with the PRHR integrated heat transfer, as represented by the IRWST fluid energy after allowing for the contribution from the ADS 1-3 inflow. Heat was transferred to the secondary side of the SGs for only the first [ ]'6# seconds of the transient, and the SG tubes drained by [ ]'6# seconds. PRHR heat removal began after the isolation valve opened [ ]'6' seconds into the test. De PRHR was responsible for only a portion of the IRWST heat-up during this transient due to ADS-1 activation at time zero. As shown in Figure 5.8.2-21, the PRHR drained completely at about [ ]'6" seconds into the transient. Figure 5.8.2-22 can be disregarded after [ ]'6' seconds since FMM-802 and -804 instruments were inoperable. ( maarom2344w.5s.non:16.too395 5.8.2-3 REVISION: 1
l l Energy Transport via the Ilreak and Automatic Depressurization System- ' There was no flow of mass out of the primary system via any simulated break in test SBl4, so Figures 5.8.2-67 and 5.8.2-68 are irrelevant. Initiation of ADS 1-3 provided a programmed, increasing l mass release path with no other release path available until ADS-4 was activated based on the CMT low-low level signal. 'Ihe initiation of ADS stages 1,2 and 3 caused the system to depressurize rapidly. At [ ]** seconds, ADS-4 initiated, and the primary system continued to depressurize to containment pressure. During the inadvertent ADS actuation transient, steam and liquid flowed through the ADS 1-3 valves at peak rates which approached [ ] and [ ]** lbm/sec, respectively, as shown on Figures 5.8.2-65 l and 66. Flow through the ADS continued at a reduced rate thereafter, increased as the accumulators emptied, then diminished once again. After [ ]** seconds, flow through ADS 1-3 was negligible for the rest of the short term transient and was replaced by flow through the lower-resistance ADS-4 paths. By the end of the short-term transient, water was flowing out of the two ADS-4 valves at approximately [ ]** lbm/sec. (Figure 5.8.2-64). Integrated mass flow out of the primary system via the ADS is shown in Figure 5.8.2-62, and the corresponding integrated energy flow is shown in Figure 5.8.2-69. The total system inventory plot (Figure 5.9.2-70) indicates that a net mass increase of [ ]** lbm occurred during the short-term transient. The energy balance components are presented in Figure 5.8.2-74. l O mAyfM2344w-58.non:IM00395 5.8.24 REVIslON: 1
b TABLE 5.8.21 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.8.2 Plot No. Component Variables Units Description 1 Pressurizer CPT-6N psia System pressure 2 RPV RPVPWR kW Core power 3 RPV TOIRPV, 'IV8RPV, 'F Core inlet / outlet temperature, ST08RPV saturation temperature 4 SG CPT 201, Cirf 2N, psia Prunary and secondary pressures in SG CirT-301, CFT-302 5 DVI-1 WWTDVIL1, Ibm /sec. Individual components and total flow in WWTIRWil, DVI l WOUTACCl, WWTIRWI3 6 DVl-2 WWTDVIL2, Ibm /sec. Individual components and total flow in WWTIRWI2, DVl-2 WOUTACC2, WWTIRW14 7 CMT AMCMTIB, Ibm Fluid mass in CMrs (excludes balance AMCMT2B lines) ! 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT MlWDVIL1, Ibm Integrated mass out of CMTs , MIWDVIL2 t , 10 CMT WWIDVIL1, Ibm /sec. Flow out of CMTs I WWTDVIL2 j 11 CMT WOUTCLB1, Ibm /sec. Flow into CMTs I WOUTCLB2 12 CMT CLDP-509, CLDP510 in. Level CL-CMT balance hnes 13 CMT UCMT1, U('MT2 Btu Fluid energy in CMTs 14 IRWST IRWST lbm Mass of fluid in IRWST 15 IRWST CLDP-701 in. Collapsed liquid level in IRWST 16 IRWST WWTIRWil, Ibm /sec. Flow from IRWST to DVI lines l WWTIRW12 17 IRWST IRWSTOR Ibm /sec. Overflow from IRWST to sump 18 IRWST ADS 13TMR lbm/sec. Total ADS flow into IRWST 19 IRWST ADS 13TIR, MIIRWII, Ibm Integrated mass out of IRWST MIIRWI2, MIIRW10 20 IRWST UIRWST Btu Fluid energy in IRWST 21 PRHR CLDP-802 in. Collapsed liquid level in PRIIR llX
/
r m:\.p600c344w.58.noo: b too395 5.8.2-5 REVISION: 1
TABLE 5.8.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.8.2 Plot No. Component Variables Units Description 22 PRHR WWOTPRHR Ibm /sec. Measured outlet flow from PRHR tube 23 Accumulator AMACCl, AMACC2 lbm Mass of fluid in accumulators 24 Accumulator CLDP-401, CLDP-402 in. Collapsed liquid level in accumulators 25 Accumulator WOUTACC1, Ibm /sec. Flow from accumulators WOUTACC2 26 Accumulator MOUTACCl, Ibm Integrated mass out of accumulators MOUTACC2 27 Accumulator UACCl, UACC2 Btu Fluid energy in accumulators 28 Pnmary sump AMPSMP lbm Pnmary sump fluid mass 29 Pnmary sump CLDP-901 in. Pnmary sump level 30 Pnmary sump UPSMP Btu Pnmary sump fluid energy 31 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG primary side MSSGOP1, MSSGOP2 inlet / outlet plena 32 SG MSSGHT1, MSSGHT2, Ibm Mass of fluid in SG primary side hot MSSGCrl, MSSGCT2 and cold tubes 33 SG/PRHR Cirr-201, Cf'r-301, psia & SGI pressure and PRHR integrated heat QPRHR1 Btu output 34 Pressurizer PZM lbm Ruid mass in pressurizer 35 Pressurizer CLDP-601 in. Collapsed liquid level in pressurizer 36 Pressurizer UPZ Btu Fluid energy in pressurizer 37 Surge line PLM lbm Fluid mass in surge line 38 Surge line CLDP-602 in. Collapsed liquid level in surge line 39 Surge line UPSL Btu Fluid energy in surge line 40 RPV MWPRV lbm Total fluid mass in reactor vessel 41 RPV DCM lbm Fluid mass in downcomer 42 RPV LDP01DC in. Collapsed liquid level in downcomer compared to various reference elevations 43 RPV MWOIRPV lbm Fluid mass in lower plenum 44 RPV MWO3RPV lbm Fluid mass in core region 45 RPV LDP03RPV in. Collapsed liquid level in core 46 RPV RPVAVDF Core exit void fraction 47 RPV RPVAQOUT Core exit quality 48 RPV MWO6RPV lbm Fluid mass in the upper plenum 49 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 50 RPV MWO8RPV lbm Fluid mass in the upper head 51 RPV LDP08RPV in. Collapsed liquid level in the upper head mAap600\2344w-58.non:lb-100395 5.8.2-6 REVISION: I
. ~ . - - .. ~..-.._ ... ..- -.
l l , l ! O) i
\_. TABLE 5.8.21 (Continued)
OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.8.2 Plot No. Comi.onent Variables Units Description i 52 RPV URPV Btu Total fluid energy in reactor vessel 53 RPV RPVXE,RPVASL in. Level of Tsat line 54 RPV RPVPab, RPVAPab, kW Heated rod power above and below RPVPWR Tsat level and total 55 RPV RPVRXV,RPVASOUT Ibm /sec. Core steam generation rate 1 56 RPV RPVALIN lbm/sec. Calculated core flow ] 57 RPV HTMXRPV, STD8RPV 'F Maximum clad temperature and I saturation temperature 58 Hot leg MWHL1, MWHL2 lbm Water mass in bot legs 59 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs , 60 Cold leg CLlWMS, CL2WMS, Ibm Water mass in cold legs l CL3WMS, CL4WMS 61 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 62 ADS and bicak BRKSTIR, ADS 13TIR, !bm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS 4s, and break 63 ADS and break BRKTIVF, AD13TIVF, Ibm Totalintegrated vapor flow for ADS AD41TIVF, AD42TIVF and break 64 ADS and break BRKTILF, AD13TILF, Ibm Totalintegrated liquid flow for ADS ; AD41TILF, AD42TILF and break l 65 ADS and break ADS 13SVR, Ibm /sec. Vapor flow out ADS 1-3 and ADS-4 ADS 41SVR, ADS 42SVR 66 ADS and break ADS 13SLR, Ibm /sec. Liquid flow out ADS 1-3 and ADS-4 ADS 41SLR, ADS 42SLR
~
67 ADS and break BRKSVLR lbm/sec. hm i flow out of break 68 ADS and break BRKSSLR lbm/sec. Liquid I5070ut of break 69 ADS and break BRKSPEi, ADS 13EI, Btu Integrated fluid energy for ADS 1-3, ADS 41EI, ADS 42EI ADS-4, and break 70 Mass balance TOTMASS lbm Total system mass inventory 71 Mass balance PRIMMASS, Ibm Measured prunary system inventory and PRIMASS2 value from mass balance 72 Mass balance MERROR lbm Mass balance error 73 Mass balance MIN, MOUT lbm Integrated mass flow in and out of SRCMASS primary system and source mass 74 Energy balance Various Btu Components of energy balance O V m:W344w.58.non:tb.ioo395 5.8.2-7 REVISION: I
l l i TIE FIGURES LISTED IN TABLE 5.8.2 I ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT O O m:\ap6aA2344w 58.noo:Ib too995 5.8.2-8 REVISION: 1
5.8.3 Long Term Transient - The long-term transient started with the initiation of IRWST injection, covered the transition from IRWST to sump injection, and provided information on the LTC response of the AP600 with sump injection. For the inadvertent ADS actuation case, Matrix Test SBl4, the long-term transient phase encompassed the time frame of [ ] seconds to the end of the test, around [ J'6# seconds. De behavior of the test facility during this period of the transient is discussed in this subsection using the plot package detailed in Table 5.8.3-1. De analysis concentrates on the components of the primary system that remained active during the LTC phase, that is the RPV, the het legs, ADS 4, the sumps and the IRWST. During the long-term transient, thermal-hydraulic phenomena of interest were: Maintenance of core cooling and removal of energy from the primary system. Level and pressure oscillations observed during test SB01 and other long-term transients that did not occur during SB14, 5.8.3.1 Maintenance of Core Cooling Mass Injected into Primary System De total DVI line flow, CMT flow, and IRWST flows are shown in Figures 5.8.3-6 and 5.8.3-7. Flow from the primary sump is shown in Figure 5.8.3-19. During the pre-sump injection phase of the long-term transient, the IRWST flow proceeded continuously at a diminishing rate. At [ J'6# seconds, flow from the primary sump started through the check valves around the main injection valves further reducing IRWST flow. From [ ]# seconds to the end of the transient, a nearly steady flow rate of [ ]'6* lbm/sec. was maintained through each DVI line. At ( ]'6# seconds, the main sump injection valves opened. His resulted in a reversal of flow through the IRWST injection line-1, while the flow in IRWST line-2 remained relatively constant. De net injection flow rate from the IRWST during sump injection was about [ ]'6# lbm/sec through both DVIlines. Reactor Pressure Vessel and Downcomer Response I he effect of water inflow on the average measured downcomer fluid, core inlet and core outlet, and peak clad temperatures during the long-term phase of the transient is shown in Figures 5.8.3-4,5.8.3-5, and 5.8.3-38. Figure 5.8.3-4 shows that there was a gradual increase in average downcomer fluid temperatures during the long-term transient once the sump began injecting. By the end of the test, j the average core inlet temperature reached a value of [ ]' 6# below saturation. Figure 5.8.3-5 1 shows that the core remained at or near saturation for the entire long-term transient after [ ]'bd maapaxu344. 5s.non:ih.ioo395 5.8.3-1 REVISION: 1
seconds. Figures 5.8.3-34 to 5.8.3-36 show that the DVI line flow method discussed in Subsection 4.11 indicated that a small level of boiling was maintained after [ ]'6# seconds into the transient. Nevertheless, the level of boiling was small, and the test showed that the inflow from the IRWST and sumps was sufficient. Figure 5.8.3-38 shows that there are no significant excursions in heater rod temperatures throughout the long-term transient. Herefore, sufficient core inventory and flow was maintained through this phase of the transient to remove the decay heat generated. For significant portions of the transient, a two-phase mixture was present in the core and upper plenum regions. De following discussion explains the variation in water level and mass throughout the reactor vessel and downcomer, ne mass and level for the core region are shown in Figures 5.8.3-28 and 5.8.3-29. He collapsed liquid level in the core indicates that the heated rods were always covered with a single- or two-phase mixture. During the later stages of the transient, the collapsed level remained just below the top of the heated rods and the core void fraction was about [ J.'b' The fall in core inventory was a result of the influx of hot water from the primary sump as it flowed through the check valves. The impact of this hot water on the system temperatures is shown in Figures 5.8.3-4 and 5.8.3-5 as a sudden increase in fluid temperature in the downcomer and at the core inlet. The hotter water also led to an increase in the calculated steam generation rate shown in Figure 5.8.3-36 and a corresponding fall in the level at which the core reached saturation temperature (Figure 5.8.3-34). The mass of water in the reactor pressure vessel is shown in Figure 5.8.3-25. After an initial decline, the reactor vessel water mass settled at an average value of 450 lbm until sump injection started, when it gradually fell to [ ]*'lbm, which is more than [ ]'6# percent of the initial vessel water inventory. The collapsed liquid level in the upper plenum region is shown in Figure 5.8.3-32. His figure shows that the collapsed liquid level initially fell, yet remained above the mid-level of the hot legs during the period before sump injection. Following a surge in level at the initial influx of hot water from the sumps, the level dropped first to the hot leg mid-elevation and then to the top of the DVI injection lines where it remained for the remainder of the transient. This corresponds to a void fraction of [ J.*b# The mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.8.3-26 and 5.8.3-27. The collapsed liquid level remained above the DVI lines for the entire long-term transient. De start of sump injection reduced the level, but this was not enough to uncover the DVI lines. SBl4 did not exhibit the RPV level / pressure oscillations observed in several of the other OSU small break LOCA tests. l
<1 l
l 9l l mAap60m2344w-58.non:n>. loo 395 5.8.3-2 REVISION: 1 l
_ _ _ . . . ~ _ . _ . _ _ _ . _ _ _ _ _ _ _ _ __ __ . _ . . _ _ _ . _ . _ . _ . 5.8.3.2 Energy Transport from the Primary System During the long-term transient, energy continued to be deposited in the primary system from the heater rods, the metal, and the fluid flowing from the primary sump. The SGs and PRHR remained inactive throughout this phase of the transient; the only paths out of the r imary system were via the ADS 1-3 and ADS-4 ADS valves. Integrated mass flow out of the primary system via the ADS is shown in Figure 5.8.343. By the end of the transient simulation, over [ ]lbm of water flowed out of the primary system. During the long-term cooling phase of the transient, the most significant outflow was through the ADS-4 valves, as confirmed by Figure 5.8.3-44, which depicts the flows through the ADS. During the sump injection phase of the transient, outflow was primarily in the form of liquid out of the ADS-4 valves; water flowed through each of the valves at an average rate of [ ]'"# lbm/sec. In addition, there was some small mass flow out of the open ADS 1-3 valves. Figure 5.8.3-36 shows the calculated steam generation rate determined using the DVI line flow method. During the sump injection phase of the transient, steam was generated at over [ J'6# Ibm /sec., but there was no evidence from the vortex meters for steam flow out of the ADS-4 valves. However, there is evidence that steam was leaving the primary system by this route: Figure 5.8.3-46 shows the measured, total system fluid inventory. During this phase of the ( transient, the interval after the start of primary sump injection is when core steam generation is most significant, and the total system inventory fell by about [ ]lbm. This corresponds to a steam flow rate of [ ]'6# lbm/sec, which would not have been detected by the vortex meters.
. Examination of the fluid thermocouples on the outlet of the ADS-4 valves indicates that the thermocouples remained at or above saturation temperature following the start of sump injection.
It was not possible for all the steam generated in the core to flow from the upper head to the downcomer via the bypass holes (Subsection 6.1.3). 'Iherefore, steam did leave the primary system via the ADS-4. Figure 5.8.3-50 shows the components of the system energy balance. Further discussion of steam loss from the primary system is provided in the mass and energy balance discussions of Subsection 6.2.2 and 6.2.3. m%r600cm-58.non:Ib.noo995 5.8.3-3 REVISION: 1
l l l 1 TABLE 5.83-1 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.8.3 9 LONG TERM TRANSIFET Plot No. Component Variables Units Description 1 RPV RPVPWR kW Core power 2 Primary sump TSMPII, TSMP12 'F Sump injection line temperatures 3 DVI TDVIL1, TDVIL2 'F DVI line temperatures 4 RPV T01DC, It)2DC, T03DC, 'F Water and saturation temperatures in STOIDC downcomer 5 RPV TOIRPV,11)8RPV, 'F Core inlet / outlet temperature, ST08RPV saturation temperature 6 DVI l WWTDVIL1, Ibm /sec. Individual components and total flow WWTIRWII, in DVI-l WWTIRWI3 7 DVI 2 WWTDVIL2, Ibm /sec. Individual components and total flow WWTIRW12, in DVI 2 WWTIRW14 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT CLDP-509, CLDPS10 in. Level CL-CMT balance lines 10 IRWST IRWST lbm Mass of fluid in IRWST 11 1RWST CLDP-701 in. Collapsed liquid level in IRWST 12 IRWST UIRWST Btu ' Fluid energy in IRWST 13 Primary sump AMPSMP lbm Pdmary sump fluid mass 14 Pnmary sump CLDP-901 in. Pnmary sump level 15 Primary sump UPSMP Btu Pnmary sump fluid energy 16 Secondary sump AMSSMP lbm Secondary sump fluid mass 17 Secondary sump CLDP-902 in. Secondary sump level 18 Secondary sump USSMP Btu Secondary sump fluid energy 19 Primary sump WSTSMPET, WWTSMPIT Ibm /sec. Pnmary sump steam and liquid injection rate 20 Pnmary sump MISMPII, MISMP12, Ibm Integrated primary sump and IRWST MISMPIT, MIIRWT flows 21 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG side inlet / outlet MSSGOP1, MSSGOP2 plena 22 Surge line PLM lbm Fluid mass in surge line 23 Surge line CLDP-602 in. Collapsed liquid level in surge line 24 Surge line UPSL Btu Fluid energy in surge line 25 RPV MWRPV lbm Total fluid mass in reactor vessel 9 m:W344w-58.non:lt> too395 5.8.3 4 REVISION: 1
I i l 1 I ( ' TABLE 5.8.31 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.8.3 LONG TERM TRANSIENT l Plot No. Component Variables Units Description 1 26 RPV DCM lbm Fluid mass in downcomer 27 RPV LDP01DC in. Collapsed liquid level in downcomer l compared to various reference I elevations l 28 RPV MWO3RPV lbm Fluid mass in core region ! 29 RPV LDP03RPV in. Collapsed liquid level in core 30 RPV P AVDF Core exit void fraction 31 RPV RPVAQOUT Core exit quality 32 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 33 RPV MWO8RPV lbm Fluid mass in the upper head 34 RPV RPVASL in. Level of Tsat line 35 RPV RPVAPab, RPVPWR kW Heated rod power above and below l Tsat level and total 1 36 RPV RPVASOUT lbm/sec. Core steam generation rate I 37 RPV RPVALIN Ibm /sec. Calculated core flow I 3 38 RPV HTMXRPV, "F Maximum clad temperature, saturation l i S11)8RPV -V temperature and delta 39 Hot leg MWHL1, MWHL2 lbm Water mass in hot legs 40 Hot leg MVHL1, MVHL2 lbm Vapor mass in hot legs 41 Cold leg CLIWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 42 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 43 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4, and break 44 ADS and break ADS 13TLR, ADS 41TLR, Ibm /sec. Liquid flow out ADS 1-3 and ADS-4 ADS 42TLR 45 ADS and break BRKSTLR lbm/sec. Liquid flow and total flow out of break l 46 Mass balance TOTMASS lbm Total system mass inventory 47 Mass balance PRIMMASS, PRIMASS2 lbm Measured prunary system inventory and valve from mass balances 48 Mass balance MERROR Ibm Mass balance error 49 Mass balance MIN, MOUT SRCMASS lbm Integrated mass flow in and out of primary system and source mass 50 Energy balance Various Blu Component of energy balance A d mAap60m2344w-58.non:1b.ioo395 5.8.3-5 REVISION: 1
e THE FIGURES LISTED IN TABLE 5.8.31 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT e e mhpuo2344w-58.non:15100995 5.8.3-6 REVISION: I
~ . . - - . _ -.- - - .- .- .. -- - . - --- . . . - - . - . - --
l r\ l Q 5.9 Analysis of Matrix Test SB15 1 1 Matrix test SB15 (OSU Test U0015) simulated a 2-in. hot-leg break LOCA with LTC and without the operation of the nonsafety-related systems. The break was at the bottom of HL-2 and except for the break location, this test was identical to SB01, including the simulated failure of one of the ADS-4 ! lines. ;
)
1 Analysis of Matrix Test SB15 is divided into three sections as follows:
- Facility performance is discussed in Subsection 5.9.1. It provides a brief outline of the response of the test facility; further details are available in the Final Data Report.*
Re short-term transient for SB15 encompassed the start of the simulation up to [ J'6# l seconds. This period includes blowdown, natural circulation, ADS and initial IRWST stages {' of the transient.
- Analysis of the long-term transient, SB15, encompassed the time frame from [ ]'6# seconds to the end of the test. His phase of the transient concluded with the IRWST injection phase to the initiation of sump injection. He long-term transient actually staned at IRWST injection, which is discussed as part of the short-term transient. Between the end of the short. )'
term transient to [ ]'6' seconds, the system remained relatively inactive with the exception O' of the CMT refill. At [ ]'6# seconds, CMF 1 began to refill and CMT-2 followed [ ]'6# seconds later. CMT refill phenomena is discussed further in Section 6.1.1 and the ! discussion of the long-term transient provided here begins at [ ]'6# seconds. l The discussion of the short- and long-term phase of the transient focuses on important thermal-hydraulic phenomena identified in the PIRT (Table 1.3-1). The mass and energy balance results are key indicators of the quality of the analysis on which this discussion is based. Ecsc are discussed in detail in Subsections 6.2.2 and 6.2.3. l l v 9 I rnvemw.59.non:1b too395 5.9-1 REVISION: 1 a w- n w w --m-w+-
i a 5.9.1 Facility Performance - i 2 , The performance of the OSU test facility during Matrix Test SB15 in reference to the five transient phases is outlined in the following :
. Blowdown
- Natural circulation
= ADS
= IRWST injection
- Sump injection The overall performance of the facility during the transient is shown in Figures 5.9.1-1 to 5.9.1-4.
Figure 5.9.1-1 shows the pressurizer pressure throughout the test with various phases and operating i components. De time scale was reduced for clarity since there were only small changes in system l pressure during the long-term phase of the transient. Figure 5.9.1-2 shows the total DVIline flow and l its composition from the various sources at each time in the transient. Figure 5.9.1-3 shows the
]
calculated core steam generation rate throughout the test and Figure 5.9.1-4, the variation in core outlet ' temperature and peak clad temperature relative to the core outlet saturation temperature. Figures 5.9.1-1 and 5.9.1-2 show that there was a continuous flow of cool water to the core from the - O passive safety systems throughout the transient. Once initiated, the ADS lines rapidly depressurized the primary system, which enhanced the CMT and accumulator injection flow rates. Ultimately, the ADS-4 valves sufficiently reduced the circuit pressure to start gravity-driven IRWST injection. De I passive injection systems overlapped so that as one source of water drained, the next was available to continue the cooling process. He level of steam generation in the core and the response of the core outlet fluid temperatures and maximum clad temperatures are shown in Figures 5.9.1-3 and 5.9.1-4. ! These figures show that the cooling flow prevented any excessive core heating. He core remained subcooled for large periods of the transient and when steam production occurred, the rate of generation remained well below the rate at which water was delivered to the core. 5.9.1.1 Blowdown Phase ne blowdown phase began at time zero when the break was initiated and continued until the primary circuit pressure was nearly in equilibrium with the secondary-side pressure at around [ ]"# seconds. Immediately following the opening of the break, primary circuit pressure decreased as the small hot-leg break removed the energy being added from the core. Pressure fell gradually until the end of the blowdown phase, During this phase of the transient, cooling flow was provided from the two CMTs, which remained in the recirculation mode until around [ ]' b'* seconds. Heat was removed from the primary circuit via the PRHR and the SGs. The pressurizer and surge line completely drained at [ ]** and [ ]** seconds, respectively. O) L. m%60m344 59.non:lt>100395 5.9.1-1 REVISION: 1
5.9.1.2 Natural Circulation Phase In this LOCA simulation, the single- and two-phase natural circulation phase was marked by a gradual reduction in system pressure rather than by the more stable pressure observed in SB01. During this phase of the transient, the SG tubes all drained by about [ ]'6* seconds and at that time, heat j I removal from the primary circuit continued via the PRHR. The steam in the SG tubes became superheated and remained so until the end of the transient. In response to voiding in CL-3, CMT-1 transitioned to draindown mode at [ ]'6# seconds, and the falling CMT level reached the ADS low-level setpoint at [ ]'*# seconds. The natural circulation phase of the transient continued to [ ]'6# I seconds when the ADS-1 valve opened. l 1 5.9.1.3 Automatic Depressurization Phase i ADS-1 actuation was followed by ADS-2 and ADS-3 actuation at [ ]'*# and [ ]'6# seconds respectively. Accumulation injection began with initiation of the ADS. The influx of cold water combined with increased venting via the ADS led to a rapid depressurization of the primary circuit. I Actuation of ADS-4 at [ l'6* seconds completed depressurization to a level that initiated IRWST j injection at [ ]'** seconds via DVI-2 and [ ]'** seconds via DVI-1. During accumulator l injection, increased circuit resistance reduced flow out of the CMTs. CMT flow resumed as the , accumulators drained. 'Ihe accumulators were fully drained [ ]'*# seconds before IRWST l injection began. The CMTs did not fully drain until [ ]'6
- and [ ]'** seconds after the start of IRWST injection. The transfer from CMT accumulator to IRWST injection was indicated by the i minimum RPV inventory of [ ]'6# lbm at [ ]'*# seconds. )
1 Actuation of ADS-1 rapidly refilled the pressurizer as water and steam flowed out of the ADS. 'Ihe pressurizer gradually drained by [ ]'*# seconds. 5.9.1.4 In-Containment Refueling Water Storage Tank Injection l IRWST injection signals the transition from the short- to long-term phase of the transient. The initial ! phase of IRWST injection involved an increase in flow through the two DVI lines, which was followed by a gradual flow reduction as the driving head between the IRWST and the RCS fell due to the reduced IRWST water level. Once maximum flow was established, the inflow of water from the i IRWST was sufficient to keep the core subcooled from about [ ]'6' seconds. Steam was subsequently generated in the core for the remainder of the transient. Following the restart of core steam generation, IRWST injection between [ ]'*# seconds, was indicated by oscillations in delivery from the refilled CMT, and in pressure and level oscillations throughout the primary system. These oscillations were also observed in the ADS-4 ligtdd flow rates. O mAap600034w-59.non:Ib-100395 5,9.1 2 REVISION: 1
- . . - - . - - - - . . - - . . - . . . . . - - - . - - . ~ - - - - . - . . . . - . . - . - - . - . . - - - . -
i l L i j, 5.9.1.5 Sump Injection i ) Injection from the primary sump via the check valves around the main sump injection valves began at [ ]** seconds when the level in the IRWST was low enough to allow flow. This reduced the
- flow rate from the IRWST. Since the test was terminated [ ]** seconds beyond unis point, no j further analysis is provided of sump injection during long-term cooling in Test SB15.
i 1 I ) i 4 i e i to i i l 2 i 1 ) i b t m:W34w 59.non:1b-100395 5,9.13 REVISION: 1
TABLE 5.9.1-1 OSU TEST ANALYSIS PLOT PACKAGE FOR SUBSECTION 5.9.1 Ol Plot No. Component Variables Units Description 1 Pressurizer CPT-604 psia System pressure and event history 2 Water WWTDV11+WWTDV12, Ibm /sec. Total of CMT, accumulator, IRWST, injection WOUTACCl+WOUTACC2, and sump injection flows WWTIRWII+WWTIRW12, WWTSMPIT 3 Reactor RPVASOUT lbm/sec. Steam generation in reactor vessel vessel 4 Reactor T08RPV, HTMXRPV, TSAT *F Reactor vessel outlet temperature, vessel maximum clad temperature 2 ) fuel exit saturation temperature O l l l 9 m:\ap60m2344w 59. con:lb 100395 5.9.1-4 REVISION: 1
s THE FIGURES LISTED IN TABLE 5.9.1-1 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT 4 1 1 i i mAap600%2344w-59.noo:lb 100995 5.9.1-5 REVISION: 1
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m 8 5.9.2 Short-Term Transient For the 2-inch hot-leg break, Matrix Test SB]5, the short-term transient encompassed the time frame up to [ l seconds. As shown in Figure 5.9.1-1, this period included full depressurization of the facility through all four stages of the ADS together with CMT and accumulator injection plus the initial stages of IRWST injection. Variations in mass, energy, pressure, and temperature throughout this stage of the transient are illustrated in the plot package outlined in Table 5.9.21. He plots concentrate on the primary system, including the accumulators, CMTs, IRWST, primary sump, and flows from the primary system via the ADS, break, and IRWST overflow. For the short-term transient there were two principal concerns: Adequate flow from the passive systems to the reactor vessel must be maintained. Adequate flow into the core must be maintained to ensure that decay heat was removed from the simulated fuel rods without a temperature excursion. These concerns are addressed in the following discussion. 5.9.2.1 Maintenance of Core Cooling f% Mass Injected to the Primary System Figures 5.9.2-5 and 5.9.2-6 show the combined effect of the injection flows for the shon-term phase of the transient. Separate plots of the individual contributions to the total flow can be located by consulting the plot package index given in Table 5.9.2-1. Figures 5.9.2-5 and 5.9.2-6 show how the CMTs, accumulators and IRWST supply an almost continuous flow of cool water to the core. During the first [ ]'6 seconds, cooling flow was provided by the CMTs. The rate of flow from the CMTs gradually reduced from a maximum value of [ ]'6"lbm/sec. as the driving head fell in response to the CMT water heat-up until ADS-1 initiation. Rapid accumulator injection temporarily reduced CMT flow, but led to an overall increase in flow to the core to a peak value of [ l'6' lbm/sec. in DVI-2. Following the end of accumulator injection, the CMTs again provided cooling flow until drained. He only period in which there was relatively little cooling flow was during the initial CMT recirculation period at the stan of the transient. Reactor Pressure Vessel and Downcomer Behavior The effect of the water flow on core inlet / outlet temperatures and peak clad temperatures during the shon-term phase of the transient is shown in Figures 5.9.2-3 and 5.9.2-57. The combined CMT and accumulator flow was sufficient to keep the core completely subcooled between [ ]'** Q seconds. The core outlet temperature then remained at the saturation level until about [ ]'6# maap600c3.w59.aoa:ib.ioo395 5.9.2-1 REVIs10N: 1
seconds when the inflow of water from the IRWST became sufficient to again subcool the core. De core then remained subcooled until the end of the short-term transient. Figure 5.9.2-57 shows that there were no significant excursions in heated rod temperatures throughout the short-term transient; therefore, sufficient core inventory and flow was maintained through this phase of the transient to remove the decay heat generated. For significant portions of the transient, a two-phase mixture was present in the core and upper plenum regions, although core boiling was kept at a low level. De following discussion tracks the variation in water level and mass throughout the reactor vessel and downcomer. The mass and level for the core region are shown in Figures 5.9.2-44 and 5.9.2-45. The collapsed liquid level in the core indicates that the heated rods remained covered with a single- or two-phase mixture. De transient minimum core inventory of [ ]'"lbm occurred at about [ ]'** seconds into the transient, and is greater than the initial steady-state mass inventory. Figure 5.9.2-45 shows that the collapsed liquid level dropped to [ ]'** in. below the top of the heated rod length during the second boil-off period in this phase of the transient. This corresponds to a core void fraction of [ ].'" By the end of the short-term transient, the effect of IRWST injection ended all core boiling (Figure 5.9.2-55), and the core was again water-solid. De collapsed liquid level in the upper plenum region and the associated fluid mass are shown in Figures 5.9.2-49 and 5.9.2-48. These figures show that at the time of minimum core mass inventory, the upper plenum collapsed level remained above the hot-leg centerline. The end of accumulator injection coincided with the start of a fall in collapsed liquid level to the minimum elevation observed in SB15. Following the start of IRWST injection, the upper plenum collapsed level increased. De upper plenum was again almost water-solid by the end of the short-term transient. Figures 5.9.2-50 and 5.9.2-51 show that the upper head also gained inventory at the time when accumulator injection began. This behavior was the result of condensation from the large influx of cool accumulator water. Accumulator injection coincided with ADS-1 initiation. The flow of water and steam through the ADS rapidly refilled the pressurizer, which then removed water from the RPV. De loss of mass to the ADS valves and to the hot-leg break was responsible for minimizing the increase in overall RPV 8nventory during accumulator injection (Figure 5.9.2-40) as the core inventory also increased (Figure 5.9.2-44). The mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.9.2-41 and 5.9.2-42. The downcomer collapsed level remained at about the cold-leg elevation until ADS-1 actuation, when the level fell to the bottom of the DVI line piping. The downcomer level refilled to the top of the hot legs during accumulator injection, then fell back to just above the DVI line entry point, where it remained until IRWST injection once again raised the level to the hot legs. O maap6002w59.noo:ib-too395 5.9.2-2 REVISION: 1
- - -_ - - . . - - - .-_ _ ~ _ - - - . - . - . - - - - - - - 5.9.2.3 Energy Transport from the Primary System Following the break, energy was deposited in the primary circuit fluid by the heater rods to simulate , decay heat and the primary circuit metal as it cooled down. Some fluid energy was lost to the ambient and out of the break. Excess energy must be removed from the primary system to prevent excessive fluid and heater rod temperature excursions. The AP600 plant is designed to remove heat by a combination of the SGs and the PRHR and ADS. Steam Generator and Passive Residual Heat Removal Heat Transfer During normal operation, most of the primary system heat was removed via the SGs; however, once the coolant pumps tripped, the reduced circuit flow decreased primary to secondary-side heat transfer. J The SGs were only available as heat sinks until the primary system pressure dropped to that of the ! secondary side, then the two sides were in thermal equilibrium. De PRHR is designed to remove heat I from the primary system once the safety signal opens the isolation valve. De PRHR continued to remove energy after the SGs were thermally isolated until ADS actuated. Once the ADS is actuated, it became the predominant path for the removal of energy from the primary circuit. Figure 5.9.2-33 shows the SG pressure equalization together with the PRHR integrated heat transfer as represented by the IRWST fluid energy after allowing for the contribution from ADS 1-3 inflow. Heat was transferred to the secondary side of the SGs for only the first [ ]"' seconds of the transient. PRHR heat removal began [ ]*# seconds into the test. De PRHR was responsible for all the IRWST heat-up until ADS-1 activation, after which the PRHR heat transfer reduced significantly. During the active phase, the PRHR transferred heat to the IRWST at an average rate of [ ]'6* Btu /sec. Energy Transport via the Break and Automatic Depressurization System The mass flow rate from the primary system via the break is shown in Figures 5.9.2-67 and 5.9.2-68. As seen from these figures, liquid flow was detected by the flow meters for a short-term transient. During the first [ ]*# seconds following the break, [ ]*# lbm of water flowed out of the primary circuit via the break at an average rate of approximately [ }"# lbm/sec. During this period, the primary circuit depressurized to around [ ] psi (Figure 5.9.2-1). With initiation of ADS 1-3, vapor flow through the break ceased, but liquid flow continued well into the long-term cooling phase. Liquid and vapor flow through the ADS 1-3 valves began immediately upon actuation and continued until around [ ] seconds. By [ ]"" seconds, the time of ADS-4 activation, ADS 1-3 caused the circuit to depressurize rapidly to about [ ]*# psia. When ADS-4 was initiated, the primary circuit continued to depressurize to containment pressure. De initiation of the ADS 1-3 did not terminate the liquid flow through the break because the break is located at '.he bottom of the hot leg, upstream of the ADS 1-3 valves. The break flow became completely liquid flow because the RPV mass was adequate to continuously cover the hot-leg break mAap600 2344w.59.non:lb.100395 5.9.2-3 REVISION: I
elevation. He ADS 1-3 valves pass vapor and liquid at peak flow rates of about [
]*** lbm/sec., respectively. Flow through ADS l-3 continued at a declining rate until about
[ J'6* seconds when it was almost completely terminated and replaced by flow through the lower resistance ADS-4 paths. By the end of the short-term transient, water was flowing out of the two ADS-4 valves at approximately [ ]'** lbm/sec. (see Figure 5.9.2-64); water was also flowing out of the break at about [ ] lbm/sec. The integrated mass flow from the primary system via the ADS and the break is shown in Figures 5.9.2-62, and the corresponding integrated energy flow is shown in Figure 5.9.2-69. The total system inventory plot given in Figure 5.9.2-70 indicates that up to [ ]'** lbm ofinventory was lost at the limiting point in time during the short-term transient. The lost inventory was steam, none of which had been detected on the flow meters; thus, this is not included in the energy balance shown in Figure 5.9.2-74. I s 9 1 l I O mAap60tA2344w 59.non:lb-100395 5.9.24 REVISION: 1
O TABLE 5.9.2.I OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.9.2 Plot No. Component Variables Units Description 1 Pressurizer CirT-6(4 psia System pressure 2 RV RPVPWR kW Core power 3 RV TOIRPV, T08RPV, 'F Core inlet / outlet temperature, ST08RPV saturation temperature 4 SG CPT-201, CPT 204, psia Pnmary and secondary pressures in SG CPT-301, CI"T-302 5 DVI-l WWTDVIL1, Ibm /sec. Individual components and total flow in WWTIRWII, DVI-l WOLTTACCl, WWTIRWI3 6 DVI-2 WWIDVIL2, Ibm /sec. Individual components and total flow in WWTIRWI2, DVI-2 WOtJTACC2, WWTIRWI4 7 CMT AMCMTIB, Ibm Fluid mass in CMTs (excludes balance AMCMT2B lines) 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT MIWDVIL1, Ibm Integrated mass out of CMTs MIWDVIL2 OI 10 CMT WWTDVILI, ihm/sec. Flow out of CMTs WWTDVIL2 11 CMT WOUTCLB1, Ibm /sec. How into CMTs WOUTCLB2 12 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 13 CMT UCMT1, UCMT2 Btu Fluid energy in CMTs 14 IRWST IRWST lbm Mass of fluid in IRWST 15 IRWST CLDP-701 in. Collapsed liquid level in IRWST 16 IRWST WWTIRWII, Ibm /sec. Flow from IRWST to DVI lines WWTIRWI2 17 IRWST IRWSTOR lbm/sec. Overflow from IRWST to sump 18 IRWST ADS 13TMR lbm/sec. Total ADS flow into IRWST 19 IRWST ADSI3TIR, MIIRWII, Ibm Integrated mass out of IRWST
.MIIRWI2, MIIRWlO 20 IRWST UIRWST Btu Fluid energy in IRWST 21 PRHR CLDP-802 in. Collapsed liquid levelin PRHR HX m:W344w-59.noa:Ib-100395 5.9.2-5 REVISION: 1
l i TABLE 5.9.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.9.2 Ol Plot No. Component Variables Units Description 22 PRHR WWOTPRHR lbm/sec. Measured outlet flow from PRHR tube 23 Accumulator AMACC1, AMACC2 lbm Mass of fluid in accumulators 24 Accumulator CLDP-401, CLDP-402 in. Collapsed liquid level in xcumulators 25 Accumulator WOUTACCl, Ibm /sec. Flow from accumulators WOUTACC2 26 Accumulator MOUTACCl, Ibm Integrated mass out of accumulators MOUTACC2 27 Accumulator UACCl, UACC2 Btu Fluid energy in accumulators 28 Pnmary sump Ah1PSMP lbm Prunary sump fluid mass 29 Pnmary sump CLDP-901 in. Pnmary sump level 30 Pnmary sump UPSMP Btu Pnmary sump fluid energy 31 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG primary side . MSSGOP1, MSSGOP2 inlet / outlet plena l 32 SG MSSGHT1, MSSGHT2, Ibm Mass of fluid in SG pnmary side bot MSSGCT1, MSSGCr2 and cold tubes 33 SG/PRHR Cirr 201, CPT-301, psia & SGI pressure and PIGIR integrated heat l QPRHRI Btu output ' 34 Pressurizer PZM lbm Fluid mass in pressurizer 35 Pressurizer CLDP-601 in. Collapsed liquid level in pressurizer l 36 Pressunzer UPZ Btu Fluid energy in pressurizer 37 Surge line PLM lbm Fluid mass in surge line 38 Surge line CLDP-602 in. Collapsed liquid level in surge line 39 Surge line UPSL Btu Fluid energy in surge line 40 RV MWRPV lbm Total fluid mass in reactor vessel 41 RV DCM lbm Fluid mass in downcomer 42 RV LDP01DC in. Collapsed liquid level in downcomer compared to various reference elevations 43 RV MWOIRPV lbm Fluid mass in lower plenum 44 RV MWO3RPV lbm Fluid mass in core region 45 RV LDP03RPV in. Collapsed liquid level in core 46 RV RPVAVDF Core exit void fraction 47 RV RPVAQOlTT Core exit quality 48 RV MWO6RPV lbm Fluid mass in the upper plenum 49 RV LDP06RPV in. Collapsed liquid level in the upper plenum 50 RV MWO8RPV lbm Fluid mass in the upper bead 51 RV LDP08RPV in. Collapsed liquid level in the upper head m:\a;fM2344w 59.noa:Ib-loo 395 5.9.2-6 REVISION: 1
0 ( TABLE 5.9.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECrlON 5.9.2 Plot No. Component Variables Units Description 52 RV URPV Btu Total fluid energy in reactor vessel 53 RV RPVXE.RPVASL in. Level of Tsat line 54 RV RPVPab, RPVAPab, kW Heated rod power above and below RPVPWR Tsat level and total 55 RV RPVRXV,RPVASOUT Ibm /sec. Core steam generation rate 56 RV RPVALIN lbm/sec. Calculated core flow 57 RV HTMXRPV, ST08RPV *F Maximum clad temperature and saturation temperature 58 Hot leg - MWHL1, MWHL2 lbm Water mass in bot legs 59 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs 60 Cold leg CLlWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 61 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 62 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discbarged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4s, and break 63 ADS and break BRKTIVF, ADl3TIVF, Ibm Total integrated vapor flow for ADS AIMITIVF, AIM 2TIVF and break [ 64 ADS and bn:ak BRKTILF, ADl3TILF, Ibm Total integrated liquid flow for ADS AD41TILF, AD42TILF and break 65 ADS and break ADS 13SVR, Ibm /sec. Vapor flow out ADS 1-3 and ADS-4 ADS 41SVR, ADS 42SVR 66 ADS and break ADS 13SLR, Ibm /sec. Liquid flow out ADS 13 and ADS 4 ADS 41SLR, ADS 42SLR 67 ADS and break BRKSSVR Ibm /sec. Vapor flow out of break 68 ADS and break BRKSSLR Ibm /sec. Liquid flow out of break 69 ADS and break BRKSPEI, ADS 13EI, Btu Integrated fluid energy for ADS 1-3, ADS 41EI, ADS 42El ADS 4, and break 70 Mass balance TOTMASS lbm Total system mass inventory 71 Mass balance PRIMMASS, Ibm Measured pnmary system inventory and PRIMASS2 value from mass balance 72 Mass balance MERROR Ibm Mass balance error 1 73 Mass balance MIN, MOUT lbm Integrated mass flow in and out of SRCMASS pnmary system and source mass 74 Energy balance Various Btu Components of energy balance b d I m:\ap600\2344w 59.noo:Ib.10o395 5.9.2-7 REVISION: 1
O i Tile FIGURES LISTED IN TABLE 5.9.21 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT l 1 l O O m:W344,.59.non:lb-100995 . . -8 REVISION: 1 i
.m 5.9.3 Long Term Transient
'Ihe long-term transient started with initiation of IRWST injection, covered the transition from IRWST to sump injection, and provided information on the LTC response of the AP600 plant. For the 2-inch )
hot leg break, Matrix Test SB15, the long-term transient analyzed runs from [ l seconds to the ; end of the test at near [ ]'** seconds. The behavior of the test facility during this period of the
)
transient is discussed in this subsection using the plot package detailed in Table 5.9.3-1. This analysis concentrates on the components of the primary system that remained active during the LTC phase, that is, the RPV, the hot legs, ADS-4, the sumps, and the IRWST. l During the long-term transient, the main thermal-hydraulic phenomena of interest were: Maintenance of core cooling and removal of energy from the primary circuit. Level oscillations (from [ l'6' seconds there were system wide level and pressure oscillations, which are discussed further in Subsection 6.1.3) 5.9.3.1 Maintenance of Core Cooling Mass Injected into Primary System l l% l Total DVI line flow, CMT flow, and IRWST flows are shown in Figures 5,9.3-6 and 5.9.3-7, and flow from the primary sump is shown in Figure 5.9.3-19. From around [ ] seconds, there was a contribution to the DVI flow from the CMTs as the CMTs reached post-refill draindown. I During the pre-sump injection phase of the transient, IRWST flow proceeded at a gradually declining rate with the effect of the primary system oscillations superimposed. At [ ] seconds, flow from the primary sump began through the check valves around the main injection valves resulting briefly in a further reduction in IRWST flow. ! Reactor Pressure Vessel and Downcomer Response i
'Ihe effect of the water inflow on the downcomer fluid temperatures, core inlet and core outlet temperatures, and peak clad temperatures during the long-term phase of the transient is shown in
! Figures 5.9.3-4,5.9.3-5, and 5.9.3-38. Figure 5.9.3-4 shows that there was general increase in average l downcomer fluid temperatures during the long-term transient. By the end of the CMT injection phase, j this average temperature reached a value about [ ]'6# 'F below saturation. It decreased fairly rapidly ! thereafter as the colder IRWST was the sole water source, then increased once again as sump liquid
- entered the RPV via the check valve flow path. Figure 5.9.3-5 implies that the core remained subcooled for the entire long-term transient. However, as shown in Figures 5.9.3-34 to 5.9.3-36, the DVI Line method indicates that a small level of boiling was maintained after [ ]'6# seconds into m
- \ap600G344w.59.noa:1b-100395 5.9.3 1 REVISION: 1
the transient. Nevertheless, the level of boiling wns small and showed that the inflow from the IRWST was sufficient. Figure 5.9.3 38 shows that there were no significant excursions in heated rod temperatures throughout the long-term transient; therefore, sufficient core inventory and flow was maintained through this phase of the transient to remove the decay heat generated. For significant portions of the transient, a two-phase mixture was present in the core and upper plenum regions. The following discussion tracks the variation in water level and mass throughout the reactor vessel and downcomer. The mass and level for the core region are shown in Figures 5.9.3-28 and 5.9.3-29. The collapsed liquid level in the core indicated that the heated rods were always covered with a single- or two-phase mixture. During the later stages of the transient, the collapsed liquid level remained above the top of the heated rods. Flow through the check valves of the hot water from the sump is shown in Figures 5.9.3-4 and 5.9.3-5 as a sudden increase in fluid temperature in the downcomer and at the core inlet. The hot water also led to an increase in the calculated steam generation rate, as shown in Figure 5.9.3 36. The collapsed liquid level in the upper plenum region is shown in Figure 5.9.3-32. The figure indicates that the collapsed liquid level initially fell but then remained at the top of the hot legs during the period before sump injection began. Following the influx of relatively hot water from the sumps, the test was terminated at [ ]'6* seconds. The mass of water in the RPV is shown in Figure 5.9.3-25. After an initial decline, the reactor vessel water mass settled at an average value of [ l'6* lbm until the refilled CMTs emptied, when it gradually rose to [ ]'*
- lbm, which is more than [ l'** percent of the initial vessel water inventory.
From [ ]'6* seconds, oscillations in vessel inventory were observed. For SB15, these oscillations are comparable in magnitude to those observed in SB01. Figures 5.9.3-51 to 5.9.3-56 illustrate these oscillations using plots on a restricted time frame from [ j'** seconds. These oscillations are observed in primary system measurements from the upper plenum to to the ADS-4 flows. The oscillations have a period [ J.'*' These oscillations and possible mechanisms for their production are discussed further in Subsection 6.1.3. The mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.9.3-26 and 5.9.3-27. The collapsed liquid level remained above the cold legs for the entire long-term transient up until the time at which the CMTs were empty. ' 5.9.3.2 Energy Transport from the Primary System During the long-term transient, energy continued to be deposited in the primary system from the heated rods, metal, and fluid flowing from the primary sump. The SGs and PRilR remained inactive m:W600\2344w 59.non:lb 100395 5,9,3 2 REVISION: 1
-._._m . _._.__.m._ . - - . _ _ _ . _ . . - . _ . _ _ _ . . . - _ _ . _ . _ _ . - . _ . _ ._ _ ___
throughout this phase of the transient and the principal paths for energy removal from the primary system were via the hot leg break and via the ADS-4 valves. Integrated mass flow from the primary system via the ADS and the break is shown in Figure 5.9.3-43. Significant amounts of flow left the primary circuit through both the ADS-4 valves and the hot leg break during the LTC phase of the transient. This is confirmed by Figures 5.9.3-44 to 5.9.3-45, which show the flows through the ADS and break. Figure 5.9.3-36 shows the calculated steam generation rate as determined by the DVI line flow method. During the IRWST injection phase of the transient, steam was generated at [ ]'** lbm/sec. Steam left the primary circuit by the ADS-4 route, as evidenced by the following:
- The ADS 4 2 vapor flow rate showed periods of significant positive flow. ;
- Examination of the fluid thermocouples on the outlet of the ADS-4 valves indicates that I
temperatures remained at or above saturation. l 1 Furthermore, it was not possible for all the steam generated in the core to flow from the upper head to the downcomer via the bypass holes (Subsection 6.1.3). Therefore steam left the primary system via ADS-4. Figure 5.9.3-50 shows all the components to the system energy balance; any contribution from steam leaving via the ADS-4 valves is not included in this figure. Further discussion of steam O- loss from the primary circuit is provided in the mass and energy balance discussions of Section 6.2. 1 O m:W3h59.noa:tb. loo 395 5.9.3-3 REVISION: 1
l TABLE 5.9.31 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.9.3 LONG TERM TRANSIENT Plot No. Component Variables Units Description l 1 RV RPVPWR kW Core power 2 Primary sump TSMPil, TSMPI2 'F Sump injection line temperatures 3 DVI TDVIL1, TDVIL2 'F DVI line temperatures 4 RPV TOIDC, 'IV2DC, T03DC, 'F Water and saturation temperatures in ST01DC downcomer 5 RPV TOIRPV, 'IU8RPV, 'F Core inlet / outlet temperature, ST08RPV saturation temperature 6 DVI.1 WW'IDVIL1, Ibm /sec. Individual components and total flow WWTIRWII, in DVI l WWTIRW13 7 DVI2 WWIDVIL2, Ibm /sec. Individual components and total flow WWTIRWI2, in DVI-2 WWTIRW14 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT CLDP-509, CLDPSIO in. Level CL-CMT balance lines 10 IRWST IRWST lbm Mass of fluid in IRWST 11 IRWST CLDP-701 in. Collapsed liquid level in IRWST 12 1RWST UIRWST Btu Fluid energy in IRWST 13 Pnmary sump AMPSMP lbm Primary sump fluid mass 14 Primary sump CLDP-901 in. Primary sump level 15 Pnmary sump UPSMP Blu Primary sump fluid energy 16 Secondary sump AMSSMP lbm Secondary sump fluid mass 17 Secondary sump CLDP-902 in. Secondary sump level 18 Secondary sump USSMP Btu Secondary sump fluid energy 19 Pnmary sump WSTSMPET, WWTSMPIT lbm/sec. Pnmary sump steam and liquid injection rate 20 Pnmary sump MISMPII, MISMPI2, Ibm Integrated primary sump and IRWST MISMPIT, MIIRWT flows 21 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG side inlet / outlet MSSGOP1, MSSGOP2 plena 22 Surge line PLM lbm Fluid mass in surge line 23 Surge line CLDP402 in. Collapsed liquid level in surge line 24 Surge line UPSL Btu Fluid energy in surge line 25 RV MWRPV lbm Total fluid mass in reactor vessel 9 mAar6062344w 59.uon:lb-100395 5.9.3-4 REVISION: 1
i TAllLE 5.9.31 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUllSECTION 5.9.3 LONG TERM TRANSIENT Plot No. Component Variables Units Description 26 RV DCM lbm Fluid mass in downcomer 27 RV LDP01DC in. Collapsed liquid level in downcomer compared to various reference elevations 28 RV MWO3RPV lbm Fluid mass in core region 29 RV LDP03RPV in. Collapsed liquid level in core 30 RV RPVAVDF Core exit void fraction 31 RV RPVAQOUT Core exit quality 32 RV LDP06RPV in. Collapsed liquid level in the upper plenum 33 RV MWO8RPV lbm Fluid mass in the upper head 34 RV RPVASL in. Level of Tsat line 35 RV RPVAPab, RPVPWR kW IIcated rod power above and lelow Tsat level and total 36 RV RPVASOUT lbm/sec. Core steam generation rate 37 RV RPVALIN lbm/sec. Calculated core flow 38 RV HTMXRPV, *F Maximum clad temperature, saturation
/fm\ ST08RPV temperature and delta G 39 Hot leg MWHL1, MWHL2 lbm Water mass in hot legs 40 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs m
41 Cold leg CLlWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 42 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 43 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4, and break 44 ADS and break ADS 13TLR, ADS 41TLR, Ibm /sec. Liquid flow out ADS 1-3 and ADS-4 ADS 42TLR 45 ADS and break BRKSTLR lbm/sec. Liquid flow and total flow out of break 46 Mass balance TOTMASS lbm Total system mass inventory 47 Mass balance PRIMMASS, PRIMASS2 lbm Measured pnmary system inventory and valve from mass balances 48 Mass balance MERROR lbm Mass balance error 49 Mass balance MIN, MOUT SRCMASS lbm Integrated mass flow in and out of primary system and source mass 50 Energy balance Various Blu Component of energy balance 51 ADS-4 ADS 41TLR, ADS 42TLR lbm/sec. Oscillations in ADS-4 liquid flow (N 52 Surge line CLDP-602 in. Oscillations in surgeline level 53 RV CIYT-107 psia Oscillations in upper head pressure m:\a@00\2344w-59.non:lb 100395 5.9.3-5 REVISION: 1
TABLE 5.9.31 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.9.3 LONG TERM TRANSIENT Plot No. Component Variables Units Description 54 RV CLDP-ll3 in. Oscillations in upper plenum level 55 RV LDP03RPV in. Oscillations in core level 56 RV LDP01DC in. Oscillations in downcomer level O O m:\a;fiXA2344w 59.non:Ib-100395 5.93-6 REVISION: 1
l i i i i k 4 ! i q 4 THE FIGURES LISTED IN TABLE 5.93-1 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT , l
)
I I 1 1 m:W3h59.non:Ibl00995 5.9.3-7 REVISION: 1
. _ _ . ~ . . . - - - . . - . . - . . _ _ - - - - - - _ _ _
5.10 Analysis of Matrix Test SB19 Matrix Test SB19 (OSU Test U0019) simulated a 2-in, break LOCA with LTC and without the l operation of the nonsafety-related systems. By automatically controlling the BAMS header pressure, I the effect of containment backpressure was siinulated. De break was located at the bottom of CL-3 ; and except for the simulation of backpressure, this test was similar to SB01, including the simulated { failure of one of the ADS-4 lines. Changes to the OSU facility between SB01 and SB19 are identified in the Final Data Report.m The analysis of Matrix Test SB19 is divided into three sections, as follows: I
. Facility performance is discussed in Subsection 5.10.1. It provides a brief outline of the response of the test facility; further details are available in the Final Data Report.*
De short-term transient for SB19 encompassed the start of the simulation up to ( ]** seconds. This period includes blowdown, natural circulation, ADS, and initial IRWST stages of the transient. . For this test, the CMT refill occurred during the short-term transient. At [ ]"# seconds, CMT-1 began to refill, and CMT-2 followed [ ]*# seconds later. This CMT refill phenomena is discussed further in Subsection 6.1.1 and is excluded from this discussion of the short-term transient. ( De analysis of the long-term transient for SB19 encompassed the time frame from ( ]*# seconds to the end of the test, his phase of the transient included IRWST injection and covered the transition to sump injection. The long-term transient actually started with IRWSTinjection, which is discussed as part of the short-term transient. Between the end of the short-term transient and [ ]*# seconds, the system remained relatively inactive, so this discussion begins at [ ]*# seconds. The discussion of the short and long-term phase of the transient focuses on important thermal-hydraulic phenomena identified in the PIRT (Table 1.3-1). The mass and energy balance results are key indicators of the quality of the analysis on which this discussion is based. Dese are discussed in detail in Subsections 6.2.2 and 6.2.3. O mv,omms t.non:ib-too395 5.10-1 REVISION: 1
5.10.1 Facility Performance - The performance of the OSU test facility during Matrix Test SB19 in reference to the five transient phases is outlined in the following:
- Blowdown
- Natural circulation l
- ADS
= IRWST injection
. Sump injection The overall performance of the facility during the transient is shown in Figures 5.10.1-1 to 5.10.1-4.
Figure 5.10.1 1 shows the pressurizer pressure throughout the test with various phases and operating compoaents delineated on the figure. He time scale was reduced for clarity since there were only small changes in system pressure during the long-term phase of the transient. Figure 5.10.1-2 shows the total DVI line flow and its composition from the various sources at each time in the transient. Figure 5.10.1-3 shows the calculated core steam generation rate throughout the test, and Figure 5.10.14 shows the variation in average measured core outlet temperature and peak clad temperature relative to the core outlet saturation temperature, t Figures 5,10.1-1 and 5.10.1-2 show that there was a continuous flow of water to the core from the s passive safety-related systems throughout the transient. Once initiated, the ADS lines rapidly depressurized the primary system, which enhanced the CMT and accumulator injection flow rates. Ultimately, the ADS-4 valves sufficiently reduced the system pressure to start gravity-driven IRWST injection. He passive injection systems overlapped so that as one source of water drained, the next was available to continue the cooling process. The level of steam generation in the core and the response of the average measured core outlet fluid temperatures and maximum clad temperatures are shown in Figures 5.10.1-3 and 5.10.1-4. These figures show that the cooling flow prevented core heatup, and the core remained covered. De core remained subcooled for large periods of the transient and when steam production occurred, the rate of generation remained well below the rate at which water was delivered to the core. 5.10.1.1 Blowdown Pluise The blowdown phase began at time zero when the break was initiated and continued until the primary system pressure was in equilibrium with the secondary-side pressure at around [ ]* seconds. Immediately following the opening of the break, the primary system pressure felt gradually to the end of the blowdown phase. Durirg this phase of the transient, cooling flow was provideri from the two CMTs, which remained in the recirculation mode, and heat was removed from the primary system via the SGs. O m%p600c3h51. wib-too395 5.10.1 1 REVISION: 1
5.10.1.2 Natural Circulation Phase in this LOCA simulation, the single- and two-phase natural circulation phase was initially marked by a (Figure 5.10.1 1) short period of stable system pressure, followed by a gradual reduction after [ ]"* seconds when the primary system pressure fell below that of the secondary side. During this phase of the transient, the SG tubes drained by about [ ]"# seconds and at this time, heat removal from the primary system continued via the PRilR. The pressurizer and surge line completely drained at [ ]*' and [ ]*' seconds, respectively. In response to voiding in CL-3, CMT-1 transitioned to draindown mode at [ ]'b' seconds, and the falling CMT level reached the ADS low-level setpoint at [ ]"# seconds. As shown in Figure 5.10.1-2, a low level of accumulator flow began at [ ]** seconds. Le natural circulation phase of the transient continued to [ ]** seconds when the ADS-1 valve opened. 5.10.1.3 Automatic Depressurization System Phase ADS-1 actuation was followed by ADS-2 and ADS-3 [ ]'** and [ ]"' seconds later. With initiation of the ADS, accumulator injection increased. De influx of cold water combined with increased venting via the ADS led to a rapid depressurization of the primary system. Actuation of ADS-4 at [ ]"# seconds completed depressurization to a level that allowed IRWST injection at [ ]'b' seconds via both DVI lines. During the rapid accumulator injection, increased flow path system resistance reduced flow out of the CMTs. CMT flow resumed as the accumulators drained. The accumulators were fully drained [ j"' seconds before IRWST injection began. De CMTs did not fully drain during the short-term transient. Flow from the CMTs ceased following the start of IRWST injection. He transfer from CMT/ accumulator to IRWST injection was indicated by the minimum RPV inventory of [ ]"# lbm at [ ]'6# seconds. His minimum inventory coincided with the refill of CMT-1. Another slightly higher minimum was observed at [ ]"' seconds when CMT-2 refilled. Actuation of ADS-2 rapidly refilled the pressurizer as water and steam flowed out of the ADS. The pressurizer gradually drained by [ ]'b' seconds. I 5.10.1.4 In Containment Refueling Water Storage Tank Injection l l IRWST injection signals the transition from the short- to long-term phase of the transient he initial phase of IRWST injection involved an increase in flow through the two DVI lines, which was followed by a gradual flow reduction as the driving head between the IRWST and the RCS fell due to I the reduced IRWST water level. Once maximum flow was established, the influx of water from the IRWST was sufficient to keep the core subcooled from [ ]"# seconds (Figure 5.10.1-4). , Sta.m was subsequently generated in the core for the remainder of the transient (Figure 5.10,1-3). I Following the restart of core steam generation, IRWST injection between [ ]"" seconds, was marked by oscillations in pressure and level throughout the primary system. These , oscillations were also observed in the ADS-4 liquid flow rates. mhp600s2%.51.non:1b-ioo395 5.10.1-2 REVIs!ON: 1
l 1 5.10.1.5 Sump Injection - When the IRWST level fell to ( ]'6'in., the main sump injection valves opened and the sump ! injection flow started (Figure 5.10.1-2). Sump flow began at [ J'6' seconds, and the driving head from the sump was sufficient for flow to the IRWST on DVI-1. On DVI 2, there was a I corresponding increase in the flow out of the IRWST, which indicated that there was no significant increase in IRWSTinventory. Note that there was no flow through the check valves around the main sump injection valves before the main valves opened for this test. l 1 l l l l l l l I v 4 ( ( mwac2m-si.noo: Sloo395 5.10.1-3 REVISION: 1
TABLE 5.10.1 1 OSU TEST ANALYSIS PLOT PACKAGE FOR SUBSECTION 5.10.1 Plot No. Component Variables Units Description 1 Pressurizer CPT-604 psia System pressure and event history 2 Water WWTDVil+WWTDVI2, Ibm /sec. Total of CMT, accumulator, IRWST, injection WOUTACCl+WOUTACC2, and sump injection flows WWITRWil+WWTIRWI2, WWTSMPIT 3 Reactor PJ)VASOU2 lbm/sec. Ster.9 generation in reactor vessel vessel 4 Reactor T08RPV, HTMXRPV, TS AT 'F Reactor vessel outlet temperature, vessel maximem clad temperature and fuel exit saturation temperature O l l I O ( i mhpaxn3h51.noa:1b-too395 5.10.1-4 REVISION: 1
_ . _ _ . . ._ . . . _ _ . _ _ _ _ . _ _ _ _ . - _ _ _ _ . _ ~ . _ _ _ _ _ . _ _ _ _ _ .-- - _ _ _ _ . _ _ _ _ _ _ _ _ - . . _ . _ k i i I i I 1 \ !, ) ( , i I i ! i 4 , l, i
- i. THE FIGURES LISTED IN TABLE 5.10.11 e
i ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT T 4 t i i i 4 ( l t 1 1 l 1 1 i s J j i m:W344w-51.non:IM00995 5.10.1-5 REVISION: 1 ! 1 __ _ - ~ . _ . _ _ _ . . . _ _ _ . . _ _ . - . _ , . . _ . . . _ . -. . , _ _ _ . _ . . _ , . .
[ 5.10.2 Short-Term Transient For the 2-in. cold-leg break with simulated containment backpressure, Matrix Test SB19, the short-term transient encompassed the time frame up to [ J'** seconds. As shown in Figure 5.10.1-1, this period included the full depressurization of the facility through all four stages of the ADS, together with CMT and accumulator injection plus the initial stages of IRWST injection. De variations in mass, energy, pressure, and temperature throughout this stage of the transient are illustrated in the plot package outlined in Table 5.10.21. He plots concentrate on the primary system, including the accumulators, CMTs, IRWST, primary sump, and flows from the primary system via the ADS, break, and IRWST overflow. There were two principal parameters to be examined for the short-term transient: Adequate flow must be maintained from the passive systems to the reactor vessel. Adequate flow into the core must be maintained to ensure that decay heat was removed from the simulated fuel rods without a temperature excursion. Rese parameters are addressed in the following discussion. l 5.10.2.1 Maintenance of Core Cooling l 1 Mass Injected to the Primary System ! l l Figures 5.10.2-6 and 5.10.2-7 show the combined effect of the injection flows for the short-term phase l of the transient. Separate plots of the individual contributions to the total flow can be located by consulting the plot package index given in Table 5.10.2-1. Figures 5.10.2-5 and 5.10.2-6 show how the CMTs, accumulators, and IRWST supplied a continuous l flow of water to the core. During the first [ J'A' seconds, cooling flow was provided by the CMTs, which was then supplemented by flow from both accumulators. De rate of flow from the CMTs had an initial value of [ ]'** lbm/sec., which increased to [ ]*** lbm/sec. when draindown commenced. As the driving head fell in response to the CMT water heat-up and draindown, the flow rate gradually decreased until ADS-1 initiation, which resulted in an increase in accumulator flow. Rapid accumulator injection temporarily reduced CMT flow, but led to an overall increase in flow to the core to a peak value of [ ]'*# lbm/sec. Following the end of accumulator injection, the CMTs again provided cooling flow until backpressure from IRWST injection ended the CMT draindown. Since IRWST injection began before the CMTs had fully drained, there was no period of the short-term transient when the passive safety systems failed to provide flow to the RPV. mw6aochst.noa:1b-loo 395 5.10.2-1 REVISION: 1
Reactor Pressure Vessel and Downcomer llehavior ne effect of water flow on the average measured core inlet / outlet temperatures and peak clad temperatures during the short-term phase of the transient is shown in Figures 5.10.2-3 and 5.10.2-57. He core outlet temperature first reached saturation at [ ]'6# seconds and remained at the saturation level until [ ]'6# seconds when the rapid accumulator injection subcooled the core. The combined CMT and accumulator flow was sufficient to keep the core completely subcooled up to [ J' 6" seconds. De core outlet temperature then remained at the saturation level for about [ ]'6# seconds until the influx of water from the IRWST was sufficient to subcool the core again. He core then remained subcooled until the end of the shon-term transient, with the exception of a short period during the reflood of CMT-2. Figure 5.10.2-57 shows that there were no significant excursions in heater rod temperatures throughout the short-term transient; therefore, sufficient core inventory and flow was maintained through this phase of the transient to remove the decay heat generated. For significant portions of the transient, a two-phase mixture was present in the core and upper plenum regions, with core boiling kept at a low level. He following discussion tracks the variation in water level and mass throughout the reactor vessel and downcomer. The mass and level for the core region are shown in Figures 5.10.2-44 and 5.10.2-45. The collapsed liquid level in the core indicates that the heater rods remained covered with a single- or two-phase mixture. He minimum core inventory of [ ]'6# Ibm occurred at [ ]'6# seconds into the transient before accumulator injection began. Figure 5.10.2-45 shows that the collapsed liquid level dropped to [ j'6# in below the top of the heater rod length during this phase of the transient. The average void fraction of the core two-phase mixture may be estimated by dividing the measured l core collapsed liquid level by the [ l'6# in. heated rod length. In this test, the minimum collapsed liquid level corresponded to a core void fraction of [ J.' 6
- During the period before IRWST flow was fully established, the core void fraction reached a maximum of [ ].* 6# By the end of the short-term transient, the effect of IRWST injection ended all core boiling (Figure 5.10.2-55), and the core was again water-solid.
The collapsed liquid level in the upper plenum region covered by LDP-113 and the associated fluid mass are shown in Figures 5.10.2-49 and 5.10.2-48. During the period before accumulator injection, the upper plenum gradually drained to the DVI line elevation. The start of accumulator injection caused an increase in collapsed liquid level to the elevation of the cold legs. Following the end of accumulator injection, the region of the upper plenum spanned by the LDP cell fully drained by [ ]'6# seconds and remained drained until IRWST injection supplied sufficient inventory to initiate a refill. A further draining and refill of the upper plenum coincided with the reflood of Chfr-2. He upper plenum was again water-solid by the end of the short-term transient. Figures 5.10.2-50 and 5.10.2-51 show that the upper head variations were similar to those observed for the upper plenum. He upper head gradually drained before the onset of accumulator injection, which m sapsocc3h51.oon:ib-too395 5.10.2-2 REVISION: 1
=- _. . - - . _ -. -- - . - _ - . . -
l caused an increase in upper head inventory. He upper head was fully drained by [ ]# seconds. IRWST injection was sufficient to resupply a level in the upper head, and at the end of the short-term
; transient, a level below the elevation of the by-pass holes was maintained.
l De mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.10.2-41 and 5.10.2-42. De downcomer level fell to the middle of the cold-leg piping by [ ]'6# seconds, where it remained for [ ]'6# seconds. Before the start of accumulator injection, the downcomer level fell further but remained at or above the middle of the cold legs. As a result, there was a level of water in the cold legs for most of the short-term transient. 5.10.2.2 Energy Transport from the Primary System Following the break, energy was deposited by the heater rods in the primary system fluid to simulate I decay heat and by the primary system metal as it cooled down. Some fluid energy was lost to l ambient and out of the break. Excess energy must be removed from the primary system to prevent excessive fluid and heater rod temperature excursions. The AP600 plant is designed to remove heat i . by a combination of the SGs and the PRHR plus the ADS. ] Steam Generator and Passive Residual Heat Removal Heat Transfer l 1 l
~'T During normal operation, most of the primary system heat was removed via the SGs; however, once . (O the RCPs tripped, the reduced system flow decreased primary- to secondary-side heat transfer. The SGs were only available as heat sinks until the time when the primary system pressure dropped to that of the secondary side; afterward, the two sides were in thermal equilibrium while the primary system had water in the SG tubes. He PRHR is designed to remove heat from the primary system once the I S signal opens the isolation valve. He PRHR continues to remove energy after the SGs were thermally isolated until ADS actuated. Once ADS actuated,it became the predominant path for the i removal of energy from the primary system.
Figure 5.10.2-33 shows the SG primary- and secondary-side pressure together with the PRHR ! integrated heat transfer, as represented by the IRWST fluid energy after alMwing for the contribution from ADS 1-3 inflow. He SGs were potential sinks for primary system heat while the primary-side pressure was above that of the secondary side, that is, before [ ]'6# seconds and before the SG tubes drained. PRHR heat removal began [ ]# seconds into the test. The PRHR was responsible for all the IRWST heat-up until ADS-1 activation, after which the PRHR heat transfer reduced significantly.
, During the active phase, the PRHR transferred heat to the IRWST at an average rate of [ ]'*d Blu/sec.
Energy Transport via the Break and Automatic Depressurization System g The mass flow rate from the primary system via the break is shown in Figures 5.10.2-67 and tj 5.10.2-68. As shown in these figures, steam and liquid flow were detected by the flow measuring m%6m2m -51..on: b-ioo395 5.10.2-3 REVISION: 1
devices for the short-term transient. With initiation of ADS 13, steam flow through the break ceased, although liquid flow continued as a level in the cold legs was maintained. This liquid flow ..as supplemented by steam and liquid flow through the ADS l-3 valves. Between [ ]'** and [ ]'*' seconds, ADS 1-3 caused the system to rapidly depressurize and at [ ]** seconds, ADS-4 was initiated and the primary system continued to the BAMS header pressure. l Flow through the ADS continued until [ ]'** seconds when the flow through the ADS 1-3 terminated and was replaced by flow through the lower resistance ADS-4 paths. By the end of the short-term transient, water was flowing out of the two ADS-4 valves at approximately [ ]'** lbm/sec. (Figure 5.10.2-61). The integrated mass flow from the primary system via the ADS and the break is shown in Figure 5.10.2-62, and the corresponding integrated energy flow is shown in Figure 5.10.2-69. The total system inventory plot given in Figure 5.10.2-70 indicates that [ ]^* lbm ofinventory was gained during the short-term transient. Components of the energy balance are shewn in Figure 5.10.2-74. l 1 l l O
~
G mwnsi.noa:id.ioo395 5.10.2 4 REVIs!ON: 1
_ m.._ _ _ _ ._.. _ . . _ . . . _ - _ _ . _ _ _ _ . _ _ _ , _ . _ _ _ _ . _ _ _ _ _ . . _ _ _ _ _ _ _ _ _ . _ _ _ _ . _ _ . _ _ . _ . . _ _ { W TABLE 5.10.2-1 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.10.2 Plot No. Component Variables Units Description 1 Pressurizer CPT-604 psia System pressure 2 RPV RPVPWR kW Core power 3 RPV TOIRPV, TD8RPV, 'F Core inlet / outlet temperature, ST08RPV saturation temperature 4 SG CPT-201, CPT-204,- psia Primary and secondary pressures in SG CIFT-301 CPT-302 5 DVI-l WWTDVIL1, Ibm /sec. Individual components and total flow in WWTIRWII, DVI-l WOUTACC1, WWTIRWI3 6 DVI-2 WWTDVIL2, Ibm /sec. Individual components and total flow in WWTIRWI2, DVI-2 WOUTACC2, WWTIRWI4 7 CMT AMCMTIB, Ibm Fluid mass in CMTs (excludes balance AMCMT2B lines) 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid levelin CMTs 9 CMT- MIWDVIL1, Ibm Integrated mass out of CMTs MlWDVIL2 O* 10 CMT WWTDVIL1, Ibm /sec. Flow out of CMTs
%WIDVIL2 11 CMT WOUTCLBI, Ibm /sec. Flow into CMTs WOUTCLB2 12 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 13 CMT UCMT1, UCMT2 Btu Fluid energy in CMTs 14 IRWST IRWST lbm Mass of fluid in IRWST 15 IRWST CLDP-701 in. Collapsed liquid level in IRWST 16 IRWST WWTIRWII, Ibm /sec. Flow from IRWST t DVI line WWTIRWI2
~
17 IRWST IRWSTOR lbm/sec. Overflow from IRWST to sump' 18 IRWST ADS 13TMR lbm/sec. Total ADS flow into IRWST 19 IRWST ADS 13TIR, MIIRWII, Ibm Integrated mass out of IRWST MIIRWI2, MIIRWIO 20 IRWST UIRWST Btu Fluid energy in IRWST 21 PRHR CLDP-802 in. Collapsed liquid level in PRHR HX O mW3h51.non:Ibl00395 5,]0.2-5 REVISION: 1
i TABLE 5.10.2-1 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECflON 5.10.2 Plot No. Component Variables Units Description 22 PRlut WWOTPRHR Ibm /sec. Measured outlet flow from PRHR tube 23 Accumulator AMACCI, AMACC2 lbm Mass of fluid in accumulators 24 Accumulator CLDP-401. CLDP-402 in. Collapsed liquid level in accumulators 25 Accumulator WOUTACCl, IbtrJsec. Flow from accumulators WOUTACC2 26 Accumulator MOUTACCl, Ibm Integrated mass out of accumulators MOUTACC2 27 Accumulator UACCl, UACC2 Btu Fluid energy in accumulators 28 Pnmary sump AMPSMP lbm Prunary sump fluid mass 29 Pnmary sump CLDP-901 in. Pnmary sump level 30 Pnmary sump UPSMP Btu Primary sump fluid energy 31 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG primary side MSSGOP1, MSSGOP2 inlet / outlet plena 32 SG MSSGliT1, MSSGirl'2, Ibm Mass of fluid in SG primary side bot MSSGCT1, MSSGCT2 and cold tubes 33 SG/PRHR CI'r-201, CI'r-301, psia & SGI pressure and PRHR integrated beat ; QPRHR1 Btu output 1 34 Pressurizer PZM lbm Fluid mass in pressurizer 35 Pressurizer CLDP-601 in. Collapsed liquid level in pressurizer 36 Pressurizer UPZ Btu Fluid energy in pressurizer 37 Surge line PLM lbm Fluid mass in surge line j 38 Surge line CLDP-602 in. Collapsed liquid level in surge line 39 Surge line UPSL Btu Fluid energy in surge line 40 RPV MWRPV lbm Total fluid mass in reactor vessel 41 RPV DCM lbm Fluid mass in downcomer 42 RPV LDP01DC in. Collapsed liquid level in downcomer compared to various reference elevations 43 RPV MWOIRPV lbm Fluid mass in lower plenum 44 RPV MWO3RPV lbm Fluid mass in core region 45 RPV LDP03RPV in. Collapsed liquid level in core 46 RPV RPVAVDF2 Core exit void fraction 47 RPV RPVAQOU2 Core exit quality 48 RPV MWO6RPV lbm Fluid mass in the upper plenum 49 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 50 RPV MWO8RPV lbm Fluid mass in the upper bead 51 RPV LDP08RPV in. Collapsed liquid level in the upper head mp344w.51.non:1b ioo395 5.10.2-6 REVISION: 1
CJ 1 BLe ,.. 2., <Co.t . d, OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECFION 5.10.2 - Plot No. Component Variables Units Description 52 RPV URPV Btu Total fluid energy in reactor vessel 53 RPV RPVXE, RPVASL2 in. Level of Tsat line 54 RPV RPVPab, RPVAPab2, kW Heated rod power above and below RPVPWR Tsat level and total 55 RPV RPVRXV,RPVASOU2 lbm/sec. Core steam generation rate 56 RPV RPVALIN2 lbm/sec. Calculated core flow 57 RPV HTMXRPV, S11)8RPV 'F Maximum clad temperature and saturation temperature 38 Hot leg MWHL1, MWHL2 lbm Water mass in bot legs 59 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs 60 Cold leg CLIWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 61 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 62 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4s, and break 63 ADS and break BRKTIVF, AD13TIVF, Ibm Totalintegrated vapor flow for ADS AD41TIVF, AD42TIVF and break 64 ADS and break BRKTILF, AD13TILF, Ibm Total integrated liquid flow for ADS AD41TILF, AD42TILF and break 65 ADS and break ADS 13SVR, Ibm /sec. Vapor flow out ADS 1-3 and ADS-4 ADS 41SVR, ADS 42SVR 66 ADS and break ADS 13SLR, Ibm /sec. Liquid flow out ADS 1-3 and ADS-4 ADS 41SLR, ADS 42SLR 67 ADS and break BRKSSVR lbm/sec. Vapor flow out of break 68 ADS and break BRKSSLR lbm/sec. Liquid flow out of break 69 ADS and break BRKSPEI, ADS 13EI, Btu Integrated fluid energy for ADS 1-3, ADS 41EI, ADS 42E1 ADS-4, and break 70 Mass balance TOTMASS lbm Total system mass inventory 71 Mass balance PRIMMASS, Ibm Measured prunary system inventory and PRIMASS2 value from mass balance 72 Mass balance MERROR Ibm Mass balance error 73 Mass balance MIN, MOUT lbm Integrated mass flow in and out of SRCMASS prunary system and source mass 74 Energy balance Various Btu Components of energy balance O { maap6tn2h51.non:1b.100395 5.10.2-7 REVISION: 1
l O 4 THE FIGURES LISTED IN TABLE 5.10.2-1 , ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT 1 i i l O O m:W344w-51. mon:n.ioo995 5.10.2-8 REVISION: 1
O () 5.10.3 Long Term Transient - De long-term transient started with initiation of IRWST injection, covered the transition from IRWST to sump injection, and provided information on the LTC response of the AP600 plant. For the 2-in. cold-leg break with simulated containment backpressure, Matrix Test SB19, the long-term transient analyzed ran from [ ]'6* seconds to the end of the test around [ ]'** seconds. The behavior of the test facility during this period of the transient is discussed in this subsection using the plot package detailed in Table 5.10.3-1. His analysis concentrates on the components of the primary system that remained active during the L'IU phase, that is, the RPV, the hot legs, ADS-4, the sumps, and the IRWST. For the long-term transient, thermal-hydraulic phenomena of interest were:
- Maintenance of core cooling and removal of energy from the primary system.
Level oscillations (from [ j'*# seconds there were system wide level and pressure oscillations, which are discussed further in Subsection 6.1.3). 5.10.3.1 Maintenance of Core Cooling O V Mass Injected into Primary System Total DVI line flow, CMT flow, and IRWST flows are shown in Figures 5.10.3-6 and 5.10.3-7, and the flow from the primary sump is shown in Figure 5.10.3-19. From around [ ]# seconds, there was a contribution to the DVI flow from the CMTs as the CMTs reached post-refill draindown. During the pre-sump injection phase of the transient, IRWST flow proceeded at a gradually declining rate with the effect of the primary system oscillations superimposed. At [ ]'6* seconds, the main sump injection valves opened, resulting in a reversal of flow through IRWST injection line-1 and an increase in IRWST flow in line-2. He net result was that an injection flow rate of [ ]'*# lbm/sec. was maintained through DVI-1 and -2, respectively. Reactor Pressure Vessel and Downcomer Response ne effect of the water inflow on the average measured downcomer fluid temperatures, core inlet and core outlet temperatures, and heater rod temperatures during the long-term phase of the transient is shown in Figures 5.10.3-4,5.10.3-5, and 5.10.3-38. Figure 5.10.3-4 shows that there was a general increase in average downcomer fluid temperatures during the long-term transient. By the end of the test, this average temperature was [ J'*# *F below saturation. Figure 5.10.3-5 shows that the core was above saturation temperature for most of the long-term transient. Figures 5.10.3-34 to 5.10.3-36 Q show that the DVI line flow method discussed in Section 4.11 indicates that a small level of core mwooc3m-51.non:ib.too395 5.10.3-1 REVISION: 1
boiling was maintained after [ J'b' seconds into the transient. Nevertheless, the level of boiling was small and showed that the inflow from the IRWST and sumps was sufficient to maintain cooling. Figure 5.10.3-38 shows that there were no significant excursions in heater rod temperatures throughout the long-term transient; therefore, sufficient core inventory and flow was maintained through this phase of the transient to remove the decay heat generated. For significant portions of the transient, a two-phase mixture was present in the core and upper plenum regions. l The following discussion tracks the variation in water level and mass throughout the reactor vessel and downcomer. He mass and level for the core region are shown in Figures 5.10.3-28 and 5.10.3-29. De collapsed liquid level in the core indicated that the heater rods, were always covered with a single- or two-phase mixture. During the later stages of the transient, the collapsed liquid level fell just below the top of the heater rods, and the maximum core void fraction was [ ].*** De fall in core inventory was a result of the influx of hot water from the primary sump as it flowed through the check valves. De impact of this hot water on the system temperatures is shco .n Figures 5.10.3-4 and 5.10.3-5 as a sudden increase in fluid temperature in the downcomer and at the core inlet. The hot water also led to an increase in the calculated steam generation rate, as shown in Figure 5.10.3-36, and a corresponding fall in the level at which the core reached saturation temperature (Figure 5.10.3-34). De collapsed liquid level in the upper plenum region covered by LDP-ll3, is shown in Figure 5.10.3-32. The figures indicate that during the period before sump injection began, the collapsed liquid level initially fell and then remained at the top of the hot legs. Following the influx of relatively hot water from the sumps, the level dropped to the bottom of the hot legs, where it remained for the rest of the transient. His level corresponded to a void fraction of [ ] .**# Re mass of water in the reactor pressure vessel is shown in Figure 5.10.3-25. After an initial decline, the reactor vessel water mass settled at an average value of [ ]'*' lbm until the time sump injection started when it fell to [ ]'6# lbm, which is [ ]'6# pm M h Mal vel we bate From [ ]'** seconds, oscillations in vessel inventory were observed. Figures 5.10.3-51 to 5.10.3-56 illustrate these oscillations using plots on a restricted time frame from [
]'6* seconds. Dese oscillations are observed in primary system measurements from the upper plenum to the ADS-4 flows. The oscillations occurred for [ ]'6# seconds. The oscillations in the ADS flow lagged behind those in the upper head pressure by around [ ]'*# seconds. Dese oscillations and possible mechanisms for their production are discussed further in Subsection 6.1.3.
The mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.10.3-26 and 5.10.3-27. De collapsed liquid level remained above the mid-elevation of the cold legs until [ ]'6' seconds. At this time the level fell below the cold legs and subsequently, the start of sump injection reduced the level to the DVI line elevation before it recovered to the hot legs. He downcomer level remained at this level for the rest of the transient. m:\ap600C344w-51.non:1b-100395 5.10.3-2 REVISION: 1
i 5.10.3.2 Energy Transport from the Primary System During the long-term transient, energy centinueri to be deposited in the primary system from the heated rods, metal, and fluid flowing from the primary sump. The SGs and PRHR remained inactive throughout this phase of the transient, and the principle path for energy out of the primary system was via the ADS 4 valves. Integrated mass flow from the primary system via the ADS and the break is shown in Figure 5.10.3-43. During the LTC pliase of the transient until nearly [ }*# seconds, there is significant outflow through the break and ADS-4 valves. This is confirmed by Figures 5.10.3 44 to 5.10.3-45, which show flow through the ADS and break. At [ ]*# seconds, flow through the break ceased as the cold legs drained. From [ ]'6# seconds, there is reverse flow through the break as indicated by the reducing integral given in Figure 5.10.3-43. During the sump injection phase of the transient, measured outflow is in the form of liquid out of the ADS-4 valves. Water flowed through each of these valves at an average rate of [ ]'6# lbm/sec. Figure 5.10.3-36 shows the calculated steam generation rate, as determined by the DVI line flow method. During the sump injection phase of the transient, steam was generated at over [ ]'6# lbm/sec., although the steam vortex meters indicate little or no flow out of the ADS-4 valves. The following two indications show that steam is leaving the primary system by this route: Figure 5.10.3-46 shows total measured system fluid inventory. During this phase of the transient after the start of primary sump injection, that is, when core steam generation was most significant, the total system inventory fell by about [ ]*# lbm. 'Ihis amount corresponds to a steam flow rate of f ]*# lbm/sec., which would not have been detected by the vortex meters. Examination of the fluid thermocouples on the outlet of the ADS-4 valves indicates that temperatures remained at or above saturation temperature following the start of sump injection. It was not possible for all the steam generated in the core to flow from the upper head to the downcomer via the bypass holes (Subsection 6.1.3). Therefore, steam was leaving the primary system via ADS-4. Figure 5.10.3-50 shows all the components to the system energy balance. Further discussion of steam loss from the primary system is provided in the mass and energy balance discussions of Section 6.2. l O manp600chs t.noc:Ib 100995 5.10.3-3 REVISION: 1
I l l l l TABLE 5.10.31 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.10 3 l LONG TERM TRANSIENT i Plot No. Component Variables Units Description 1 RPV RPVPWR kW Core power l 2 Pnmary sump TSMPII, TSMPI2 'F Sump injection line temperatures 1 1 3 DVI TDVIL1, TDVIL2 *F DVIline temperatures ' 4 RPV T01DC, T02DC, T03DC, 'F Water and saturation temperatures m l ST01DC downcomer l 5 RPV TOIRPV,108RPV, 'F Core inlet / outlet temperature, ST08RPV saturation temperature 6 DVI-l WWTDVIL1, Ibm /sec. Individual components and total flow WWTIRWil, in DVI-l WWTIRWI3 7 DVI-2 WWTDVIL2, Ibm /sec. Individual components and total flow WWTIRWI2, in DVI-2 WWTIRW14 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs i 9 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 10 IRWST IRWST lbm Mass of fluid in IRWST 11 IRWST CLDP-701 in. Collapsed liquid level in IRWST 12 IRWST UIRWST Btu Fluid energy in IRWST 13 Pnmary sump AMPSMP lbtn Pnmary sump fluid mass 14 Pnmary sump CLDP-901 in. Pnmary sump level 15 Pnmary sump UPSMP Btu Pnmary sump fluid energy 16 Secondary sump AMSSMP lbm Secondary sump fluid mass 17 Secondary sump CLDP-902 in. Secondary sump level 18 Secondary sump USSMP Btu Secondary sump fluid energy 19 Pnmary sump WSTSMPET, WWTSMPIT lbm/sec. Primary sump steam and liquid injection rate 20 Primary sump MISMPII, MISMPI2, Ibm Integrated primary sump and IRWST MISMPIT, MIIRWT flows 21 SG MSSGIPI, MSSGIP2, Ibm Mass of fluid in SG side inlet / outlet MSSGOP1, MSSGOP2 plena 22 Surge line PLM lbm Fluid mass in surge line 23 Surge line CLDP-602 in. Collapsed liquid level in surge line 24 Surge line UPSL Blu Fluid energy in surge line 25 RPV MWRPV lbm Total fluid mass in reactor vessel e m: sap 6ao2344w-51.noo:isioo395 5.10.3-4 REVISION: 1
i 1 i t i TABLE 5.10.3-1 (Continued) 5 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.10.3 i LONG TERM TRANSIENT Plot No. Component Variables Units Description i b 26 RPV DCM lbm Fluid mass in downcomer l 27 RPV LDP01DC - in. Collapsed liquid levelin downcomer ! cour. pared to various reference , elevations 28 RPV MWO3RPV lbm Fluid mass in core region l 29 RPV LDP03RPV in. Collapsed liquid levelin core 1 30 RPV RPVAVDF2 Core exit void fraction
- 31 RPV RPVAQOU2 Core exit quality
> 32 RPV LDP06RPV in. Collapsed liquid level in the upper f plenum 33 RPV MWO8RPV lbm Fluid mass in the upper head l 34 RPV RPVASL2 in. Level of Tsat line l 35 RPV RPVAPab2, RPVPWR kW Heated rod power above and below l Tsat level and total
- 36 RPV RPVASOU2 lbm/sec. Core steam generation rate
! 37 RPV RPVALIN2 lbm/sec. Calculated core flow j 38 RPV HTMXRPV, 'F Maximum clad temperature, saturation ST08RPV temperature and delta 39 Hot leg MWHL1, MWHL2 lbm Water mass in hot legs j~ 40 Hot leg MVHL1, MVHL2 lbm Vapor mass in hot legs i 41 Cold leg CLlWMS, CL2WMS, Ibm Water mass in cold legs j CL3WMS, CL4WMS 3 42 Cold leg CLIVMS, CL2VMS, Ibm - Vapor mass in cold legs
, CL3VMS, CL4VMS l 43 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 13, ADS 41TIR. ADS 42T1R ADS-4, and break 44 ADS and break ADS 13TLR, ADS 41TLR, Ibm /sec. Liquid flow out ADS 1-3 and ADS-4 ADS 42TLR 45 ADS and break BRKSTLR lbm/sec. Liquid flow and total flow out of break 46 Mass balance TO1 MASS lbm Total system mass inventory 47 Mass balance PRIMMASS, PRIMASS2 lbm Measured primary system inventory and valve from mass balances 48 Mass balance MERROR lbm Mass balance error 49 Mass balance MIN, MOUT SRCMASS lbm Integrated mass flow in and out of primary system and souret mass 50 Energy balance Various Btu Component of energy balance 51 ADS-4 ADS 41TLR, ADS 42TLR lbm/sec. Oscillations in ADS-4 liquid flow p 52 Surge line CLDP-602 in. Oscillations in surgeline level 53 RPV Cirr-107 psia Oscillations in upper head pressure mAng600(2344w-51.non:Ib-10o395 5.10.3-5 REVISION: 1
TABLE 5.10.31 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.10.3 LONG TERM TRANSIENT Plot No. Component Variables Units Description 54 RPV CLDP-ll3 in. Oscillations in upper plenum level 55 RPV LDP03RPV in. Oscillations in core level 56 RPV LDP01DC in. Oscillations in downcomer level e O O m:\ap60cc344.-51.non:Ib-ioo395 5.10.3-6 REVISION: 1
i I i I d ) I l 1 N . 4 l 1
- l l
1 i THE FIGURES LISTED IN TABLE 5.10.31 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT J- ; i J i I i 4 J k 4 m:Wm51.non:thloo995 5.10.3-7 REVISION: 1
- , . . . . . , - . ., . . . - . -- . . . - . . - - . . . - . . - . ~ . . - .
4 ) 5.11 Analysis of Matrix Test SB21 - i Matrix Test SB21 (OSU Test U0021) simulated a double 4-in, cold-leg break LOCA with LTC and l without the operation of the nonsafety-related systems, ne breaks were located on the top and . bottom of CL-3 and except for the break size, this test was similar to SB01, including the simulated
- failure of one of the ADS-4 lines. Changes to the OSU facility since the performance of SB01 are I noted in the Final Data Report.m ;
.l De analysis of Matrix Test SB21 is divided into three sections as follows: l l
i
- Re facility performance is discussed in Subsection 5.11.1, which provides a brief outline of
{ the response of the test facility; further details are available in the Final Data Report.* l j I The shon-term transient for SB21 encompassed the stan of the simulation up to l [ ]*# seconds. His period included the blowdown, natural circulation, ADS, and initial 4 IRWST stages of the transient. 3 )
- He analysis of the long-term transient for SB21 encompassed the time frame from
[ ]*d seconds to the end of the test. His phase of the transient includes IRWST injection , and covered the transition to sump injection. He long-term transient actually started with ! IRWST injection, which is discussed as part of the short-term transient. Between the end of i' the short-term transient and [ ]*# seconds, the system remained relatively inactive with the l exception of the CMT refill. At [ ]** seconds, CMT-1 began to refill, and CMT-2 followed [ - ]*' seconds later. CMT refill phenomena is discussed further in i Subsection 6.1.1, and the discussion of the long-term transient provided here begins at [ ]*' seconds. The discussion of the short- and long-term phase of the transient focus, . on important thermal-hydraulic phenomena identified in the PIRT (Table 1.31). The mass anu energy balance results are j key indicators of the quality of the analysis on which this discussion is based. Dese are discussed in detail in Subsections 6.2.2 and 6.2.3. 1 j 1 4 I l 4
! b 'V l
rn:\ap60m2344w-51.non:lb 100395 5,11 1 REVISION: 1
5.11.1 Facility Performance - He performance of the OSU test facility during Matrix Test SB21 in reference to the five transient phases is outlined in the following :
- Blowdown
- Natural circulation
- ADS
- IRWST injection
- Sump injection The overall performance of the facility during the transient is shown in Figures 5.11.1-1 to 5.11.1-4.
Figure 5.11.1 1 shows the pressurizer pressure throughout the test with various phases and operating components delineated on the figure. The time scale was reduced for clarity since there were only small changes in system pressure during the long-term phase of the transient. Figure 5.11.12 shows the total DVI line flow and its composition from the various sources at each time in the transient. Figure 5.11.1-3 shows the calculated core steam generation rate throughout the test, and Figure 5.11.1-4 shows the variation in average measured core outlet temperature and peak clad temperature relative to the core outlet saturation temperature. O \/ Figures 5.11.1-1 and 5.11.12 show that there was a continuous flow of water to the core from the passive safety-related systems throughout the transient. Once initiated, the ADS lines rapidly depressurized the primary system, which enhanced the CMT and accumulator injection flow rates. Ultimately, the ADS-4 valves sufficiently reduced the system pressure to start gravity-driven IRWST injection. The passive injection systems overlapped so that as one source of water drained, the next l was available to continue the cooling process. The level of steam generation in the core and the response of the average measured core outlet fluid temperatures and maximum clad temperatures are shown in Figures 5.11.1-3 and 5.11.1-4. These figures show that the cooling flow prevented core heatup and the core remained covered. The core remained subcooled for large periods of the transient and when steam production occurred, the rate of generation remained well below the rate at which water was delivered to the core. 5.11.1.1 Blowdown Phase I The blowdown phase began at time zero when the break was initiated and continued until the primary system pressure was in equilibrium with the secondary-side pressure at about [ ]* seconds. During this phase of the transient, cooling flow was provided from the two CMTs, which remained in O mhp600(2mw-51.noo:1b-too395 5.11.1-1 REVISION: 1
recirculation. Heat was removed from the primary system via the SGs. He pressurizer and surge line completely drained at [ ]'b' and [ ]'6# seconds, respectively. 5.11.1.2 Natural Circulation Phase After a brief period of stability, the single- and two-phase natural circulation phase was marked by a (Figure 5.11.1 1) gradual reduction in system pressure in this LOCA simulation, rather than by the more stable pressure observed in SB01. During this phase of the transient, the SG tubes drained by about [ ]'6# seconds and at this time, heat removal from the primary system continued via the PRHR. In response to voiding in CL-3, CMT-1 transitioned to draindown mode at [ ]'** seconds before the end of the blowdown phase, and the falling CMT level reached the ADS low-level setpoint at [ ]'6# seconds. CMT-2 recirculation flow ceased at [ ]'6# seconds, which was after CMT-1, but the ADS low level setpoint was reached before CMT-1 at [ ]'6' seconds. At [ ]'6# seconds, the accumulators began (Figure 5.11.1-2) to inject at a steady rate of over [ ]'6
- lWsm without affecting the CMT outflow. De natural circulation draindown phase of the transient continued to
[ ]'*# seconds when the ADS-1 valve opened. The minimum RPV inventory was estimated as [ ]'6# lbm at around [ ]'*# seconds. This is an overestimate of the actual mass present by up to [ l'*# ibm because LDP413 spanning the upper plenum was inoperable during SB21. 5.11.1.3 Automatic Depressurization System Phase ADS 1 actuation was followed by ADS-2 and ADS-3 [ ]'6# and [ ]'6' seconds later. With initiation of the ADS, a rapid phase of accumulator injection began. The influx of cold water combined with increased venting via the ADS led to a rapid depressurization of the primary system. Actuation of ADS-4 at [ ]'*# seconds completed depressurization to a level that allowed IRWST injection at [ ]'6* seconds via DVI-2 and [ ]'6# seconds via DVI-1. During the rapid accumulator injection, increased flow path resistance reduced flow out of the CMTs. CMT flow resumed as the accumulators drained. The accumulators were fully drained [ ]'*# seconds before IRWST injection began. Because of the relative elevation heads, CMT flow essentially ceased [ ]'*# seconds after the start of IRWST injection, before the CMTs were fully drained. Actuation of ADS-2 rapidly refilled the pressurizer as water and steam flowed out of the ADS. The pressurizer gradually drained by [ ]'6* ceconds. Note that following the actuation of ADS-1 in Test SB21, a series of irregular oscillations discussed in the Final Data Reportm were observed in pressure, level, flow, and temperature readings. These continued until around [ ]'*# seconds when the CMT-1 balance line was refilling and were believed to have resulted from continued low level irregular flow through the SG tubes. O mAar60'A2344w-51.non:1b-100395 5.11.1-2 REVISION: 1
. . ._. . ~ ~ _ . . . . . --- . . - . -
O I 5.11.1.4 In-Containment Refueling Water Storage Tank Injection IRWST injection signals the transition from the short to long-term phase of the transient. The initial phase of IRWST injection involved an increase in flow through the two DVI lines, which was followed by a gradual flow reduction as the driving head between the IRWST and the RCS fell due to the reduced IRWST water level. Once maximum flow was established, the influx of water from the IRWST was sufficient to keep the core subcooled until ( J*# seconds (Figure 5.11.1-4). Steam was subsequently (Figure 5.11.1-3) generated in the core for the remainder of the transient. The period of IRWST injection between [ ]*** seconds was marked by oscillations in pressure and level throughout the primary system. These oscillations were also observed in the ADS-4 liquid flow rates. 5.11.1.5 Sump Injection When the IRWST level fell to [ ]'b# in., the main sump injection valves opened and the sump injection flow started (Figure 5.11.1-2). Sump flow began at [ ]'6# seconds and the driving head from the sump was sufficient for flow to the IRWST on DVI-1. On DVI-2. there was a corresponding increase in the flow out of the IRWST, which indicated that there was no significant increase in IRWST inventory until later in the transient. Note that there was no flow through the check valves around the main sump injection valve before those valves opened for this test. l 0 l 1 i l (O/ maap600cus.-5i.no :it>.ioo395 5.I1.1-3 REVISION: 1
TABLE 5.11.1-1 OSU TEST ANALYSIS PLOT PACKAGE FOR SUBSECTION 5.11.1 Plot No. Component Variables Units Description 1 Pressurizer CIFT-604 psia System pressure and event history 2 W ater WWIDVil+WWTDV12, Ibm /sec. Total of CMT, accumulator, IRWST, injection WOUTACCl+WOUTACC2, and sump injection flows WWTIRWI1+WWTIRW12, WWTSMPIT 3 Reactor RPVASOU2 lbm/sec. Steam generation in reactor vessel vessel 4 Reactor T08RPV, HTMXRPV, TS AT 'F Reactor vessel outlet temperature, vessel maximum clad temperature and fuel exit saturation temperature O O 1 l I m:\ap600c344w-51.noa:ib-too395 5.11.1 4 REVISION: 1 (
.... _ . - _ . . ~. . .... . . - . - --.--- .. - -. - - - - .- - - - -. - -. - -.. - - .. - - . ..-... - .-.
1 s l M i i i 1 l 1 4 1 \ J 1 f 1 i i; THE FIGURES LISTED IN TABLE 5.11.1-1 f ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT d 1 e I 4 i i ,
- i t
i !' i ! l k 4 e i s l i 1 i l ) i i 1 i l l ! l 5 i i i ' \ i i 4 4 4' f 4 4 1 l i
- i. A 4
1-\ i i . l l m:\np600ch51.non:1b too995 5.11.1-5 REVISION: I d
, _,y - w - -
l lt 5.11.2 ' Short-Term Transient
\
For the largest cold-leg break LOCA simulation, Matrix Test SB21, the short-term transient
- encompassed the time frame up to [ ]'6' md M hwn in Figure 5.11.1 1, this period l included the full depressurization of the facility through all four stages of the ADS, together with CMT and accumulator injection plus the initial stages of IRWST injection. De variations in mass.
j energy, pressure, and temperature throughout this stage of the transient are illustrated in the plot J package outlined in Table 5.11.2-1. The plots concentrate on the primary system, including the
)
accumulators, CMTs, IRWST, primary sump, and flows from the primary system via the ADS, break, i
- and IRWST overflow.
i j Rere were two principal parameters to be examined for the short-term transient: .
- Adequate flow must be maintained from the passive systems to the reactor vessel.
!
- Adequate flow into the core must be maintained to ensure that decay heat was removed from j the simulated fuel rods without a temperature excursion.
2 Dese parameters are addressed in the following discussion. 5.11.2.1 Maintenance of Core Cooling j Mass Injected into the Primary System i l Figures 5.11.2-6 and 5.11.2-7 show the combined effect of the injection flows for the short-term phase I of the transient. Separate plots of the individual contributions to the total flow can be located by
- consulting the plot package index given in Table 5.11.2-1.
- Figures 5.11.2-5 and 5.11.2-6 show how the CMTs, accumulators, and IRWST supply a continuous
- j. flow of water to the core. During the first [ ]'6* seconds, cooling flow was provided by the CMTs, i which was then supplemented by flow from both accumulators. The rate of flow from the CMTs
, gradually reduced from an initial value of ( ]*** lbm/sec. as the driving head fell in response to the CMT water heat-up and draindown until ( ]'6' seconds when the depressurization following ADS-1 i- initiation generated rapid accumulator injection. De rapid accumulator flow temporarily stopped ! CMT flow, but led to an overall increase in flow to the core to a peak value of ( ]*** lbm/sec.
- Following the end of accumulator injection, the CMTs again provided cooling flow until the flow from IRWST injection ended CMT draindown. Since IRWST injection began before the CMTs had fully
} drained, there was no period of the short-term transient when the passive safety systems failed to
< provide flow to the RPV.
1 iO 4 maap60m23u 51.noa:1b-ioo395 5.11.2-1 REVISION: 1 4
l Reactor Pressure Vessel and Downcomer Behavior l l The effect of the water flow on the average measured core inlet / outlet temperatures and peak clad temperatures during the short-term phase of the transient is shown in Figures 5.11.2-3 and 5.11.2-57. De core outlet temperature reached the saturation point at [ J'6' seconds and remained at the i saturation level until the end of the short-term transient. l l Figure 5.11.2-57 shows that there were no significant excursions in heater rod temperatures throughout l l the short-term transient; therefore, sufficient core inventory and flow was maintained through this phase of the transient to remove the decay heat generated. For significant portions of the transient, a two-phase mixture was present in the core and upper plenum regions, with core boiling kept at a low level. , l The following discussion tracks the variation in water level and mass throughout the reactor vessel and downcomer. The mass and level for the core region are shown in Figures 5.11.2-44 and 5.11.2-45. l The collapsed liquid level in the core indicates that the heater rods remained covered with a single- or j two-phase mixture throughout the short-term transient. He minimum core inventory of [ J'6 ' lbm occurred at [ ]'6* seconds into the transient before accumulator injection began. Figure 5.11.2-45 shows that the collar.cd liquid level dropped to [ ]'6* in. below the top of the heated rod length during this phase of the transient. De average void fraction of the core two-phase mixture may be estimated by dividing the measured core collapsed liquid level by the [ ]'6* In. heated rod length. In this test, the minimum collapsed liquid level corresponded to a core void fraction of [ ] .'6
- By the end of the short-term transient, the effect of IRWST injection ended all core boiling (Figure 5.11.2-55), and the core was again water-solid.
1 he collapsed liquid level in the upper plenum region and the associated fluid mass are shown in Figures 5.11.2-49 and 5.11.2-48. He LDP spanning the upper plenum (LDP-ll3) was inoperable for all of the transients and indicated that the upper plenum remained full of water. His resulted in an overestimation of the minimum vessel water inventory of up to [ ]'6* lbm. The precise error depends on how full the upper plenum was at the time. l l Figures 5.11.2-50 and 5.11.2-51 show that the upper head drained rapidly during SB21. De upper head remained essentially drained for the entire short-term transient. He mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.11.2-41 and 5.11.2-42. The downcomer collapsed liquid level fell to the bottom of the cold-leg piping during the first [ ]'6* seconds. IRWST injection maintained the collapsed liquid level at the center of the cold legs. Cold-leg refill began at around [ ]'6* seconds. Following ADS-1 actuation, there were irregular oscillations in the downcomer level, which probably resulted from continued flow through the
- SG tubes.
O mup6cou3w.51.oon:1b-ioo395 5.I1.2-2 REVis]ON: I
5.11.2.2 Energy Transport from the Primary System - Following the break, energy was deposited by the heater rods in the primary circuit fluid to simulate decay heat and by the primary circuit metal as it cooled down. Some fluid energy was lost to the ambient and out of the break. The energy must be removed from the primary system to prevent excessive fluid and heater rod temperature excursions. The AP600 plant is designed to remove heat by a combination of the SGs and the PRHR plus the ADS. Steam Generator and Passive Residual Heat Removal Heat Transfer During normal operation, most of the primary system heat was removed via the SGs; however, once the RCPs tripped, the reduced system flow decreased primary- to secondary-side heat transfer. The SGs were only available as heat sinks until the primary system pressure dropped to that of the secondary side; afterward, the two sides were in thermal equilibrium. The PRHR is designed to remove heat from the primary system once the S signal opens the isolation valve. He PRHR continues to remove energy after the SGs are thermally isolated until ADS actuates. Once ADS actuates, it becomes the predominant path for the removal of energy from the primary system. Figure 5.11.2-33 shows the SG primary- and secondary-side pressure together with the PRHR integrated heat transfer, as represented by the IRWST fluid energy after allowing for the contribution O from ADS 1-3 inflow. The SGs were potential sinks for primary system heat, while the primary-side pressure was above that of the secondary side, that is, before [ ]** seconds. PRHR heat removal began [ ] seconds into the test, and the PRHR was responsible for all the IRWST heat-up until ADS-1 activation. Following actuation of the ADS, PRHR heat transfer continued for [ ]*' seconds. This was different from the behavior observed in the smaller break cases, as discussed in Subsection 7.1.2. During the active phase, the PRHR transferred heat to the IRWST at an average rate of[ ] Btu /sec. Energy Transport via the Break and Automatic Depressurization System The mass flow rate from the primary system via the break is shown in Figures 5.11.2-67 and 5.11.2-68. During the first [ J seconds following the break, nearly [ ]*# lbm of water appeared to flow out of the primary system via the break (Figure 5.11.2-62). Figure 5.11.2-71 shows that there was no evidence of such a dramatic loss of primary system inventory from the mass balance. He apparent mass flow resulted from the indicated increase in break separator inventory, caused by a rise in pressure above the fluid in response to the opening of the break. He first [ ]** seconds should taerefore be ignored in reference to mass loss from the break. At [ ]"# seconds, the operator (as instructed) isolated the 8-in. vent line as the volumetric steam flow rate dropped below the [ ]*# cfm level. His caused a small increase in break separator pressure, which increased the steam and liquid flow at this time. ) i m:\ap600(2344w-51.non:1b-100395 5.11.2-3 REVISION: 1
1 l l Following the initiation of ADS 1-3, flow through the break continued and was augmented by steam and liquid flow through the ADS 1-3 valves. Between [ ]'6" and [ ]*' seconds, ADS 1-3 l l rapidly depressurized the system and at [ ]*# seconds, ADS-4 was initiated and the primary system continued to depressurize to the HAMS header pressure. Beyond [ ]*# seconds, there was continued flow through the break as cold-leg refill was occurring. i Flow through the ADS continued at a declining rate until [ ]'6* seconds when flow through l ADS 1-3 terminated and was replaced by flow through the lower resistance ADS 4 paths for the rest of the short-term transient. Integrated mass flow from the primary system via the ADS and the break is shown in Figure 5.11.2-62, and the corresponding integrated energy flow is shown in Figure 5.11.2-69. The j total system inventory plot given in Figure 5.11.2-70 indicates that only [ ]*# lbm of inventory left { the system during the short-term transient. Components of the energy balance are shown in i Figure 5.11.2-74. 1 Ol1 1 O mAapm2ws1.non:t5100395 5.I1.2 4 REVISION: I
O TABLE 5.11.21 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.11.2 Plot No. Component Variables Units Description i Pressurizer CI'T-604 psia System pressure 2 RPV RPVPWR kW Core power 3 RPV TOIRPV, T08RPV, *F Core inlet / outlet temperature, ST08RPV saturation temperature 4 SG Cirr-201, CPT-204, psia Primary and secondary pressures in SG CI'T-301, CPT-302 5 DVI-1 WWTDVILI, Ibm /sec. Individual components and total flow in l WWTIRWII, DVI l ! WOUTACC1, , WWTIRW13 6 DVI-2 WWTDVIL2, Ibm /sec. Individual components and total flow in WWTIRW12, DVI-2 WOUTACC2, WWTIRWI4 7 CMT AMCMTIB, Ibm Fluid mass in CMTs (excludes balance AMCMT2B lines) 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT MIWDVIL1, Ibm Integrated mass out of CMTs MIWDVIL2 ) w/ 10 CMT WWTDVIL1, Ibm /sec. Flow out of CMTs WWTDVIL2 11 CMT WOUTCLBI, Ibm /sec. Flow into CMTs WOUTCLB2 12 CMT CLDP-509, CLDPSIO in. Level Cl CMT balance lines 13 CMT UCMT1, UCMT2 Btu Fluid energy in CMTs 14 IRWST IRWST lbm Mass of fluid in IRWST 15 IRWST CLDP-701 in. Collapsed liquid level in IRWST 16 IRWST WWTIRWII, Ibm /sec. Flow from IRWST to DVI lines WWTIRWI2 17 IRWST IRWSTOR lbm/sec. Overflow from IRWST to sump 18 IRWST ADS 13TMR Ibm /sec. Total ADS flow into IRWST 19 IRWST ADS 13TIR, MIIRWII, Ibm lategrated mass out of IRWST MllRWI2, MIIRWIO 20 IRWST UIRWST Btu Fluid energy in IRWST 21 PRHR CLDP-802 in. Collapsed liquid levelin PRHR HX
/ i %.)
m:\ap600cw51.nn:1b.100395 5.I1.2-5 REVISION: 1
l TABLE 5.11.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 511.2 Plot No. Component Variables Units Description 22 PRHR WWOTPRHR Ibm /sec. Measured outlet flow from PRHR tube 23 Accumulator AMACCl, AMACC2 lbm Mass of fluid in accumulators 24 Accumulator CLDP-401, CLDP-402 in. Collapsed liquid level in accumulators 25 Accumulator WOUTACC1, Ibm /sec. Flow from accumulators WOUTACC2 26 Accumulator MOlJTACCl, Ibm Integrated mass out of accumulators MOUTACC2 27 Accumulator UACC1, UACC2 Btu Fluid energy in accumulators 28 Prunary sump AMPSMP lbm Pnmary sump fluid mass 29 Pnmary sump CLDP-901 in. Primary sump level 30 Prtmary sump UPSMP Btu Pnmary sump fluid energy 31 SG MSSGIP1, MSSGIP2. Ibm Mass of fluid in SG primary side MSSGOP1, MSSGOP2 iulet/ outlet plena 32 SG MSSGHT1, MSSGHT2, Ibm Mass of fluid in SG pnmary side hot MSSGCTI, MSSGCT2 and cold tubes 33 SG/PRHR CFT-201, CPT-301, psia & SGI pressure and PRHR integrated beat QPRHRI Btu output 34 Pressurizer PZM lbm Fluid mass in pressurizer 35 Pressunzer CLDP-601 in. Collapsed liquid level in pressurizer 36 Pressurizer UPZ Blu Fluid energy in pressurizer 37 Surge line PLM lbm Fluid mass in surge line 38 Surge line CLDP-602 in. Collapsed liquid level in surge line 39 Surge line UPSL Btu Fluid energy in surge line 40 RPV MWRPV lbm Total fluid mass in reactor vessel 41 RPV DCM lbm Fluid mass in downcomer 42 RPV LDP0lDC in. Collapsed liquid level in downcomer compared to various reference elevations 43 RPV MW0lRPV lbm Fluid mass in lower plenum 44 RPV MWO3RPV lbm Fluid mass in core region 45 RPV LDP03RPV in. Collapsed liquid levelin core 46 RPV RPVAVDF2 Core exit void fraction 47 RPV RPVAQOU2 Core exit quality 48 RPV MWO6RPV lbm Fluid mass in the upper plenum 49 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 50 RPV MWO8RPV lbm Fluid mass in the upper head 51 RPV LDP08RPV in. Collapsed liquid level in the upper head m:wp60&2h51.noa:1b-too395 5.11.2-6 REVISION: 1
4 4 4 l
/'
TABLE 5.11.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.11.2 Plot No. Component Variables Units Description 52 RPV URPV Btu Total fluid energy in reactor vessel 53 RPV RPVXE, RPVASL2 ft Level of Tsat line 54 RPV- RPVPab, RPVAPab2, kW Heated rod power above and below RPVPWR Tsat level and total 55 RPV RPVRXV, RPVASOU2 lbm/sec. Core steam generation rate 56 RPV RPVALIN2 lbm/sec. Calculated core flow 57 RPV HTMXRPV, S'II)8RPV 'F Maximum clad temperature and saturation temperature 58 Hot leg MWHL1, MWHL2 lbm Water mus in hot legs 59 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs 60 Cold leg CLIWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 61 Cold leg CL1VMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 62 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4s, and break 63 ADS and break BRKTIVF, ADl3TIVF, Ibm Totalintegrated vapor flow for ADS O 64 ADS and break AD41TIVF, AIM 2TIVF BRKTILF, ADl3TILF, AD41TILF, AD42TILF' Ibm . and break Total integrated liquid flow for ADS and break 65 ADS and break ADS 13SVR, Ibm /sec. Vapor flow out ADS 1-3 and ADS 4 ADS 41SVR, ADS 42SVR 66 ADS and break ADS 13SLR, Ibm /sec. Liquid flow out ADS 1-3 and ADS-4 ADS 41SLR, ADS 42SLR 67 ADS and break BRKSSVR lbm/sec. Vapor flow out of break 68 ADS and break BRKSSLR lbm/sec. Liquid flow out of break 69 ADS and break BRKSPEl, ADS 13EI, Btu integrated fluid energy for ADS 1-3, ADS 41El, ADS 42El ADS-4, and break 70 Mass balance TOTMASS lbm Total system mass inventory 71 Mass balance PRIMMASS, Ibm Measured primary system inventory and PRIMASS2 value from mass balance 72 Mass balance MERROR Ibm Mass balance error l 73 Mass balance MIN, MOUT lbm Integrated mass flow in and out of SRCMASS primary system and source mass 74 Energy balance Various Btu Components of energy balance l O m: sap 600s2m51..o :1b.too395 5.11.2-7 REVISION: 1
l l 1 l 9 THE FIGURES LISTED IN TABLE 5.11.2-1 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT l l l l l O' O m:w3uw.51.non:is100995 5.I1.2-8 REVISION: 1
l 1 j i 5.11.3 Long Term Transient - The long-term transient started with initiation of IRWST injection, covered the transition from IRWST l to sump injection, and provided information on the LTC response of the AP600 plant. For the large cold-leg break, Matrix Test SB21, the long-term transient analyzed runs from [ ]'** seconds to the , end of the test around [ ]'6* seconds. The behavior of the test facility during this period of the i transient is discussed in this subsection using the plot package detailed in Table 5.11.3-1. This } analysis concentrates on the components of the primary system that remained active during the LTC phase, that is, the RPV, the hot legs, ADS-4, the sumps, and the IRWST. ; I During the long-term transient, the main thermal-hydraulic phenomena of interest were: l Maintenance of core cooling and removal of energy from the primary system. I 1
- Level oscillations (from [ ]'** seconds there were system wide level and pressure oscillations, which are discussed further in Subsection 6.1.3).
f ) 5.11.3.1 Maintenance of Core Cooling j Mass Injected into Primary System 1 (n) b
\
Total DVI line flow, CMT flow, and IRWST flows are shown in Figures 5.11.3-6 and 5.11.3-7, and
- the flow from the primary sump is shown in Figure 5.11.3-19. From around [ J'6' i
- seconds, there was a contribution to the DVI flow from the CMTs as the CMTs reached post-refill
! draindown. l During the pre-sump injection phase of the transient, IRWST flow proceeded at a gradually declining rate with the effect of the primary system oscillations superimposed. At [ ]'6* seconds, flow 1 from the primary sump began through the main injection valves, which opened as the IRWST reached ' { tv low-low level setpoint. This resulted in a reversal of flow through IRWST injection line-1 and an ' inci.Ae in IRWST flow in line-2. The net result was that an injection flow rate of [ ]'6
- lbm/sec.
was maintained through both DVI lines. Reactor Pressure Vessel and Downcomer Response The effect of the water inflow on the downcomer fluid temperatures, core inlet and core outlet I temperatures, and heater rod temperatures during the long-term phase of the transient is shown in Figures 5.11.3-4,5.11.3-5, and 5.11.3-38. Figure 5.11.3-4 shows that there was a general increase in average downcomer fluid temperatures during the long-term transient. By the end of the test, this average temperature reached an equilibrium [ ]a,b,c *F below saturation. Figure 5.11.3-5 shows that g) ( the core outlet temperature remained at or above saturation for most of the long-term transient. Figures 5.11.3-34 to 5.11.3-36 show that the DVI line flow method described in Section 4.11 indicates maarmocm-51. on:1b.too395 .5.11,3-1 REVISION: 1
I that a small level of boiling was maintained after [ ]'6' seconds into the transient. Nesertheless, l the level of boiling was small and showed that the inflow from the IRWST and sumps was sufficient to maintain cooling. Figure 5.11.3-38 shows that there were no significant excursions in heater rod temperatures throughout l the long-term transient; therefore, sufficient core inventory and flow was maintained through this phase i of the transient to remove the decay heat generated. For significant portions of the transient, a two-phase mixture was present in the core and upper plenum regions. De followitt discussion tracks the variation in water level and mass throughout the reactor vessel and downcomer. De mass and level for the core region are shown in Figures 5.11.3-28 and 5.11.3-29. He collapsed liquid level in the core indicated that the heater rods were always covered with a single-or two-phu;e mixture. During the later stages of the transient (beyond [ ]'6# seconds), the collapsed liquid level remained just below the top of the heater rods, and the core void fraction was [ ] .' 6
- hm was a reduction in the core collapsed liquid level following the start of sump injection as the DVI flow settled at a new value. There is, however, no impact on primary system fluid temperatures in this test since the sump water was relatively cold (Figures 5.11.3-4 and 5.11.3-5).
During sump injection, the calculated steam generation rate was at a maximum of about [ l'6' lbm/sec. (Figure 5.11.3-3). The collapsed liquid level in the upper plenum region is shown in Figure 5.11.3 32. For this test, the LDP for the upper plenum (LDP-113) was inoperable; thus, variations in upper plenum mass and level were not tracked. Figure 5.11.3-33 shows the mass of water in the upper head, which remained below [ ]'6'lbm until the end of the test. The mass of water in the reactor pressure vessel is shown in Figure 5.11.3-25. For the entire long- j term transient, the reactor vessel water mass remained at an average value of [ ]'6# lbm, which is [ ]'6# percent of the initial vessel water inventory. From [ ]' 6d seconds, oscillations in vessel inventory were observed. Figures 5.11.3-51 to 5.11.3-56 illustrate these oscillations using plots on a restricted time frame from [ ]'6* seconds. These oscillations are observed in primary system measurements from the upper plenum to the ADS-4 flows. De oscillations in the ADS flow lagged behind those in the upper head pressure. Ecse oscillations and possible mechanisms for their production are discussed further in Subsection 6.1.3. He mass of fluid and collapsed liquid level in the RPV downcomer is shown in Figures 5.11.3-26 and 5.11.3-27. He collapsed liquid level remained above the cold legs until [ ]'6# seconds, then decreased to the center of the hot legs. His occurred at about the same time CMT-1 completed draindown and the cold legs drained (see Figure 5.10.3-41). Here was no effect on downcomer level resulting from the start of sump injection. O m:\ap60cc3h5i.noo:1w00395 5.11.3-2 REVISION: 1
M
! 5.11.3.2 Energy Transport from the Primary System -
i During the long-term transient, energy continued to be deposited in the primary system from the l heated rods, metal, and fluid flowing from the primary sump. The SGs and PRHR remained inactive l throughout this phase of the transient, and the primary path for energy out of the primary system was l via the ADS-4 valves. l Integrated mass flow from the primary system via the ADS and the break is shown in Figure 5.11.3-43. During the LTC phase of the transient, the only significant outflow la through the ADS-4 valves and, until [ ]'6' seconds, the break. This is confirmed in Figures 5.11.3-44 and 5.11.3-45, which show flow through the ADS and break. After [ ]'6* seconds, there was reverse l flow through the break, as indicated by the reducing integrated flow shown in Figure 5.11.3-43. l During the sump injection phase of the transient, the only measured flow was liquid out of the ADS-4 valves. By the end of the test, water flowed through these valves at a combined average rate of [ ]"# lbm/sec. Figure 5.11.3 36 shows the calculated steam generation rate, as determined by the DVI line flow l method. During the sump injection phase of the transient, steam was generated around [ ]' 6 *
^
lbrn/sec., although the steam vortex meters indicate little or no flow out of the ADS-4 valves. The following two indications show that steam is leaving the primary system by this route: p , i (
- Figure 5.11.3-46 shows total measured system fluid inventory. During this phase of the transient after the start of primary sump injection (from [ ]"# seconds), that is, when core steam generation was most significant, the total system inventory fell by about l
[ ]"# lbm. This amount corresponds to a steam flow rate of [ ]"# Ibm /sec., which ) would not have been detected by the vortex meters. I
- Examination of the fluid thermocouples on the outlet of the ADS-4 valves indicates that temperatures remained at or above saturation temperature following the start of sump injection.
It was not possible for all the steam generated in the core to flow from the upper head to the downcomer via the bypass holes (Subsection 6.1.3). It can therefore be concluded that steam was leaving the primary system via ADS-4. Figure 5.11.3-50 shows all the components to the system energy balance. Further discussion of steam loss from the primary system is provided in the mass and , energy balance discussions of Section 6.2. I l l O ma.psoccmst.non:tb.100995 5.I1.3-3 REVISION: 1
TABLE 5.11.31 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.11.3 LONG-TERM TRANSIENT Plot No. Component Variables Units Description 1 RPV RPVPWR kW Core power 2 Pnmary sump TSMPII, TSMPI2 'F Sump injection line temperatures 3 DVI TDVILI, TDVIL2 'F DVI line temperatures 4 RPV T01DC,102DC, T03DC, *F Water and saturation temperatures in ST0lDC downcomer 5 RPV TOIRPV,TD8RPV, 'F Core inlet / outlet temperature, STOSRPV saturation temperature 6 DVl-1 WWTDVIL1, Ibm /sec. Individual components and total flow WWTIRWII, in DVI-l WWTIRWI3 7 D VI-2 WWTDVIL2, Ibm /sec. Individual components and total flow WWTIRWI2, in DVI-2 W%TIRWI4 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 10 1RWST IRWST lbm Mass of fluid in IRWST 11 1RWST CLDP-701 in. Collapsed liquid level in IRWST l 12 IRWST UIRWST Btu Fluid energy in IRWST ! 13 Pnmary sump AMPSMP lbm Prunary sump fluid mass 14 Prunary sump CLDP-901 in. Pnmary sump level l 15 Pnmary sump UPSMP Btu Pnmary sump fluid energy 16 Secondary sump AMSSMP lbm Secondary sump fluid mass h 17 Secondary sump CLDP-902 in. Secondary sump level 18 Secondary sump USSMP Btu Secondary sump fluid energy 19 Pnmary sump WSTSMPET, WWTSMPIT Ibm /sec. Pnmary sump steam and liquid injecGon rate 20 Pnmary sump MISMPII, MISMPI2, Ibm Integrated primary sump and IRWST MISMPIT, MIIRWT flows 21 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG side inlet / outlet MSSGOPI, MSSGOP2 plena 22 Surge line PLM lbm Fluid mass in surge line 23 Surge line CLDP-602 in. Collapsed liquid level in surge line 24 Surge line UPSL Btu Fluid energy in surge line 25 RPV MWRPV lbm Total fluid mass in reactor vessel 9 mangotA2344w 51.non:Ib-100395 5.11.3-4 REVISION: 1
l TABLE 5.11.31 (Continued) l OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.11.3 ; LONG-TERM TRANSIENT l Plot No. Component Variables Units Description 26 RPV DCM lbm Fluid mass in downcomer , 27 RPV LDP0lDC in. Collapsed liquid level in do'icomer l compared to various reference l elevations l 28 RPV MWO3RPV lbm Fluid mass in core region i 29 RPV LDP03RPV in. Collapsed liquid level in core 30 RPV RPVAVDF2 Core exit void fraction 31 RPV RPVAQOU2 Core exit quality 32 RPV LDPOVPV in. Collapsed liquid level in the upper plenum 33 RPV MWO8hPV lbm Fluid mass in the upper head 34 RPV RPVASL2 in. Level of Tsat line 35 RPV RPVAPab2, RPVPWR kW Heated rod power above and below Tsat level and total 36 RPV RPVASOU2 lbm/sec. Core steam generation rate 37 RPV RPVALIN2 lbm/sec. Calculated core flow I l 38 RPV HTMXRPV, 'F Maximum clad temperature, saturation [s\ STT)8RPV temperature and delta V 39 Hot leg MWHL1, MWHL2 lbm Water mass in bot legs 40 Hot leg MVHL1, MVHL2 lbm Vapor mass in bot legs 41 Cold leg CLIWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 42 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 43 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4, and break 44 ADS and break ADS 13TLR, ADS 41TLR. Ibm /sec. Liquid flow out ADS 1-3 and ADS-4 ADS 42TLR 45 ADS and break BRKSTLR lbm/sec. Liquid flow and total flow out of break 46 Mass balance TOTMASS lbm Total system mass inventory 47 Mass balance PRIMMASS, PRIMASS2 lbm Measured prunary system inventory and valve from mass balances 48 Mass balance MERROR lbm Mass balance error 49 Mass balance MIN, MOUT SRCMASS lbm Integrated mass flow in and out of pnmary system and source mass 50 Energy balance Various Btu Component of energy balance 51 ADS-4 ADS 41TLR, ADS 42TLR lbm/sec. Oscillations in ADS-4 liquid flow A 52 Surge line CLDP-602 in. Oscillations in surgeline level V 53 RPV CI'T-107 psia Oscillations in upper bead pressure m:wm2mst.non:1b-ioo395 5.11.3-5 REVISION: 1
TABLE 5.11.31 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.11.3 LONG TERM TRANSIENT Plot No. Component Variables Units Description 54 RPV CLDP-ll3 in. Oscillations in upper plenum level 55 RPV LDP03RPV in. Oscillations in core level 56 ,RPV LDP01DC in. Oscillations in downcomer level I l l l l l l l I e i i O m:\ psu23w.5i. on:15100395 5.I1.3-6 REVISION: 1
(.,..--.---.-......~.-......---- - l THE FIGURES LISTED IN TABLE 5.11.3-1 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT j a ] l T l J J m:W2344w.51.non:Ibl00995 5.11.3-7 REVISION: 1 4
- r. . . 3- --. - - - --- v,.-e- - . - .~.---w,. . - - . . - , , . . - - - ~ - ..w- --- , - - - --- ..-v,-, we.-,- m.w--, . , - - - - 7rm' r --
5.12 Analysis of Matrix Test SB23 - Matrix Test SB23 (OSU Test U0023) simulated an 0.5-in. break LOCA with LTC and without the operation of the nonsafety-related systems. The break was located at the bottom of CL-3 and except for the break size, this test was id:ntical to SB01, including the simulated failure of one of the ADS-4 lines. As noted in Section 1.5, th: original scaling methodology used indicates that the selected OSU break area was larger than necess.try for a true simulation of a 0.5-in. break. However, the results are acceptable to validate codes since the calculations account for this variation when predicting test results. The analysis of Matrix Test SB23 is divided into three sections, as follows: )
- The facility performance is discussed in Subsection 5.12.1. It provides a brief outline of the ' l response of the test facility; further details are available in the Final Data Report.* j
- The short-term transient for SB23 encompassed the start of the simulation up to
[ ]'6# seconds. This period includes the blowdown, natural circulation,' ADS, and initial IRWST stages of the transient.
- The analysis of the long-term transient for SB23 encompassed the time frame from
[ ]'6# seconds to the end of the test. This phase of the transient includes IRWST injection ! and covered the transition to sump injection. The long-term transient actually started with l IRWST injection, which is discussed as part of the short-term transient. Between the end of l the short-term transient and [ J'6# seconds, the system remained relatively inactive with the exception of the CMT refill. At [ J'6# seconds, CMT-1 began to refill, and CMT-2 followed [ ]*6* seconds later. CMT refill phenomena is discussed further it. Subsection 6.1.1, so the discussion of the long-term transient presented begins at [ ]'6* suod The discussion of the short and long-term phase of the transient focuses on important thermal-hydraulic phenomena identified in the PIRT (Table 1.3-1). The mass and energy balance results are key indicators of the quality of the analysis on which this discussion is based. These are discussed in detail in Subsections 6.2.2 and 6.2.3. O v m:\ p60m23w.51.non:tb.too395 5.12-1 REVIs!ON: 1
4 ! i l i I f~ ( 5.12.1 Facility Performance he performance of the OSU test facility during Matrix Test SB23 in reference to the five transient j phases is outlined in the following:
- Blowdown
- Natural circulation -
i
- ADS l
- IRWST injection
- Sump injection l
1 The overall performance of the facility during the transient is shcwn in Figures 5.12.1-1 to 5.12.1-4. Figure 5.12.1 1 shows the pressurizer pressure throughout the test with the various phases and ) operating components delineated on the figure. He time scale was reduced for clarity since there were only small changes in system pressure during the long-term phase of the transient. Figure 5.12.1-2 shows the total DVI line flow and its composition from the various sources at each time in the transient. Figure 5.12.1-3 shows the calculated core steam generation rate throughout the test, and Figure 5.12.1-4 shows the variation in average measured core outlet temperature and peak clad l temperature relative to the core outlet saturation temperature. Figures 5.12.1-1 and 5.12.1-2 show that there was an almost continuous flow of water to the core from the passive safety-related systems throughout the transient. ,Once initiated, the ADS lines rapidly depressurized the primary system, which enhanced the CMT and accumulator injection flow rates. Ultimately, the ADS-4 valves sufficiently reduced the system pressure to start gravity-driven IRWST injection. He passive injection systems overlapped so that as one source of water drained, the next was available to continue the cooling process. The level of steam generation in the core and the response of the average measured core outlet fluid temperatures and maximum clad temperatures are shown in Figures 5.12.1-3 and 5.12.1-4. These figures show that the cooling flow prevented core heatup, and the core remained covered. He core remained subcooled for large periods of the transient and when steam production occurred, the rate of generation remained well below the rate at which water was delivered to the core. 5.12.1.1 Blowdown Phase he blowdown phase began at time zero when the break was initiated and continued until the primary system pressure was in equilibrium with the secondary-side pressure at around [ ]'6 m ods (Figure 5.12.1-1). Immediately following the opening of the break, the primary system pressure increased since the small break was incapable of removing the energy being added from the core O mW60m* 51.non:1b.ioo395 5.12.1-1 REVISION: 1
heater rods. This caused the pressure relief valves to open. After the valves closed, the system pressure fell gradually almost to the end of the blowdown phase. During this phase of the transient, cooling flow was provided from the two ChfTs, which remained in the recirculation mode, and heat was removed from the primary system via the SGs. The pressurizer and surge line completely drained at [ ]'6" and [ ]'6' seconds respectively. 5.12.1.2 Natural Circulation Phase In this LOCA simuladon, the single- and two-phase natural circulation phase was marked by a gradual reduction in system pressure rather than by the more stable pressure observed in SB01. During this phase of the transient, the SG tubes drained by about [ ] seconds and at this time, heat removal from the primary system continued via the PRHR. De steam in the SG tubes became superheated and remained so until the end of the transient. In response to voiding in CL-3, ChfT-1 transitioned to draindown mode at [ J'6' seconds, and the falling Chit level reached the ADS low-level setpoint at [ ]'6' seconds. The natural circulation phase of the transient continued to [ ]'6# seconds when the ADS-1 valve opened. 5.12.1.3 Automatic Depressurization System Phase ADS-1 actuation was followed by ADS-2 and ADS-3 [ ]'6' and [ ]'6# seconds later. With the initiation of the ADS, accumulator injection began (Figure 5.12.1-2). 'Ihe influx of cold water combined with increased venting via the ADS led to a rapid depressurization of the primary system. Actuation of ADS-4 at [ ]'6# seconds completed depressurization to a level that allowed IRWST $ injection at [ ]'6" seconds via DVI-2 and [ ]'6# seconds via DVI-1. During accumulator injection, increased system flow path resistance reduced flow out of the ChfTs. ChfT flow resumed as the accumulators drained. The accumulators were fully drained [ ]'6' seconds before IRWST injection began. The ChfTs did not fully drain until [ ]*6# and [ ]'6# seconds after the start of IRWST injection. The transfer from Chit / accumulator to IRWST injection was indicated by the minimum RPV inventory of [ ]'6# lbm at [ ]'6# seconds. Actuation of ADS-1 rapidly refilled the pressurizer as water and steam flowed out of the ADS. The pressurizer gradually drained by [ ]'6* seconds. 5.12.1.4 In-Containment Refueling Water Storage Injection IRWST injection signals the transition from the short- to long-term phase of the transient. The initial phase of IRWST injection involved an increase in flow through the two DVI lines, which was followed by a gradual flow reduction as the driving head between the IRWST and the RCS fell due to O mMp60m2344w 51.non:lb-100395 5.12.1-2 REVISION: 1
( g the reduced IRWST water level. Once maximum flow was established, the influx of water from the IRWST was sufficient to keep the core subcooled from [ ]'6" seconds (Figure 5.12.1-4). Steam was subsequently generated in the core for the remainder of the transient (Figure 5.12.13). Following the restart of core steam generation, IRWST injection between [ ] seconds, was marked by oscillations in pressure and level throughout the primary system. 'Diese oscillations were also observed in the ADS-4 liquid flow rates. 5.12.1.5 Sump Injection Injection from the primary sump via the check valves around the main sump injection valves began at [ ] seconds when the level in the IRWST was low enough to allow flow. This reduced the flow rate from the IRWST. When the IRWST level fell to [ ]'6* in., the main sump injection valves opened and the sump injection flow rate increased (Figure 5.12.1-2). This increase occurred at [ ]'6" seconds and the driving head from the sump was sufficient for flow to the IRWST on DVI-1. On DVI-2, there was a corresponding increase in the flow out of the IRWST, which meant that there vias only a small initial increase in IRWST inventory. O I l l l m:W%1w-51.non:lb-100395 5.12.1-3 REVISION: 1
TABLE 5.12.1 1 OSU TEST ANAINSIS PLOT PACKAGE FOR SUBSECTION 5.12.1 Plot No. Component Variables Units Description 1 Pressurizer Cirr-604 psia System pressure and event history 2 Water WWTDVil+WWIDV12, Ibm /sec. Total of CMT, accumulator, IRWST, injection WOUTACC1+WOUTACC2, and sump injection flows WWTIRWII+WWTIRWI2, f WWTSMPIT 3 Reactor RPVASOU2 lbm/sec. Steam generation in reactor vessel vessel 4 Reactor T08RPV, HTMXRPV, TSAT 'F Reactor vessel outlet temperature, vessel Inaximum cla:1 temperature and fuel exit saturation temperature O O mp344w 51.noo:tb-too395 5.12.1 4 REVISION: 1
,.~~-.. - - . . _ . , . . - - - - - . - - - . . - . . . - . . - - . - . - . -.
. _ - . ~ . - .
O THE FIGURES LISTED IN TABLE 5.12.11 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT O O mA p6aMh-51.non:lb 100995 5.12.1-5 REVISION: 1
t" i 5.12.2 Short Term Transient - For the 0.5-in. cold-leg break, Matrix Test SB23, the short term transient encompassed the time frame up to [ ]**# seconds. As shown in Figure 5.12.1-1, this period included the full depressurization of the facility through all four stages of the ADS, together with CMT and accumulator injection plus the initial stages of IRWST injection. The variations in mass, energy, pressure, and temperature throughout this stage of the transient are illustrated in the plot package outlined in Table 5.12.2-1. The plots concentrate on the primary system, including the accumulators, CMTs, IRWST, primary sump, and flows from the primary system via the ADS, break, and IRWST overflow. For the short-term transient there were two principal parameters to examine: Adequate flow must be maintained from the passive systems to the reactor vessel. Adequate flow into the core must be maintained to ensure that decay heat was removed from the simulated fuel rods without a temperature excursion. a These parameters are addressed in the following discussion. 5.12.2.1 Maintenance of Core Cooling Mass Injected to the Primary System Figures 5.12.2-6 and 5.12.2-7 show the combined effect of the injection flows for the short-term phase of the transient. Separate plots of the individual contributions to the total flow can be located by consulting the plot package index given in Table 5.12.21. Note that the flow measurements for the outflow from ACC-1 were incorrect for this test. Due to the similarity in the level behavior for the two accumulators, (Figure 5.12.2-24) ACC-2 outflows have been used for ACC-1 in the mass balance calculations and in the flow figures. The level calculations confirm the flow rates. Figures 5.12.2-5 and 5.12.2-6 show how the CMTs, accumulators, and IRWST combined to supply an almost continuous flow of water to the core. During the first [ J'6' seconds, cooling flow was provided by the CMTs. The rate of flow from the CMTs gradually reduced from an initial value of [ ]' lbm/sec. as the driving head fell in response to the CMT water heat-up and draindown until ADS-1 initiation, which resulted in en increase in CMT flow. Rapid accumulator injection temporarily reduced CMT flow, but led to an overall increase in flow to the core to a peak value of [ ]# lbm/sec. Following the end of accumulator injection, the CMTs again provided cooling flow until they drained. The only period in which there was relatively little cooling flow was for [ ] seconds between CMT draining and the start of IRWST injection. mangooamst.noastoo395 5.12.2-1 REVISION: 1
Reactor Pressure Vessel and Downcomer liehavior He effect of water flow on the average measured core inlet / outlet temperatures and peak clad temperatures during the short-term phase of the transient is shown in Figures 5.12.2-3 and 5.12.2-57. The combined CMT and accumulator flow was sufficient to keep the core completely subcooled up to [ ]** seconds. De core outlet temperature then remained at the saturation level for about [ ]*' seconds until the influx of water from the IRWST was sufficient to subcool the core again. i The core remained subcooled until the end of the short-term transient. l Figure 5.12.2-57 shows that there were no significant excursions in heater rod temperatures throughout the short-term transient; therefore, sufficient core inventory and flow was maintained through this phase of the transient to remove the decay heat generated. For significant portions of the transient, a two-phase mixture was present in the core and upper plenum regions, with core boiling kept at a low level. l De following discussion tracks the variation in water level and mass throughout the reactor vessel and downcomer. He mass and level for the core region are shown in Figures 5.12.2-44 and 5.12.2-45. The collapsed liquid level in the core indicated that the heater rods remained covered with a single- or two-phase mixture throughout the shcrt-term transient. He minimum core inventory of [ ]** lbm occurred at [ ]** seconds into the transient before IRWST injection was fully established. Figure 5.12.2-45 shows that the collapsed liquid level dropped to [ ]** in. below the top of the heated rod length during this phase of the transient. He average void fraction of the core two-phase mixture may be estimated by dividing the measured core collapsed liquid level by the [ ]** in. heated rod length. In this test, the minimum collapsed liquiu level corresponded to a core void fraction of [ ).** By the end of the short-term transient, the effect of IRWST injection ended all core boiling (Figure 5.12.2-55), and the core was again water-solid. The collapsed liquid level in the upper plenum region covered by LDP-113 and the associated fluid mass are shown in Figures 5.12.2-49 and 5.12.2-48. During the period before accumulator injection, the upper plenum was water-solid. The start of accumulator injection coincided with a fall in collapsed liquid level to the elevation of the cold legs. Following the end of accumulator injection, the region of the upper plenum spanned by the LDP cell fully drained and remained drained until IRWST injection supplied sufficient inventory to initiate a refill. The upper plenum was again water-solid by the end of the short-term transient. Figures 5.12.2-50 and 5.12.2-51 show that the upper head also lost inventory at the time when accumulator injection began. His behavior is the opposite of what might be expected as a result of the influx of cooler accumulator water. Accumulator injection coincided with ADS-1 initiation. The flow of water and steam through the ADS rapidly refilled the pressurizer, which removed water from the RPV. The reduction in upper plenum and upper head inventory reduced the overall RPV inventory during accumulator injection (Figure 5.12.2-40) even though the core inventory increased (Figure 5.12.2-44). mAap600\2344w 51.non:Ib.100395 5.12.2-2 REVis!ON: 1
He mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.12.2-41 l and 5.12.2-42. The downcomer remained water-solid until ADS-1 actuation when the level fell to the bottom of the cold-leg piping, where it remained until IRWST injection once again raised the level above the cold legs, and cold leg refill began. 5.12.2.2 Energy Transport from the Primary System l l Following the break, energy was deposited by the heater rods in the primary system fluid to simulate l decay heat and by the primary system metal as it cooled down. Some fluid energy was lost to ambient and out of the break. Excess energy must be removed from the primary system to prevent ! excessive fluid and heater rod temperature excursions. De AP600 plant is designed to remove heat i by a combination of the SGs and the PRHR plus the ADS. i Steam Generator and Passive Residual Heat Removal Heat Transfer During normal operation, mos' of the primary system heat was removed via the SGs; however, once l the RCPs tripped, the reduced sys r.1 flow decreased primary- to secondary-side heat transfer. He SGs were only available as heat sinks entil the time when the primary system pressure dropped to that of the secondary side; afterward, the two sides were in thermal equilibrium. De PRHR is designed to remove heat from the primary system once the S signal opens the isolation valve. De PRHR continues to remove energy after the SGs are thermally isolated until ADS actuates. Once ADS actuates, it becomes the predominant path for the removal of energy from the primary system. Figure 5.12.2 33 shows the SG primary- and secondary-side pressure together with the PRHR integrated heat transfer, as represented by the IRWST fluid energy after allowing for the contribution from ADS 1-3 inflow. The SGs were potential sinks for primary system heat, while the primary-side pressure was above that of the secondary side, that is, before [ ]'6# seconds. PRHR heat removal began [ ] seconds into the test. He PRHR was responsible for all the IRWST heat up until ADS 1 activation, after which the PRHR heat transfer reduced significantly. During the active phase, the PRHR transferred heat to the IRWST at an average rate of [ J'6# Btu /sec. Energy Transport via the Break and Automatic Depressurization System He mass flow rate from the primary system via the break is shown in Figures 5.12.2-67 and 5.12.2-68. As shown in these figures, liquid flow was detected by the flow measuring devices for the short-term transient. During the first [ ]'6# seconds following the break, [ ]'6# lbm of water flowed out of the primary system via the break at an average rate of approximately [ ]'6' lbm/sm During this period, the primary system depressurized to around [ J'6# psi (Figure 5.12.21). With the initiation of ADS 1-3, flow through the break stopped and was replaced by steam and liquid flow through the ADS 13 valves. Between [ ]'6# and [ ]'6# seconds, ADS 13 caused the system f to depressurize rapidly and at [ ]'6# seconds, ADS 4 was initiated and the primary system continued to depressurize to the BAMS heater pressure, mangoouwst.non:tb-too395 5.12.2-3 REVISION: 1 _ ._- _ _ _ _ . _ . _ _ _ _ _ _ _ _ ,~ _
l Initiation of the ADS terminated the flow through the break. During ADS 1-3 depressurization, break flow was replaced by steam and liquid flow through the ADS 1-3 valves at a peak rate of over [ ]'" lbm/sec. Flow through the ADS continued at a declining rate until [ J seconds when the flow through the ADS 1-3 terminated and was replaced by flow through the lower resistance ADS-4 paths. l By the end of the short-term transient, water was flowing out of the two ADS-4 valves at approximately [ J'" lbm/sec. (Figure 5.12.2-64). The integrated mass flow from the primary system via the ADS and the break is shown in Figure 5.12.2-62, and the corresponding integrated energy flow is shown in Figure 5.12.2-69. The inventory plot given in Figure 5.12.2 70 indicates that there was little or no steam flow out of the primary system during the short-term transient. Components of the energy balance are shown in Figure 5.12.2-74. l l l l l l l O1l l 9 mup600c3hst.noo:ib-ioo395 5.12.2-4 REVISION: 1
O h TABLE 5.12.21 OSU TEST ANALYSIS S1 ANDARD PLOT PACKAGE FOR SUBSECTION 5.12.2 Plot No. Component Variables Units Description 1 Pressurizer Cirr-604 psia System pressure 2 RPV RPVPWR kW Core power 3 RPV TOIRPV,708RPV, "F Core inlet / outlet temperature, ST08RPV saturation temperature 4 SG Cirr 201, Cfrr 204, psia Pnmary and secondary pressures in SG Cirr-301, CPT-302 5 DVI l WWTDVIL1, Ibm /sec. Individual components and total flow in WWTIRWII, DVI-l WOUTACC1, WWTIRWI3 6 DVI-2 WW'IDVIL2, Ibm /sec. Individual components and total flow in WWTIRWI2, DVI-2 WOUTACC2, WWTIRWI4 7 CMT AMCMTIB, Ibm Fluid mass in CMTs (excludes balance AMCMT2B lines) 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs 9 CMT MIWDVIL1, Ibm Integrated mass out of CMTs MIWDVIL2 ( 10 CMT WWTDVILI, Ibm /sec. Flow out of CMTs WWTDVIL2 11 CMT WOUTCLBI, Ibm /sec. Flow into CMTs WOUTCLB2 12 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines 13 CMT UCMT1, UChf!2 Btu Fluid energy in CMTs 14 IRWST IRWST lbm Mass of fluid in IRWST 15 IRWST CLDP-701 in. Collapsed liquid level in IRWST 16 IRWST W%TIRWII, Ibm /sec. Flow from IRWST to DVI lines WWTIRWI2 17 IRWST IRWSTOR lbm/sec. Overflow from IRWST to sump 18 IRWST ADS 13TMR lbm/sec. Total ADS flow into IRWST 19 IRWST ADS 13TIR, MIIRWII, Ibm Integrated mass out of IRWST MIIRWI2, MIIRWIO 20 IRWST UIRWST Btu Fluid energy in IRWST 21 PRHR CLDP-802 in. Collapsed liquid level in PRHR HX O m:vm2n4w 51.non:isico395 5.12.2-5 REVISION: 1
TABLE 5.12.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECrlON 5.12.2 Plot No. Component Variables Units Description 22 PRifR WWOTPRHR Ibm /sec. Measured outlet flow from PRHR tube 23 Accumulator AMACCI, AMACC2 lbm Mass of Huid in accumulators 24 Accumulator CLDP-401, CLDP-402 in. Collapsed liquid level in accumulators 25 Accumulator WOUTACCl, Ibm /sec. Flow from accumulators WOUTACC2 26 Accumulator MOUTACCl, Ibm Integrated mass out of accumulators MOUTACC2 27 Accumulator UACCl, UACC2 Btu Fluid energy in accumulators 28 Prunary sump AMPSMP lbm Prunary sump fluid mass 29 Primary sump CLDP-901 in. Pnmary sump level 30 Pnmary sump UPSMP Btu Pnmary sump fluid energy 31 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG primary side MSSGOP1, MSSGOP2 inlet / outlet plena 32 SG MSSGHT1, MSSGHT2, Ibm Mass of fluid in SG primary side hot MSSGCT1, MSSGCT2 and cold tubes 33 SG/PRHR Cirr-201, CI'r-301, psia & SGI pressure and PRHR integrated heat QPRHR1 Btu output 34 Pressurizer PZM lbm Fluid mass in pressurizer 35 Pressurizer CLDP-601 in. Collapsed liquid level in pressurizer 36 Pressurizer UPZ Btu Fluid energy in pressurizer 37 Surge line PLM lbm Fluid mass in surge line 38 Surge line CLDP-602 in. Collapsed liquid level in surge line 39 Surge line UPSL Btu Fluid energy in surge line 40 RPV MWRPV lbm Total fluid mass in reactor vessel 41 RPV DCM lbm Fluid mass in downcomer 42 RPV LDP01DC in. Collapsed liquid level in downcomer compared to various reference elevations 43 RPV MW0lRPV lbm Fluid mass in lower plenum 44 RPV MWO3RPV lbm Fluid mass in core region 45 RPV LDP03RPV in. Collapsed liquid level in core 46 RPV RPVAVDF2 Core exit void fraction 47 RPV RPVAQOU2 Core exit quality 48 RPV MWO6RPV lbm Fluid mass in the upper plenum 49 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 50 RPV MWO8RPV lbm Fluid mass in the upper head 51 RPV LDP08RPV in. Collapsed liquid level in the upper head m:up600awsi.non:1b-ioo395 5.12.2-6 REVISION: 1
O TABLE 5.12.21 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.12.2 Plot No. Component Variables Units Description 52 RPV URPV Btu Total fluid energy in reactor vessel 53 RPV RPVXE, RPVASL2 in. Level of Tsat line 54 RPV RPVPab, RPVAPab2, kW Heated rod power above and below RPVPWR Tsat level and total 55 RPV RPVRXV,RPVASOU2 lbm/sec. Core steam generation rate 56 RPV RPVALIN2 lbm/sec. Calculated core flow 57 RPV HTMXRPV, S'!1)8RPV 'F Maximum clad temperature and saturation temperature 58 Hot leg MWHL1, MWHL2 lbm Water mass in bot legs 59 Hot leg MVHL1, MVHL2 lbm Vapor mass in hot legs 60 Cold leg CLlWMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 61 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 62 ADS and break BRKSTIR, ADS 1311R, Ibm Total discharged mass for ADS 1-3. ADS 41TIR, ADS 42TIR ADS-4s, and break 63 ADS and break BRKTIVF, AD13TIVF, Ibm Total integrated vapor flow for ADS AD41TIVF, AD42TIVF and break -O. 64 ADS and break BRKTILF, ADl3TILF, Ibm Totalintegrated liquid flow for ADS AD41TILF, AD42TILF and break 65 ADS and break ADS 13SVR, Ibm /sec. Vapor flow out ADS 1-3 and ADS-4 ADS 41SVR, ADS 42SVR 66 ADS and break ADS 13SLR, Ibm /s-c. Liquid flow out ADS 1-3 and ADS-4 ADS 41SLR,' ADS 42SLR 67 ADS and break BRKSSVR lbm/sec. Vapor flow out of break 68 ADS and break BRKSSLR lbm/sec. Liquid flow out of break 69 ADS and break BRKSPEl, ADS 13El, Blu Integrated fluid energy for ADS 1-3, ADS 41El, ADS 42EI ADS-4, and break 70 Mass balance TOTMASS lbm Total system mass inventory 71 Mass balance PRIMMASS, Ibm Measured pnmary system inventory and PRIMASS2 value from mass balance 72 Mass balance MERROR Ibm Mass balance error 73 Mass balance MIN, MOUT lbm Integrated mass flow in and out of SRCMASS pnmary system and source mass 74 Energy balance Various Btu Components of energy balance O mAap6002344w 51.non:lb.100395 5.12.2-7 REVISION: 1
O TIIE FIGURES LISTED IN TABLE 5.12.21 ARE NOT INCLUDED IN TIIIS NONPROPRIETARY DOCUMENT O O m:ww.51..oo:id. ion 995 5.12.2-8 REVISION: 1
5.12.3 Long-Term Transient The long-term transient started with initiation of IRWST injection, covered the transition from IRWST to sump injection, and provided information on the LTC response of the AP600 plant. For the 0.5-in. cold-leg break, Matrix Test SB23, the long-term transient analyzed ran from [ ]'6' seconds to the end of the test around [ ]'6# seconds. De behavior of the test facility during this period of the transient is discussed in this subsection using the plot package detailed in Table 5.12.3-1. His analysis concentrates on the components of the primary system that remained active during the LTC phase, that is, the RPV, the hot legs, ADS-4, the sumps, and the IRWST. For the long-term transient, thermal-hydraulic phenomena of interest were:
- Maintenance of core cooling and removal of energy from the primary system.
=
Level oscillations (from [ ]'6# seconds; there were system wide level and pressure oscillations, which are discussed further in Subsection 6.1.3). 5.12.3.1 Maintenance of Core Cooling Mass Injected into Primary System Total DVI line flow, CMT flow, and IRWST flows are shown in Figures 5.12.3-6 and 5.12.3-7, and the flow from the primary sump is shown in Figure 5.12.3-19. From around [ ]'6* seconds, there was a contribution to the DVI flow from the CMTs as the CMTs reached post-refill draindown. During the pre-sump injection phase of the transient, IRWST flow proceeded at a gradually declining rate with the effect of the primary system oscillations superimposed. At [ ]' 6d seconds, flow from the primary sump began through the check valves around the main injection valves, further reducing IRWST flow. From [ ]'6' seconds to the end of the transient, a nearly steady flow rate of[ ]'6# lbm/sec. was maintained through each DVI line. At [ ]'** seconds, the main sump ) injection valves opened, reversing flow through IRWST injection line-1 and increasing IRWST flow in l line-2. He net result was that an injection flow rate of [ ]*6* lbm/sec. was maintained through both DVI lines. l Reactor Pressure Vessel and Downcomer Response l The effect of the water inflow on the average measured downcomer fluid temperatures, core inlet and I core outlet temperatures, and heater rod temperatures during the long-term phase of the transient is shown in Figures 5.12.3-4,5.12.3-5, and 5.12.3 38. Figure 5.12.3-4 shows that there is a general O increase in average downcomer fluid temperatures during the long-term transient. By the end of the U test, this average temperature reached an equilibrium of [ ]'6' 'F below saturation. Figure 5.12.3-5 maar60m2m-si.no :ib-too395 5.12.3-1 REVISION: 1
implies that the core remained at or near saturation for all of the long-term transient after [ ]' 6' seconds. However, Figures 5.12.3-34 to 5.12.3-36 show that the DVI line flow method described in Section 4.11 indicates that a small level of boiling was maintained after [ ]'6* seconds into the transient. Nevertheless, the level of boiling was small and showed that the inflow from the IRWST and sumps was sufficient to maintain cooling. Figure 5.12.3-38 shows that there were no significant excursions in heater rod temperatures throughout the long-term transient; therefore, sufficient core inventory and flow was maintained through this phase of the transient to remove the decay heat generated. For significant portions of the transient, a two-phase mixture was present in the core and upper plenum regions. De following discussion tracks the variation in water level and mass throughout the reactor vessel and downcomer. De mass and level for the core region are shown in Figures 5.12.3-28 and 5.12.3-29. The collapsed liquid level in the core indicated that the heater rods were always covered with a single-or two-phase mixture. During the later stages of the transient, the collapsed liquid level remained just below the top of the heater rods, and the core void fraction was [ ].**# The fall in core inventory was a result of the influx of hot water from the primary sump as it flowed through the check valves. The impact of this hot water on the system temperatures is shown in Figures 5.12.3-4 and 5.12.3-5 as a sudden increase in fluid temperature in the downcomer and at the core inlet. The hot water also led to an increase in the calculated steam generation rate, as shown in Figure 5.12.3-36, and a corresponding fall in the level at which the core reached saturation temperature (Figure 5.12.3 34). The collapsed liquid level in the upper plenum region covered by LDP-113 is shown in Figure 5.12.3-32. The figures indicate that during the period before sump injection began, the collapsed liquid level initially fell and then remained at the top of the hot legs. Following the influx of hot water from the sumps, the level dropped first to the hot-leg mid-elevation and then to the top of the DVI injection lines, where it remained for the remainder of the transient. This level corresponded to a void fraction of ( ].'6" The mass of water in the reactor pressure vessel is shown in Figure 5.12.3-25. After an initial decline, the reactor vessel water mass settled at an average value of [ ]'6# lbm until the time sump injection started when it gradually fell to [ ]*** lbm, which is [ ]'6' percent of the initial vessel water inventory. From [ ]'6' seconds, oscillations in vessel inventory were observed although, these oscillations are not as marked for SB23 as those observed in SB01. Figures 5.12.3-51 to 5.12.3-56 illustrate these oscillations using plots on a restricted time frame from [ ] seconds. These oscillations are observed in primary system measurements from the upper plenum to the ADS-4 flows. The oscillations occurred for [ ]'6' seconds. The oscillations in the ADS flow lagged behind those in the upper head pressure by around [ J'6' seconds. These oscillations and ! possible mechanisms for their production are discussed further in Subsection 6.1.3. The mass of fluid and collapsed liquid level in the RPV downcomer are shown in Figures 5.12.3-26 and 5.12.3-27. The collapsed liquid level remained above the cold legs for the entire long-term m%p60m2m5i.oco:ib-ioo395 5.12.3-2 REVISION: 1
transient. 'Ihe start of sump injection reduced the level, but this was not sufficient to uncover the cold legs. 5.12.3.2 Energy Transport from the Primary System During the long-term transient, energy continued to be deposited in the primary system from the heated rods, metal, and fluid flowing from the primary sump. 'Ihe SGs and PRHR remained inactive .
- throughout this phase of the transient and the primary path for energy out of the primary system was via the ADS 4 valves.
Integrated mass flow from the primary system via the ADS and the break is shown in Figure 5.12.343. During the L'IC phase of the transient, the only significant outflow is through the l ADS 4 valves. 'Ihis is confirmed by Figures 5.12.3-44 to 5.12.345, which show flow through the ADS and break. During the sump injection phase of the transient, measured outflow is in the form of liquid out of the ADS 4 valves. Water flowed through each of the valves at an average rate of ! [ ]'6# lbm/sec. l Figure 5.12.3-36 shows the calculated steam generation rate, as determined by the DVI line flow method. During the sump injection phase of the transient, steam was generated at over [ ]' *# lbm/sec., although the steam vortex meters indicate little or no flow out of the ADS 4 valves. The following two indications show that steam is leaving the primary system by this route: Figure 5.12.346 shows total measured system fluid inventory. During this phase of the transient after the start of primary sump injection, that is, when core steam generation was most significant, the total system inventory fell by about [ ]'6# lbm. This amount corresponds to a steam flow rate of [ ]'6# lbm/sec., which would not have been detected by the vortex meters. I
. Examination of the fluid thermocouples on the outlet of the ADS 4 valves indicates that j temperatures remained at or above saturation temperature following the start of sump injection. I It was not possible for all the steam generated in the core to flow from the upper head to tim downcomer via the bypass holes (Subsection 6.1.3). Therefore, steam was leaving the primary system via ADS 4. Figure 5.12.3-50 shows all the components to the system energy balance. Further discussion of steam loss from the primary system is provided in the mass and energy balance discussions of Section 6.2.
O G m:\ p6 col 2344w.si.=:ib.too995 5.12.3-3 REVISION: 1
1 i TABLE 5.12.31 OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.12 3 . LONG-TERM TRANSIENT I l Plot No. Component Variables Units Description j 1 RPV RPVPWR kW Core power 2 Pnmary sump TSMPII, TSMPI2 *F Sump injection line temperatures i 3 DVI TDVIL1, TDVIL2 'F DVI line temperatures 4 RPV TOIDC, TD2DC,1D3DC, *F Water and saturation temperatures in ST01DC downcomer 5 RPV TOIRPV, TD8RPV, *F Core inlet / outlet temperature, ST08RPV saturation temperature l 6 D VI-l WWTDVIL1, Ibm /sec. Individual components and total flow i WWTIRWII, in DVI-l I WWTIRWI3 7 DVI-2 WWTDVIL2, Ibm /sec. Individual components and total flow WWTIRWI2, in DVI-2 , WWTIRWI4 l 8 CMT CLDP-502, CLDP-507 in. Collapsed liquid level in CMTs l 9 CMT CLDP-509, CLDP510 in. Level CL-CMT balance lines ) 10 IRWST IRWST Ibm Mass of fluid in IRWST I1 IRWST CLDP-701 in. Collapsed liquid level in IRWST ) 12 IRWST UIRWST Btu Fluid energy in IRWST 13 Pnmary sump AMPSMP lbm Pnmary sump fluid mass 14 Primary sump CLDP-901 in. Pnmary sump level 15 Primary sump UPSMP Btu Pnmary sump fluid energy 16 Secondary sump AMSSMP lbm Secondary sump fluid mass 17 Secondary sump CLDP-902 in. Secondary sump level 18 Secondary sump USSMP Btu Secondary sump fluid energy 19 Primary sump WSTSMPET, WWTSMPIT lbm/sec. Pnmary sump steam and liquid injection rate 20 Pnmary sump MISMPII, MISMPI2, Ibm Integrated primary sump and IR'WST MISMPfr. MIIRWT flows 21 SG MSSGIP1, MSSGIP2, Ibm Mass of fluid in SG side inlet /wtjet MSSGOP1, MSSGOP2 plena 22 Surge line PLM lbm Fluid mass in surge line 23 Surge line CLDP-602 in. Collapsed liquid level la :. urge line 24 Surge line UPSL Btu Fluid energy in surge line 25 RPV MWRPV lbm Total fluid mass in reactor vessel e m:\apsom2m.5t.non:ib-too395 5.12.3-4 REVISION: 1
. - . - - . . . - - - - - - - . . - - - - - . . . - . . _ - - - - - . - - . _ . - - . _ _ = .
I O TABLE 5.12.31 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUllSECTION 5.12.3 LONG TERM TRANSIENT Plot No. Component Variables Units Description 26 RPV DCM lbm Fluid mass in downcomer 27 RPV LDP01DC in. Collapsed liquid levelin downcomer compared to various reference elevations 28 RPV MWO3RPV lbm Fluid mass in core region 29 RPV LDP03RPV in. Collapsed liquid level in core 30 RPV RPVAVDF2 Core exit void fraction j 31 RPV RPVAQOU2 Core exit quality 32 RPV LDP06RPV in. Collapsed liquid level in the upper plenum 33 RPV MWO8RPV lbm Fluid mass in the upper head 34 RPV RPVASL2 ft. Level of Tsat line 35 RPV RPVAPab2, RPVPWR kW Heated rod power above and below - Tsat level and total 36 RPV RPVASOU2 lbm/sec. Core steam generation rate 37 RPV RPVALIN2 lbm/sec. Calculated core flow 38 RPV HTMXRPV, 'F Maximum clad temperature, saturation O 39 Hot leg ST08RPV MWHL1, MWHL2 lbm temperature and delta Water mass in hot legs 40 Hot leg MVHL1, MVHL2 lbm Vapor mass in hot legs i 41 Cold leg CL1WMS, CL2WMS, Ibm Water mass in cold legs CL3WMS, CL4WMS 42 Cold leg CLIVMS, CL2VMS, Ibm Vapor mass in cold legs CL3VMS, CL4VMS 43 ADS and break BRKSTIR, ADS 13TIR, Ibm Total discharged mass for ADS 1-3, ADS 41TIR, ADS 42TIR ADS-4, and break 44 ADS and break ADS 13TLR, ADS 41TLR, Ibm /sec. Liquid flow out ADS 1-3 and ADS-4 ADS 42TLR 45 ADS and break BRKSTLR lbm/sec. Liquid flow and total flow out of break 46 Mass balance TOTMASS lbm Total system mass inventory 47 Mass balance PRIMMASS, PRIMASS2 lbm Measured prtmary system inventory and valve from mass balances 48 Mass balance MERROR lbm Mass balance error 49 Mass balance MIN, MOUT SRCMASS lbm Integrated mass flow in and out of primary system and source mass 50 Energy balance Various Btu Component of energy balance 51 ADS-4 ADS 41TLR, ADS 42TLR lbm/sec. Oscillations in ADS-4 liquid flow 52 Surge line CLDP-602 in. Oscillations in surgeline level (. 53 RPV CPT 107 psia Oscillations in upper head pressure maap600G344w.51.non:lb 100395 5.12.3-5 REVISION: 1
l TABLE 5.12.3-1 (Continued) OSU TEST ANALYSIS STANDARD PLOT PACKAGE FOR SUBSECTION 5.12.3 LONG. TERM TRANSIENT Plot No. Component Variables Units Description 54 RPV CLDP-113 in. Oscillations in upper plenum level 55 RPV LDP03RPV in. Oscillations in core level 56 RPV LDP01DC in. Oscillations in downcomer level i i 1 1 O l l l l l l l O m:W344w.51.non:1b too395 5.12.3-6 REVISION: 1
g.._._.______.____ _ . _ _ _ _ _ - . _ _ _ _ . _ _ _ . . _ _ _ . _ _ _ . _ _ _ _ - - - o 1
- I 4
i l i ! 1 4 l 4 4 l i , 4 i
- i 5
3 e a 2 THE FIGURES LISTED IN TABLE 5.123-1 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT j 1 i 1 4 5 1 i i 4 e : 4-I 1 1 1 , i 4 t l I a i ? ? 4 ) I 4 : I l e t, t i l , 1 1 4 1-1 i i 2 m:ww51.non:it loo 995 5.12.3-7 REVISION: 1 j l
4 6.0 TEST FACILITY PERFORMANCE - This section discusses the following test anelysis issues common to the matrix testing performed at Oregon State University (OSU): Observed thermal-hydraulic behavior, including the phenomena associated with core makeup tank (CMT) refill and oscillation in the reactor
- Passive residual heat removal (PRHR) performance
- Results of the mass and energy balance
- Effects of nitrogen dispersal from accumulator injection S
1 CD m:ur60m2N60.non:tt>.too395 6-1 REVISION: 1
t 6.1 Observed Thermal Hydraulic Phenomena The following thermal-hydraulic phenomena, observed during the matrix test program and identitled in the OSU Final Data Report (" are evaluated: CMT refill, which appeared to be the result of condensation in the CMT after initial injection and draindown
. PRHR performance Causes of pressure and liquid level oscillations in the reactor . Effects of accumulator nitrogen I
O O m:W344w-mon:ib too395 6.1-1 REVISION: 1
l l I 6.1.1 Core Makeup Tank Refill Response - l Upon initiation of CMT draindown, incoming steam from the reactor coolant system (RCS) via the , cold leg balance line heated the CMT inside wall surface. Figure 6.1.1 1 represents the average fluid ' and through-thickness metal temperatures and inside surface heat flux for one metal segment in the upper head of CMT-1 based on Matrix Test SB18; Figure 6.1.1-2 represents the same parameters for CMT-2. Given the sign convention used in the OSU data analysis code, the negative heat flux I represents heat flow into the me,tal CMT walls from fluid in the enclosed volume. ) Following initial draindown, the CMTs and respective cold-leg balance lines contained steam and, possibly, some quantity of noncondensable gas, ne back pressure from the IRWST flow caused the CMT check valve to close and remain closed such that the CMT tank was isolated from the DVI line. As the uninsulated CMTs continued to transfer heat to the environment, steam condensed in the CMTs, resulting in a decrease in CMT pressme relative to primary system pressure. At this time, both the cold leg and the cold-leg balance line connection at the top of the CMT were liquid solid. As the condensation continued, the decrease in CMT pressure resulted in an increase in the cold-leg balance line level (Figure 6.1.1-3). When the cold-leg balance line completely refilled, water entered ; the CMT through the inlet diffuser, resulting in rapid condensation of steam in the CMT and a corresponding rapid additional decrease in CMT pressure (Figure 6.1.1-4). He sudden decrease in pressure resulted in a quick refill of the CMT through the balance line (Figure 6.1.1-5). Retill of subcooled fluid resulting in a CMT condensation /depressurization event is reflected in the ' ~ reactor vessel mass (Figure 6.1.1-6). De event is also reflected in the cold leg water levels (Figure 6.1.1-7). l Re CMT refill event continued until the cold legs uncovered in the reactor vessel downcomer. Figure 6.1.18 shows that the end of the re1111 event correlates to the downcomer level and, additionally, that this level corresponds to the top of the cold-leg nozzles located at about the 71 in. elevation. De system water inventory, available prior to uncovering the cold legs, was l sufficient to refill the CMT to about two-thirds full: therefore, the CMT refill event was terminated prior to the entire CMT refilling. CMT refill, drawing fluid from the cold leg, also caused a rapid decrease in the measured CMT l internal fluid temperatures which caused local fluid temperatures to fall below local CMT metal temperatures. His resulted in the heat transfer process reversing itself as seen by the sign change of l the heat flux from negative to positive; heat flowed from the now heated CMT walls into the adjacent subcooled liquid, he CMT metal temperatures are seen to rapidly decrease as energy stored in the metal previously heated by steam was now transferred into the adjacent liquid in the CM l', The phenomena that drive the CMT refill response are identical for both CMTs and whichever CMT refills first appears to be random. However, once the refill event was initiated in a particular CMT, the accompanying decrease in system fluid levels precluded the event from taking place in the may6cocm*60. on:lt>.too395 6.1.1 1 REVislON: 1
remaining CMT until the first refill event terminated and the system fluid levels were restored by IRWST injection. This characteristic is illustrated in Figure 6.1.1-9, which shows that the CMT-1 balance line level was depressed as a result of CMT-2 refill, immediately, following recovery of the downcomer level, the CL-3/CMT-1 balance line began to refill; as soon as the balance line level recovered to completely full, the CMT-1 refill event was initiated.. l 1 This refill phenomenon generally occurred in all tests where ** t cold-leg balance line was intact. Additional discussion and comparison of the CMT refill phenomenon is given in Subsection 7.2.2. The 1/4 scale of the OSU test facility contributes to the facility's ability to refill. The decrease in ) CMT pressure, which lifted the liquid up and filled the cold-leg balance line, was only about ) [ ]'6' psi (Figure 6.1.1-3). This corresponds to an absolute CMT pressure of about [ ]' 6d psia in the OSU test. In comparison, the pressure decrease in the AP600 would have to be about ; [ ]'6" psi. Since the reactor system is nearly at containment pressure, about [ ]*6# psia for a i SBLOCA during this period, the CMTs would have to be at an absolute pressure of about [ ]*** psia. Other factors that could influence the potential for, as well as the timing of, a CMT refill include: energy stored in the CMT metal, heat loss to the environment, and presence of noncondensables in the CMT. While it is possible that CMT refill could occur in the AP600, and evidence of the CMT balance line l beginning to refill was observed in SPES-2,* the case with which the CMTs refilled in the OSU test i is believed to be a result of a scaling distortion. Therefore, refill of the CMTs would be less likely to l occur in the AP600 than in the OSU test. l I l 1 O mAap60m2344w-60.non:1b-100395 6,1,1 2 REVISION: 1
__.. ..- . ~ - - . - - - --- O TABLE 6.1.11' l OSU TEST ANALYSIS PLOT PACKAGE FOR SUBSECTION 6.1.1 - 1%t No. Component Variables Units Description 1 CMT - T Btu /sec. CMT-1 Average fluid and through-thickness metal temperatures and upper head beat flux 2 CMT - T Btu /sec. CMT-2 Average fluid and through-
]
thickness metal temperatures and upper head beat flux , 1 3 CMT - ec passe on hw psia in. !ine level l 4 , Effect of flow from CMT balance line Ibm /sec. psia on CMT pressure e - msponse to NT 5 CMT - psia in. pressure decrease 1 6 CMT - in. Ibm Effect of CMT refill on RPV mass 7 CMT - in. Effect of CMT refill on cold-leg level g _ Effect of CMT refill on downcomer in. level 9 CMT - in. Effect of CMT refill on balance line and downcomer levels ; I l l l I O l mAsp(m2344w-60.noo:1b-loo 395 6.1.1-3 REVISION: 1
O THE FIGURES LISTED IN TABLE 6.1.1-1 ARE NOT INCLUDED IN THIS NONPROPRIETARY DOCUMENT l l l i i l 9 m:\ap60(A2344w40.non:Ib-100905 6.l.1-4 REVISION: 1
6.1.2 Passive Residual Heat Removal System Performance - he passive residual heat removal (PRHR) system is designed to remove core decay heat in the event that the active safety-related systems are not available. He PRHR heat exchanger (HX) is located inside the in-containment refueling water storage tank (IRWST). It consists of a series of C-tubes, each end connected to an inlet and outlet header, which are connected to the hot leg (inlet) and pump suction (outlet), respectively. The inlet header enters the IRWST slightly below the water surface and the outlet header exits the tank near the bottom. At normal operating conditions, the PRHR is isolated from the primary system. In the event of an accident, the isolation valves are opened and natural circulation ilow is established, driven by the density difference between the hot fluid entering the PRHR and the cold fluid exiting the PRHR. Heat is transferred from the tubes to the IRWST water by either subcooled boiling or free convection. His section describes the heat transfer characteristics of the OSU representation oithe PRHR/IRWST. Also included is a discussion of the PRHR heat removal from the primary system and the interaction between the PRHR and other heat removal mechanisms, including the energy out of the break, the heat transfer through the steam generators, and the automatic depressurization system (ADS). The calculations presented in this section were performed for Test SB18. However, the conclusions are valid and describe the PRHR performance in all of the LOCA tests based on calculations for these other matrix tests. He heat transfer from the PRHR tubes to the IRWST can be calculated by either of the following two methods: Inferred from the increase in the internal energy of the IRWST water a Heat transfer calculations from the external tube walls to the IRWST water Passive Residual Heat Removal System Performance Inferred from the In-Containment Refueling Water Storage Tank Energy Increase The most accurate method for determining the heat transfer from the PRHR tubes to the IRWST is to infer the heat transfer by observing the increase in the internal energy of the IRWST water. He PRHR is the sole source of energy input into the IRWST until actuation of the ADS 1-3. After this time, the PRHR heat transfer is determined by: Qpaga = (dU/dt)mwsr - QADSn 6.1.2-1 where: Qpana = Heat transferred from the PRHR, Btu /sec. (dU/dt)mwrr = Time rate of change of the IRWST water energy, Bru/sec. QAoso = Heat transferred from the ADS stage 1-3 flow, Btu /sec. mAap600\2344-612.non:1b-100395 6.1.2-1 REVISION: 1
Equation 6.1.2-1 can be integrated:
!Qrana d t = ( U - Uo )mwsr l - QAosti t 6.1.2-2 The water temperature in the IRWST was measured at axial locations. By associating a volume with each of these measurements, a total fluid internal energy was calculated for each time step. The difference between this value and the internal energy of the tank at the start of the test is used in Equation 6.1.2-2 to determine the total PRHR energv transferred until actuation of ADS 1-3. No evaporative heat loss from the IRWST was considered in this calculation, as the IRWST is highly subcooled during the period ofinterest (<1000 sec). The results of these calculations for Matrix Test SB18 are shown in Figures 6.1.2-1 to 6.1.2-3.
Passive Residual Heat Removal System Performance Determined from the External Tube Heat Transfer The PRHR heat exchanger consists of 88 tubes in a C-tube arrangement. Two of the tubes, the longest (the top tube in the inlet tube bundle), and the shortest (the bottom tube in the inlet tube bundle) have been instrumented. The tubes are fed from a common header; flow in the long and short tubes may be different depending on conditions in the header. Four thermocouples are located on each ~ instrumented tube to determine the tube external wall temperature at the center of the horizontal inlet tube, one-fourth of the way down the vertical tube, three-fourths of the way down the vertical tube, and at the center of the horizontal outlet tube. The temperatures in the IRWST are determined at twelve axial locations from the bottom to the top of the tank. These instmments are summarized in Table 6.1.2-1. Heat transfer calculations were performed for both tubes at the four thermocouple locations. One-half of the tubes (44) were assumed to behave as short tubes, while the remainder were assumed to behave as long tubes. Previous studies of the PRHR performance
- indicate that the condensation of a high temperature two-phase mixture inside the inlet portion of the tubes results in high heat fluxes, which promote subcooled boiling on the external tube surface. However, the tube wall temperatures at the PRHR inlet for OdUm indicate that the boiling did not occur for any of the tests. This is due to the lower primary system pressure in OSU, which reduces the energy of the flow into the PRHR relative to the SPES tests which were performed at full pressure. Consequently, the external heat transfer in the PRHR at OSU is characterized solely by free convection for all test conditions.
Horizontal Tube Heat Transfer The heat transfer coefficient is calculated using the free convection correlation for horizontal tubes from Holman:"U h = k/d 0.53 ( GrPr )'" 6.1.2-3 mhp600c3*612.nondb. loo 395 6.1.2-2 REVISION: 1
O () where: k = Liquid thermal conductivity in the tank, Btu /sec.-ft. 'F d = Tube outer diameter, ft. Pr = Liquid Prandtl number Gr = Liquid Grashof number given by Gr = g$ (T, - T% ,) d'/ v2 6.1.2-4 where: g = Gravitational constant, ft/sec.2 ! v = Liquid kinematic viscosity, ft.2/sec. T, = Wall temperature of the tube section, 'F T, %
= IRWST water temperature in the vicinity of the tube, 'F l
= Volumetric expansion coefficient given by
@ = ( p%a - p, ) / [ p,( T - T% , )] 6.1.2-5 where:
p%, = Liquid density evaluated at the bulk tank temperature, Ibm /ft.' I p, = Liquid density evaluated at the wall temperature, Ibm /ft.' Finally, the heat transfer from the tube section is given by: Q, = hL ,xd ( T - T% , ) N,w 6.1.2-6 where: L, = Length of the tube segment, ft. N,w = Number of tubes represented in the calculation i Vertical Tube Heat Transfer The heat transfer rate in the vertical segment was evaluated in the same way except that the free convection correlation for vertical cylinders was used. The free convection heat transfer coefficient is given by Holman"0 as: I h = k 0.13 [ g ( T - Tw ) Pr / v2 jus 6.1.2-7 m:\ap600'.2344-612.nos.;1b-100395 6.1.2 3 REVISION: 1
The heat transfer from the vertical tube sections is given by equation 6.1.2-6. External Tube Heat Transfer Results De results of this study are presented in plots that are shown in Table 6.1.2-3. The heat transfer rate as calculated for the long tubes is shown in Figure 6.1.2-1. His figure shows that during the period of steady operation, nearly [ ]* percent of the total heat transfer occurred in the inlet horizontal section, while [ ]* percent occurred in the vertical section, and [ ]* percent occurred in the outlet ' horizontal section. Rese results are due to the high heat flux associated with the two-phase mixture condensing in the inlet tube, and are consistent with observation from the SPES tests.* Re heat transfer rate remained essentially constant at [ ]* Btu /sec once natural circulation flow had been established until the initiation of ADS 1-3 at approximately ( ]* seconds. The heat transfer for the short tubes is shown in Figure 6.1.2-2. The heat transfer rate was much lower for the short tubes than for the long tubes during the most significant period of PRHR operation (up until ADS actuation). This is due to density stratification in the inlet header, which preferentially feeds two-phase mixture to the top (long) tubes, and single-phase liquid to the bottom (short) tubes. His is verified by the inlet plenum differential pressure cell, which indicates that the plenum drains at ADS actuation. During the pre-ADS period, the heat transfer was equally split between the inlet, vertical, and outlet tube segments. Figure 6.1.2-2 also shows that the short tube heat transfer increases after ADS actuation. His is due O to the draining of the PRHR header to the point where only the tubes at the bottom of the bundle continue to receive flow. A Figure 6.1.2-3 shows a heat balance on the IRWST and includes the integrated heat transfer from all the PRHR tubes, the integrated heat flow into the IRWST from the ADS 1-3, and the change in the IRWST internal energy relative to the beginning of the test. Also shown is the energy removed from the primary system by flow out the 2-in. break. His figure shows good correlation between external tube heat transfer calculation and the IRWST internal energy calculation for times up to the initiation of ADS 1-3 at [ ]* sec. After this time, the PRHR continues to transfer heat at a reduced rate until initiation of ADS-4 at [ ]* seconds, when the PRHR heat transfer ceases. During the early stages of a SBLOCA, the primary system sensible heat and the reactor decay heat are removed by the flow of steam and waMr out of the break, heat transfer in the SGs from the primary to the secondary side, and heat transfer in the PRHR. As was observed in both the OSU and SPES tests, the SG heat transfer is significantly reduced after the secondary-side is isolated and the primary side flow transitions to natural circulation and ends when the primary side pressure reaches the secondary side pressure. De two sides are then in thermal equilibrium, and the steam generator tubes are drained. At this time, the primary heat removal paths become the break and the PRHR. In the case of the 2-in. break, flow out of the break accounts for six ti,mes as much energy removal as the PRHR. J Rese results indicate that the PRHR is far more effective at removing energy from the primary system m%p60tn234+612.non:ib-too995 6.1.2-4 REVISION: 1
A in non-LOCA events such as the steam generator tube rupture (SGTR), where'the primary system pressure and inventory remain high. 4 The results observed for SB-18 were typical of all the LOCA tests. 1 J. o 1 i s i l l 4 4 4 4 i a lI ' d j 4 m4*n23n612.noo:ib.too395 6.1.2-5 REVISION: 1
'-"-
- v--. ..w. - - ,-w.. _,_-. , , , , _ _ ., , _ . _ , , _ . _ _ , _ , , , , , , , , _ _, .
TABLE 6.1.2-1 O i INSTRUMENTATION FOR CALCULATING THE PRHR/IRWST IIEAT BALANCE Description Instrumentation Tag PRHR Shon Tube Wall - Inlet Temperature TW-807 PRHR Shon Tube Wall - Venical Temperature TW-805, 'IW-8N PRHR Shon Tube Wall - Outlet Temperature TW-802 . l PRHR Long Tube Wall - Inlet Temperature TW-808 l PRHR Long Tube Wall - Vertical Temperature TW-806, TW-803 l 1 PRHR Long Tube Wall - Outlet Temperature TW-801 l l IRWST Temperatures TF-701, TF-702, TF-703, TF-7N, TF-705, TF-706, l TF-707, TF-708, TF-709, TF-710, TF-711, TF-712 IRWST Pressure IT-701 IRWST Water Level LDP-701 l TABLE 6.1.2 2 9 KEY PARAMETERS FOR CALCULATING THE PRHR/IRWST HEAT BALANCE Description Value PRHR Short Tube - Inlet / Outlet Length 18.26 in. PRHR Sbon Tube - Venical Length 43.93 in. PRHR Long Tube - Inlet / Outlet Length 25.01 in. PRHR Long Tube - Venical Length 57.43 in. Tube Outer Diameter 0.375 in. Number of Long Tubes 44 Number of Shon Tubes 44 O m:\ap60m2344-612.non:lt>100395 6,1,26 REVISION: 1
(o3 TABLE 6.1.2 3 OSU TEST ANALYSIS PLOT PACKAGE FOR SECTION 6.1.2 Plot No. Description 1 Breakdown of Heat Transfer for Long Tubes, SB18 2 Breakdown of Heat Transfer for Short Tubes, SB18 3 Integrated Energy from PRHR and IRWST Heat-up for SB18 O , O m:'ap600N234612.non:Ib-100395 6.1.2-7 REVISION: 1
l O THE FIGURES LISTED IN TABLE 6.1.2-3 ARE NOT INCLUDED IN TIIIS NONPROPRIETARY DOCUMENT O O m: sap 600c344412.non:iti-tow >5 6.1.2-8 REVISION: 1
i 1 6.l.3 Flow Oscillations During Long Term Cooling Cyclic flow, pressure, level, and temperature oscillations were observed in about 65 percent of the OSU tests. These oscillations occurred during the latter stage of the IRWST injection phase of l SBLOCA simulations. Several possible causes of these oscillations, evaluation of test data, and the extrapolated effect in the AP600 plant are identifled in this section. He most likely cause of the oscillations is also presented in detail. 6.1.3.1 Introduction De flow paths in the reactor vessel during IRWST injection are schematically illustrated in Figure 6.1.3-1. Liquid from the IRWST entered the downcomer through the two DVI lines, which are located 180 degrees apart, flowed down through the downcomer and up through the core, where heat addition from the electrical heater rods generated steam. A two-phase water / steam mixture exited the core. Liquid (and possibly some steam) was released from the reactor vessel through the hot legs and ADS-4 lines, which are open during this portion of the test. Steam was separated from the liquid in the upper head and flowed into the downcomer through a series of ten small holes in the downcomer top plate. His steam condensed on the surface of the cooler water in the downcomer. The key features of the oscillations are summarized in Table 6.1.3-1. De oscillations exhibit several common characteristics:
- Re oscillations begin after net steam was generated in the core during IRWST injection.
. He oscillations were regular with a period between 110 and 135 seconds.
. The oscillations started gradually and end gradually (Figure 6.1.3-2).
For all the testr, die measured upper plenum collapsed liquid level oscillated around the top ( f the hot l leg nozzles, and in all tests the measured downcomer collapsed liquid level covered the cold-leg l nozzles. He difference in the oscillation start and stop time and the oscillation period is related to the break size and location. These two parameters affect the amount of the steam generated in the core and eventually, the amount of steam condensed in the downcomer. Five mechanisms have been postulated to explain these oscillations. Each hypothesis is described in detail in Subsection 6.1.3.2. A detailed analysis of the oscillations for SB01, the reference transient, is presented in Subsection 6.1.3.3. He oscillations observed in the other tests are also briefly discussed j in Subsection 6.l.3.3 in comparison with the reference transient SB01. ! O ! maap6002w64.noa:ib. loo 395 6.1.3-1 REVISION: } }
?
6.1.3.2 Proposed liypotheses Flow oscillations were observed to occur during FLECHT-SET gravity-feed reflood tests"U and have been analyzed. The oscillations were found to be natural oscillations, which are U-tube (manometer) oscillations due to the gravity force alone, with an oscillation period of three seconds. The vessel in the OSU test is shorter than that in the FLECHT-SET test. Since the period of natural oscillation is proportional to the height of the vessel, the period of natural oscillations in the OSU test is shorter than the three second period observed in the FLECHT-SET test. However, the observed oscillation period in the OSU test was [ ]'6' seconds. Therefore, the oscillations in the OSU test are not natural oscillations, but are forced oscillations due to force induced by the steam generation and l condensation. In the OSU test, the driving force of the oscillations was determined to be the steam generated in the vessel, which caused the pressure in the vessel to increase. To have oscillations with repetitive cycles, this pressure build-up in the vessel was relieved by venting the steam. The possible vent paths are: (a) through the hot legs and the ADS-4 lines, (b) through the upper head bypass holes and the cold legs, and (c) through the upper head bypass holes with condensation at the top of the downcomer, with the injected DVI line flow as shown in Figure 6.1.3-1. Five candidate hypotheses for the oscillations observed during LTC were investigated and are i described and evaluated in this section. The candidate hypotheses are: 1
- 1. Level fluctuations in the upper plenum opened and closed the steam vent path at the hot-leg i nozzle. Level fluctuations were driven by pressure changes resulting from alternately covering and uncovering the hot-leg nozzle. No condensation occurs in the downcomer. .
l
- 2. Slug flow in ADS-4 lines caused pressure surges when the steam slugs discharged into the separator by changing the two-phase flow regime and pressure drop in these lines.
- 3. The observed pressure fluctuations were driven by changes in condensation rates for steam flowing from the upper head and condensing in the downcomer.
- 4. The observed pressure and level fluctuations were due to alternately covering and uncovering of the hot leg nozzle and steam condensation in the downcomer.
- 5. The observed pressure and level fluctuations were due to alternately covering and uncovering of the upper head bypass holes.
The following sections describe assessment of the data and will show only Hypothesis 4 can be the probable cause of the oscillations. O mw60ccmaoon:tb.too395 6.1.3-2 REVISION: 1
O V 6.1.3.2.1 Hypothesis 1: _ Fluctuating Level About the Hot-Leg Nozzle I i This hypothesis holds that steam pressure in the reactor head increased because all the steam generated . in the core was not released when the two-phase flow vented through the ADS-4 lines. As pressure in the reactor vessel head increased, the level in the reactor vessel upper plenum decreased until the hot-leg nozzle became uncovered by the two-phase mixture. When the nozzle became uncovered, steam was vented from the upper plenum and reactor head through the hot leg to ADS-4. As steam was j vented, pressure in the reactor head decreased and the mixture level in the upper plenum rose until the l
- hot-leg nozzle was covered. Without considering condensation in the downcomer, when the hot-leg
] nozzle became covered again, pressure again began to rise. This cycle continued to repeat until the fluid level in the upper plenum decreased below the hot leg so that steam was continuously vented i through the ADS-4 lines. When steam could flow through the hot leg and ADS-4 lines without restriction, steam was not available to accumulate in the reactor head, and the pressure increase / level decrease fluctuations halted. I Figures 6.1.3 2 shows the collapsed levels in the upper plenum for Matrix Tests SB01 and SB18. For this particular level instmment, the relative position of the hot leg nozzle is: Top of hot leg ( ]# in. Centerline of hot leg [ ]^^* in. Bottom of hot leg [ ]'A' in. The average collapsed level did not decrease until about [ ]'A' seconds after the oscillations stopped. Derefore, the hot-leg steam venting path did not change significantly when the oscillations stopped. Based on a pressure change of ( ]'6* psi in the upper head and the calculated core steam generation rate ([ ]'A' lb/sec.), it would require only a net accumulation of [ ]'** percent of the steam generated to provide this pressure rise in [ ]'** seconds (one-half the period, i.e., the time during a cycle when the pressure was increasing). De net steam accumulation in the reactor head was the difference between the steam generated less the steam vented through the ADS-4 lines and the steam flowing into the downcomer, it is more likely that this steam accumulation was a large fraction of the steam production based on the flow analysis of the ADS-4 line which is discussed later. A larger rate of steam accumulation would result in a much smaller period for the oscillations than that which was observed.
; This hypothesis considers only the level fluctuation about the hot-leg nozzle without considering the steam condensation in the downcomer. The resulting pressure rise would be very fast and the oscillation period would be very short, which is not consistent with the observed oscillation period.
Therefore, this hypothesis is not a probable cause of the flow oscillation. O mAap600G344w-64.nondb 100395 6.1.3-3 REVISION: 1
6.1.3.2.2 IIypothesis 2: Slug Flow in ADS-4 Lines his hypothesis holds that oscillations could be produced if slug flow was induced in the ADS-4 lines. When alternate slugs of vapor and liquid discharge into the separator, pressure pulses could be generated that would be propagated through the system. To investigate this hypothesis, several approaches were used. First, the flow regimes in the hot legs and ADS-4 lines were estimated using a flow pattern map (Figure 6.1.3-3) published by Baker." Based on the measured liquid flow from the ADS-4 separators, the calculated steam flow through the ADS-4 lines (calculated from the total steam generated in the core less the steam flow to the downcomer which was calculated from the pressure difference across the flow holes in the downcomer top), and the flow areas of the piping, flow in the [ ]'6* in. diameter hot leg and [ -
]'6' in. diameter ADS-4 line were in the stratified and slug regime, respectively (Figure 6.1.3-3). He calculated transport time through the hot leg and ADS-4 line based on an average volume for the steam / liquid mixture, is about [ ]'6# seconds.
Several analyses were performed to determine the flow characteristics in the ADS-4 lines during the long-term oscillations. Rese analyses are discussed below. Steam flow measured downstream of ADS-4 steam / water separator is shown in Figure 6.1.3-4. Figure 6.1.3-4 shows that, before [ ]'6' seconds, there was steam flow in the line. Between j'6' and [ J'6' seconds, there was no measured steam flow in the ADS-4 line because the [ core did not generate steam. After [ ]*6* seconds, the core did generate steam; however, Figure 6.1.3-4 shows that the steam flow rate in the ADS-4 line was very small, except for a few spikes with very short durations that may have been slugs of steam. Figure 6.1.3-5 shows that a spike of steam flow occurred at [ ]'6' seconds. He magnitude of the spike was smaller than one half of the peak flow rate in Figure 6.1.3-4. Dat is, the steam flow meter used to measure steam flow in the i ADS 4 line was sensitive enough to indicate the steam flow from the vessel. Since the magnitude of steam flow is only indicative because the vortex flow meter for the ADS-4 is overranged, it is possible that some slugs of steam have been undetected. Ilowever, the data analyses for transport time confirm that steam flow was intermittent and small, so that transport time was unaffected. He transport time of fluid in the ADS-4 line, computed by assuming that flow was single-phase liquid, agreed with the time lag between the flow oscillations in the vessel and downstream of the ADS-4 steam / water separator. He oscillation is a marker that can be used to trace fluid. That is, the time lag between flow oscillations in the vessel and downstream of the ADS-4 steam / water separators should be equal to the transport time of the fluid between these two locations. Since the vessel flow ! rate has been computed for steam but not for liquid, the steam flow rate will be used for the vessel, ! while the liquid flow rate will be used for the ADS-4 line in the oscillation time-lag calculation. Figure 6.1.3-19 shows the oscillation of steam flow in the vessel, and Figure 6.1.3-18 shows the oscillation of liquid flow of ADS-4. He time lag between the oscillations was about [ j'6' seconds. Conversely, assuming that flow in the ADS-4 line and the hot leg were single-phase liquid and that t'2 l mAap600cw64.noa:1b-too395 6.1.3-4 REVIs10N: 1 I,
- . _ . - -~ . -. - - - _ . . -- -. .- - _- .- - -.
Q fluid was travelling at the measured average flow rate of [ ]'6# lbm/sec., the transport time is calculated to be [ ]'6' seconds. His transport time is about the same as the observed time lag between the two oscillations which was about [ ]*6* seconds. His suggests that the hot leg flow was primarily single-phase liquid. Further evidence of this is that the level measurement (LDP-ll2) at the inner barrel wall above the core plate and the fluid temperature (TF-170) indicated this region was liquid-solid with an annular liquid region above the upper core plate. Therefore, this liquid is more likely to flow into the hot leg and the ADS-4 line. He above discussion suggests that the hot-leg flow contained primarily liquid and that an occasional vapor bubble flowed down the hot leg and out of the ADS-4 valve. Slug Flow
- It is extremely unlikely that the slugs of steam in the two ADS-4 lines would be exactly synchronized so that the period would be as regular as observed.
. De measured time lags for the ADS-4 oscillations agreed with the calculated time lag for single-phase liquid flow.
' . Two-phase flow pressure variations would be random and would occur at a shorter period than the observed [ ]'6' seconds. Liauld Flow i i
. If the flow through the ADS-4 lines were completely liquid, there would be no pressure pulses because of vapor / liquid slugs. In this case, the observed flow oscillations would have originated from pressure pulses driven by another source.
As a result of the above analysis, this hypothesis is not thought to be a probable mechanism for the observed oscillation. 6.1.3.2.3 Hypothesis 3: Downcomer Condensation / Fluctuations Under this hypothesis, a large fraction of the steam in the upper head was postulated to have flowed through the ten holes ([ ]'6# in. diameter) in the top of the downcomer and condensed on the surface of the coolant in the downcomer. Steam flow into the downcomer has been estimated by the three following methods:
- Pressure drop measurement (DP-130) across the downcomer top plate (i.e., across the flow
(,, holes) mAap60m23hWon:Ib-100395 6.1.3-5 REVISION: 1
i 1 i !
. Heat balance based on the temperature rise of flow through the downcomer
. Total steam generation assuming there was no steam flow through the ADS-4 line (based on previous discussion of transport analysis of flow in the ADS-4 line) l De results of these methods provide a range of estimated steam flow into the downcomer as summarized below: !
l Steam Flow to Downcomer ' Method (Ib/sec.) Pressure drop across downcomer top [ ]' 6
- Heat balance [ ]**
Core heat generation [ ]a 6
- i During IRWST injection, the temperature of the injection flow into the downcomer slowly increased with time (Figure 6.1.3-9) because the water in the IRWST had been heated first by operation of the PRHR and, as the test progressed, by steam released through the ADS 1-3 lines. Steam production in the core started at about [ ]** seconds (Figure 6.1.3-11). The existing superheated steam bubble ,
in the upper head gradually cooled by heat loss through the reactor head. When steam produced in the core mixed with superheated steam in the upper head, the steam temperature in the upper head began to decline. De start of this temperature reduction, which occurred at about [ ]** seconds, indicated the initiation of steam production in the core. Oscillations were observed to begin about [ ]** seconds after the initiation of steam production in the simulated core, herefore, the initiating event for the oscillations was not simply the onset of l boiling in the simulated core. However, the temperature difference between liquid (TF-167) (Figure 6.1.3-9) and steam downcomer vapor space decreased about [ ]** percent during the period from the initiation of steam production and the start of the osciliations. Oscillations began when the temperature of the downcomer liquid became so high that not all the steam flowing from the upper plenum to the downcomer could be condensed. He pressure then increased in the downcomer causing flow from the upper head to decrease. Since steam continued to be generated, pressure in the upper head increased, raising flow through the ADS-4 line and decreasing IRWST injection. Increased flow in the ADS-4 line reduced steam pressure which decreased steam flow to the downcomer. He level changes in the downcomer, because of these pressure changes also caused some mixing ofliquid at the steam / liquid interface, and enhanced the condensation rate. De enhanced condensation tended to further reduce pressure in the downcomer and upper head. As the steam-downcomer liquid interface increased in temperature from the collecting condensate, the condensation rate decreased below the rate of steam flowing into the downcomer, and pressure began to rise again. His cycle was repeated, producing the observed oscillations. maa;an2m64.noodb.too395 6.1.3-6 REVISION: 1
It is possible that a layer of saturated liquid formed on the liquid interface cassing condensation to stop completely. If so, the increase in steam saturation temperature resulting from the pessure increase, and mixing from the level change would tend to initiate condensation again. This process would result in the observed oscillating behavior. The oscillations ceased when the level in the downcomer decreased sufficiently that the cold leg was uncovered, and steam could be released through the cold leg instead of increasing pressure in the downcomer vapor volume and reactor upper head. Since the level changes required to release steam through the cold leg were too small to be determined from the level measurement, the presence of steam in the CMTs was used to indicate initial release of steam through the cold leg. Figure 6.1.3-8 shows a rise in CMT temperature slightly before the oscillations stopped, supports the uncover:ng of the cold leg as the termination mechanism for the oscillations. The uncovering of the cold-leg nozzles is also supported by evidence that the thermocouples at the top of the cold-leg nozzles reached saturation temperature, which indicated the presence of steam at the top of the cold legs. The downcomer condensation model is supported by the following observations: Both the start and the end of the oscillations were gradual. This is consistent with the slow increase in IRWST temperature at the onset of the oscillations and the gradual decrease in downcomer liquid level at the end of the oscillations. O = The oscillations were regular and smooth, which would result from changes in a process, such as condensation, where a very small driving temperature exists. Oscillations stopped when steam was vented through the cold leg to the CMT (and also through the break). The level and pressure oscillations were 180 degrees out of phase. The heat balance indicated that up to [ ]'6* lbm/sec. of steam was condensed in the downcomer. The estimated average condensation heat transfer coefficient for [ ]'b* lb/sec. of steam flow was about [ ]'** Bru/hr., ft.2, 'F, which is consistent with values for condensation of steam. This was a bounding estimate using the downcomer liquid temperature (TF-167), which is lower than the temperature of the liquid at the interface. This hypothesis was considered a valid mechanism in the Revision 0 of this report."* Based on l further investigations this hypothesis, which considers only the downcomer condensation alone, is not I considered a probable mechanism for the observed oscillations, although downcomer condensation is thought to be an important factor centributing to the oscillations. mn p600(2344w-64.non:tb-too395 6.1.3-7 REVISION: 1
6.1.3.2.4 Ilypothesis 4: Covering and Uncovering of Hot Leg with Downcomer Condensation To facilitate the explanation of this hypothesis for the flow oscillations, the idealized behaviors of the flow variables are schematically shown in Figure 6.1.3-7. Figure 6.1.3-7(a) shows the mixture level, L, in the upper plenum, where the top of the hot leg, nozzle Lut, is labeled. Figure 6.1.3-7(b) shows the total flow rate out of the hot leg, th gt. Figure 6.1.3-7(c) shows the steam generation rate from the core, rh,. Figure 6.1.3-7(d) shows the pressure in the upper plenum. Figure 6.1.3-7(e) shows the sum l of the injection flow rate from two DVI lines, rh,. Note that when the pressure in the vessel is low, l the pressure difference between the IRWST and the vessel is larger, resulting in a larger DVI injection flow rate. The large injection rate in turn causes the steam generation rate to be low. Thus, m, oscillates in phase with the pressure, P, while th, oscillates out of phase with P and rh,. Consider that initially the mixture level was at the hot-leg level so that the steam can vent through the hot leg. Hence, the pressure in the vessel was at a low value (point A in Figure 6.1.3-7(d). Because l of low pressure in the vessel, the injection flow rate was high, and the steam generation rate and the hot-leg flow rate were low. As a result, rh,j is larger than th at and the mixture level moved up. As the mixture level moved above the hot-leg level, the steam cannot be vented through the hot leg. Therefore, the pressure increased and the injection rate decreased. As the injection rate decreased, the steam generation rate increased, which further caused the pressure to increase. The increase of the pressure forced more fluid to go out of the hot leg and thus increased the hot leg flow rate, that. When the hot-leg flow rate became equal to the injection flow rate (point B) and continued to i increase, the mixture level started to decrease. When the mixture level decreased below the top of the I hot leg (point C in Figure 6.1.3-7(a), the steam vented through the hot leg and the pressure dropped l from point C to point D. As the pressure drops, the injection rate, thg, increased and the hot-leg flow rate, that, decreased so that thy became larger than th at. As a result, the mixture level increased. l When the mixture level increased past the top of the hot leg (point E), the hot leg is covered and the steam cannot leave the hot leg. Therefore, the pressure increased again and the cycle repeats. Note that in a real case, the pressure, P, does not drop instantaneously when the hot leg is uncovered (point C to point D in Figure 6.1.3-7(d), rather it decreases gradually as shown by the dashed line in j the Figure 6.1.3-7(d). Other variables, that, rh,, and rhy, also do not change abruptly, these variables change gradually as indicated by the dashed lines in Figures 6.1.3 7(b), (c), and (e). Also note that l when the hot leg is covered, only a small amount of generated steam is available to pressurize the vessel. Most of the steam goes through the upper head bypass holes to the top of the downcomer and is condensed there. Detailed calculations and comparisons with the data are given in Subsection 6.1.3.3. In comparing the level data, keep in mind that the data represents the collapsed liquid level, not the two phase mixture level, since the mixture level was not measured. The collapsed liquid level is generally lower than the m Aap600c344w-M.non:lb.100395 6.1.3-8 REVISION: 1
O mixture level. The calculations show that the high hot leg flow rate, thgt, during the hot leg covering period is single-phase liquid flow. Since the hot-leg flow rate is not measured, the liquid flow rate downstream of the ADS-4 steam / water separator is used, which has a time lag of about [ ]** seconds due to transportation time from the entrance to the hot leg to the downstream location of the ADS-4 steam / water separator. 6.1.3.2.5 Hypothesis 5: Covering and Uncovering of Upper Head Bypass Holes When the mixture level is always above the hot-leg nozzle so that the hot-leg nozzle is never uncovered, Hypothesis 4 does not apply. In this case, flow oscillations may be due to the alternate, covering and uncovering of the upper head bypass holes. Initially, the mixture level is above the top of the upper head bypass holes. Since the bubbles can only move upward, the bubbles cannot go to the bypass holes. Herefore, only water will go to the bypass holes, that is, the bypass holes are covered with water. With the water covering the bypass holes, the steam cannot go through the bypass holes easily. Therefore, the pressure in the vessel increases. As the pressure increases, the DVI injection rate, rhy, decreases and the hot-leg flow rate, rtigt, increases. The decrease thyin turn causes the steam generation rate to increase, which causes the pressure to further increase. When th at becomes larger than thy, the mixture level drops. As the i mixture level drops, the bypass holes are uncovered. As the bypass holes are uncovered, the pressure drops. The decrease in the pressure causes the injection rate to increase and the hot leg flow rate to decrease. The increase of the injection rate in turn causes the steam generation rate to decrease, which causes the pressure to further decrease. As the injection rate becomes larger than the hot-leg flow rate, the mixture level rises. When the mixture level rises above the top of the bypass holes, the bypass holes are covered with water and the steam cannot go through the bypass holes easily. Therefore, the pressure increases. Then, the flow conditions return to the hiitial state and the cycle repeats. When the bypass holes are covered, pressurization of the upper head is very fast because of the high steam flow rate in comparison with the small upper head volume as discussed in Subsection 6.1.3.3. High pressure will cause the mixture level in the upper head to drop quickly, which will cause the bypass holes to be covered in a short time. As a result, the predicted oscillation period is much shorter than the observed oscillation period. Thus, this hypothesis is not a probable cause for the observed flow oscillations. 6.1.3.3 Examination of the Test Data to Evaluate the Postulated Oscillation Mechanism As discussed in Section 6.1.3, flow, pressure, level and temperature oscillations have been observed in most of the OSU tests. Hypothesis 4, described in Subsection 6.1.3.2, is the one consistent with all the tests. The oscillation process is described in detail for Test SB01, which is the reference transient. The other tests are discussed briefly because of similarities with Test SB01. I ewmmwanon:1b 100395 6.1.3-9 REVis!ON: 1
1 l l l 6.1.3.3.1 Oscillations in Test SB01 ; During IRWST injection, the temperature of the injection flow into the downcomer slowly increases (Figure 6.1.3-9). De injection flow rate also decreases because the IRWST liquid level decreases (Figure 6.1.3-10). At about [ ]'6# seconds, steam production in the core starts (Figure 6.1.3-11) and the upper plenum collapsed liquid level starts to drop (Figure 6.1.312). A small amount of steam generated in the core accumulated in the upper plenum, but most ofit was discharged from the upper head through the bypass holes into the downcomer where it condensed with the cold water. The ADS-4 high flow rate and subcooled temperature (Figure 6.1.3-13 and 6.1.3-14) indicate that during this time period (from 8000 to [ ]'6# seconds), the discharge is essentially pure liquid. At about l [ l'6* seconds, the upper plenum collapsed liquid level reaches the top level of the hot-leg nozzles l (Figure 6.1.3-12) and steam starts to a discharge through the ADS-4 valves (Figure 6.1.3-13) and the oscillations start. i In the following, one period of oscillation corresponding to the time interval from [ ]' 6
- to
[ ]'6# seconds, starting when the upper plenum collapsed liquid level is at its minimum, is j analyzed. Conditions at [ l*6* seconds (Figures 6.1.3-15 to 6.1.3-19) are:
- Upper plenum collapsed liquid level = [ ]'6# in.
- Upper plenum pressure = [ ]'*# psia
. Total injection flow rate is = [ ]'6# lbm/sec.
- Total ADS-4 discharge = [ ]'6# lbm/sec.
. Steam generation in the core = [ ]'6# lbm/sec.
The hot-leg nozzle is partially uncovered and steam is discharged easily through ADS-4. He pressure decreases consistently as injection increases and steam generation decreases as a result ofincreased injection. At about [ ]*6# seconds, the ADS-4 discharge starts to decrease when the pressure in the upper plenum is at [ ]'6# psia. At [ ]'6# seconds, injection from the DVIs is higher than the discharge from ADS-4 and steam provided by the core reaches its minimum value ([ ]'6* lbm/sec.); thus, the level in the upper plenum recovers. At about [ ]'6# seconds, the upper plenum collapsed liquid level reaches its maximum ([ ]'b'in.) and completely covers the hot-leg nozzles. De steam flow rate to ADS-4 is prevented and the steam generated is preferentially discharged into the upper head. He pressure in the upper plenum and upper head, which reaches its minimum ([ ]'6# psia) at about [ ]'6# seconds starts to increase as a consequence of the reduction of the steam flow to ADS-4. De increase of upper plenum pressure is responsible for the decrease of the injection rate and as a result, steam generation rate increases. He generated steam flows into the upper head and is discharged into the downcomer where it condenses with the cold liquid. Steam flow from the upper mAap600cmw.6tnon:1b.ioows 6.1.3-10 REVISION: 1
_ _ -- . . . .- ~. - - - . - . - . . - . - . - - . . . l A l head to the downcomer is driven by the pressure difference between the upper head and downcomer ) (Figure 6.1.3-20). The upper head pressure oscillation amplitude is [ ]'b* psi. Pressurization of the upper head / plenum is due to the difference between the steam generation rate in the core and the condensation rate in the downcomer when the hot-leg and cold-leg nozzles are covered. It takes only a small amount of steam accumulation to pressurize the upper head / plenum. De upper head / plenum
)
is pressurized when the downcomer cannot condense all the generated steam. When the cold leg is uncovered, the generated steam flows through the cold legs, then the upper head / plenum cannot be pressurized and the oscillations stop. At about [ ]'* seconds, the pressure reaches [ ] psia. He pressure in the upper plenum is , higher than the gravity head of the liquid column in the ADS-4 line ([ ]'6* psia). The ' upper plenum and hot leg liquid is pushed into the ADS-4 lines. His is consistent with a sharp increase in the ADS-4 discharge as soon as the liquid reaches the ADS valves after a short time lag. He upper plenum collapsed liquid level then uncovers the hot-leg nozzles again and reaches its minimum at about [ ]'* seconds. Steam provided by the core is again vented to the ADS-4 lines and the upper plenum pressure starts to decrease; injection increases and the cycle is repeated. t The period of oscillation is about [ ]** seconds and is consistent with a global mass balance, where periodically the water injected ([ ]'* lbm) is at first accumulated in the upper plenum, hot legs, surge line and SGs inlet plenums, then is rapidly discharged through the ADS-4 valves. This is consistent with the upper plenum collapsed liquid level being out of phase with the ADS-4 mass flow rate. At about [ ]'" seconds, the injection temperature has been increased and the injection flow rate has slowly decreased. He average steam generation rate in the core increased slightly. At this time, I the thermocouples at the top of the cold-leg nozzles have reached saturation temperature ! (Figure 6.1.3-6), which indicates the presence of steam at the top of the cold legs. Therefore, the cold-leg nozzle was uncovered and steam was vented from the downcomer to the cold legs preventing the upper head / plenum from pressurization. Consequently, the flow oscillations stopped. 6.1.3.3.1.1 The Quasi steady State Before Oscillations Start A quasi-steady state is established before the oscillations occur (at [ ] seconds). Cold water is provided by injection, no steam is generated and warmer water is discharged by the ADS-4 valves and the break. At this time, both cold-leg and hot-leg nozzles are covered. O maap600c34w-64non:1bloo395 6.1.3-11 REVISION: 1
The upper plenum pressure which is able to maintain this quasi-steady state (Figure 6.1.3-21) can be calculated as folitws: The measured total injection rate th, is the sum of two DVI flow rates, thovi,i and shovt-2 ; th,,, = rh ovi.: + rh ovi2 =[ ]'** lbtn/sec. 6.1.3-1 The measured total discharge flow rate rhm, is the sum of the flow rates from ADS 4-1, ADS 4-2 ano b:eak thAos. 4-i, thAos 4-2, and rh,,, respectively: Iha + IhADS 4-2 + Ih Brd ,[ ]b,clbm/seC. n " hl ADS. 4-1 6.1.3-2 Note that the fluid discharged is subcooled liquid and no steam is generated at this time. From these quantities, the pressure drops across the ADS-4 valves can be evaluated in the following. The measured mass flow rate rh in ADS 4-1 is [ ]'** lbm/sec. ([ ]'A' kg/sec.). Thus from the ADS 4-1 orifice diameter ([ ]'** mm) and the liquid density p, the velocity V inside the orifice can be evaluated: Orifice flow area = A = [ ]'A' 6.1.3-3 V = th/(p A) = [ ] ]'A' Me M34 The differential pressure, APi nsgi, across the valve ADS 4-1 is then evaluated from the orifice K C factor. The orifice K factor is evaluated from Idelchick Manual O and is equal to 1.1. APios,4.i = KpV2 /2 = [ ]'** psia 6.1.3-5 For ADS 4-2, the measured mass flow rate is 0.8 lbm/sec. (0.363 kg/sec.) and the discharge flow area A is double [ ]'**. The velocity, V, is then: V = th/(p A) = [ ] ]'A' m/sec. 6.1.3-6 and the differential pressure, AP40s, 2, across the valve ADS 4-2 is: APAos.42 = KpV 2/2 = (with K=[ ]'AS = [ ]'6* psia 6.1.3 7 O mAap600Q344w-64.non:lb-100395 6.1.3-12 REVISION: 1
The average differential pressure, AP30s,4, of ADS 4-1 and ADS 4-2 is then: - AP,os,4 = [ ]"# psia 6.13-8 he upper plenum pressure at the hot-leg nozzle elevation can be evaluated assuming that the distributed pressure drops are negligible. 4 he gravity head AP, of the ADS line (full of liquid) is: (h=( ['#'m) AP, = p g h = [ ]"' psia 6.13-9 De collapsed liquid level in the upper plenum is [ }"# in above the hot-leg elevation which corresponds to a gravity head, APgg, of [ ]"' psia. Bus, the pressure of the upper plenum in the steam volume is: 3 l l Pup = P. + AP ro s,4 + AP, - AP s.ur = [ 6.13 10 l"# Psia , where P is the atmospherical pressure. He oscillations start some time after steam is generated in l the core. i ! D j j3 6.133.1.2 Steam Generation and the Onset of the Oscillations he oscillations start at [ ' ]'6* seconds and the steam generated by the core (rh,) is: 4 rh,=[ ]'b# lbm/sec. 6.13 11 ] 4 i After about [ ] seconds, the generated steam oscillates between [ ]"' and [ ]*'lbm/sec.with an average value of about [ ]"' lbm/sec. l When the collapsed liquid level in the upper plenum covers the hot-leg nozzles and the ADS-4 steam
- flow is essentially prevented, the steam flows into the upper head and from the upper head to the
- downcomer where it condenses with the subcooled IRWST injection. When the hot leg nozzles are uncovered, the steam generated is vented through the ADS-4 valves.
The oscillation of the steam generation rate is a result of the injection flow rate oscillation. When the hot-leg nozzles are covered and the upper plenum is pressurizing, injection is decreasing, and r consistently the steam generation rate is increasing. The steam generation rate reaches its maximum , just before the hot-leg nozzles are uncovered, then as steam is released through ADS-4 and the pressure decreases, injection increases and consequently, steam generation rate decreases. mAap600s2%44.noa:tb-ioo395 6.13-13 REVIs10N: 1
In conclusion, the steam generation rate is out of phase with the injection flow rate because core power is constant. The maximum injection rate ([ ]lbm/sec.) corresponds to the minimum steam generation rate ([ ]'6# lbm/sec.) and the minimum injection rate ([ ] lbm/sec.) corresponds the mar.imum steam generation rate ([ ]'6# lbm/sec.). The oscillation phenomena is subdivided into the following phases:
. Downcomer steam condensation when hot leg is covered [ ]'6# seconds
. Upper head pressurization and upper plenum draining [ ]'6' seconds
. Upper head depressurization [ ]# seconds
. Upper plenum level recovery [ ] seconds These phases are described in the following.
6.133.13 Steam Condensation The steam condenses and the saturated condensate liquid layer collects at the top of the downcomer liquid level. As the thickness of the condensate increases, the condensation rate decreases. After while, if mixing does not occur, condensation stops because the saturated layer thickens and the steam liquid interface becomes insulated from the subcooled water. Only conduction through the condensate layer cools the interface, but the conduction process becomes less and less effective as the condensate layer develops. The condensation rate can be calculated by the following basic equation where the mixing effect into the downcomer is considered negligible. Assuming no-mixing occurs at the interface, the condensation rate can be evaluated based on the conduction of the liquid across the interface layer as the only significant heat transfer resistance: I The interfacial heat exchange coefficient, h, is: h = k; / d 6.13-12 where: ki= liquid conductivity d = condensate layer thickness The condensate layer thickness is related to the steam flow th, with the equation: d = rh,t/(pf A) 6.1 3-13 O m:w344w-64.noo:1b.ioo395 6.13-14 REVIs!ON: 1
where: p, = Liquid density A = Downcomer flow area = steam-liquid interface area = [ ] ft.2 t = Time (assuming d = 0 at t=0) The basic equation of heat transfer notes: heat exchange at the interface = steam condensation rate h A AT = rh,h,, 6.1.3-14 th, is obtained from the previous three equations (6.1.3-12), 6.1.3-13), and (6.1.3-14): rh, = (k,p,A2 AT/h r ,)"/t" 6.1.3 15 Pressure =[ ]'6d psia gives: k, = [ ]'6# Btu /ft. hr.*F 6.1.3-16 p, = [ ]' 6
- lbdft.' 6.1.3-17 h,, = [ ] Btu /lbm 6.1.3-18 AT = T. - Tw = 212 - [ - ] = [ ]'6" 'F 6.1.3-19
'Thus; di,=[ l'A*/t" 6.1.3-20 t=[ ]'6# sec. th,=[ ] lbm/sec. 6.1.3-21 t=[ ]'6# sec. th,=[ ]'6# lbm/sec. 6.1.3-22 t=[ ]'6d sec. rh,=[ ] lbm/sec. 6.1.3-23 t=[ ]'6# sec. th,=[ ]'6d Ibm /sec. 6,1.3-24 From these calculations, the interfacial heat transfer in the downcomer condenses all the steam generated in the core during the first [ - J'6# seconds. At [ ]# seconds, the condensation capability is reduced by a factor [ ]# and only [ ]'6d percent of tne steam generated is condensed. In the test, the hot-leg nozzles are covered (the upper plenum collapsed liquid level is greater than 11 in.) for about [ ]'6* seconds.
b 'd m:W344w-64.noa:t b.100395 6.1.3 15 REVISION: 1
Considering the integral over [ ]"' seconds: kom f 0 th,dt 6.1.3-25 The average flow rate, rh is: E y=[ ] lbm/sec. 6.1.3-26 Ris is only about [ ]**' percent of the steam generated in the same time period. De observed upper head, downcomer and upper plenum pressurization ([ J'6* psia)in h se une period is due to about [ ]'6'lbm of accumulated steam at constant volume. During pressurization, the steam volume in the upper plenum expands to about [ j ft. and this corresponds to an accumulation of [ ]lbms. The total accumulation is [ ]'6' lbm, which corresponds to an average mass flow rate in the same time period of [ ] lb/sec., which is only [ ]'6" percent of the steam generation rate. It is concluded that mixing occurring in the downcomer increases the steam condensation rate by at least a factor of [ j'6* . He mixing effect reduces the layer thickness and presumably during the time period ([ ] seconds) when the hot-leg nozzles are covered, all the steam generated is condensed. His implies that the average layer thickness during these [ ]'b* seconds is: d = k,A AT/(th,h,,) = [ ]'be=[ Jtb' ft. = [ ] in. 6.1.3-27 1
~
Assuming that all condensate is collected at the top without any mixing, the layer thickness should have been: d = rh,t/(p,A) = [ ] ]*6' in. 6.1.3-28 After [ J'b' seconds, the layer thickness is high enough (also with the mixing) to shut off the condensation. De imbalance between steam generated and condensed pressurizes the downcomer, upper head and upper plenum. O mMescocm44.non:ib.icm95 6.1.3 16 REVIslON: 1
6.1.3.3.1.4 Upper Head pressurization and Upper Plenum draining - He upper plenum draining follows the upper head pressurization. He water accumulated in upper head, hot leg, SG plenums and pressurizer surge line is discharged out of the system through the ADS 41 and 4-2. In the actual test, there is a complex feedback between the upper plenum pressurization, the upper plenum level and the steam generation. He upper head pressure increases because of the steam provided by the core. As the pressure increases, the upper plenum level decreases which means that a larger volume is available for the steam expansion. Conversely as the pressure increases, inbetion decreases and as a result, the steam generation rate increases. Rus, more Neam is provided to the upper uad during pressurization. He steam volume provided by the core i much greater than the new steam volame evallable as a consequence of the upper plenum draining. He result is that the pressure in the upper nead increases during the upper plenum draining, in order to evaluate the maximum flow rate from the upper plenum to t' e ADS-4 an instantaneaus pressurization without the upper plenum level decreasing, followed by tne upper plenum drainlag at constant pressure is assured. When the pressure in the upper plenum reaches [ ]'6' psia and the upper plenum collapsed liquid level at [ ]'6" in.: AP,(upper plenum from [ ]'6* ia to [ ]'6
- in] = [ ]'6' psia 6.1.3-29 (Pressure upstream of ADS 4-1 or ADS 4-2 valves) = [ ]'*#+[ ]'A' - 1.12 = [ ]'6' psia 6.1.3-30 Ris means that the differential pressure across the ADS-4 valves is:
APios,, = [ ]' 6 ' - 14.'l = [ ]'6
- psia 6.1.3-31 From this value, the mass flow rates th,33,,,iandrt h 42 can be evaluated with the assumption that until the hot-leg nozzles are uncovered, no steam is discharged by the ADS-4 valves. In this case, the ADS-4 valves and lines pressure drops are single-phase liquid:
The new ADS 4-1 and ADS 4-2 mass flow rate are evaluated using the previously (old) calculated values for the quasi-steady state before the oscillations start: rig ,,, = [(AP,/AP,)thilas , [ ja5jts' = 0.9 lbm/sec. 6.1.3-32 rth 4-2= [(APJAP,)thil*5 = [ ]*5] = 1.5 lbm/sec. 6.1.3-33 O V maap60m234sw-unon:1b.too395 6.1.3-17 REVISION: 1
1 The total discharge is [ ]'6* IWa Ms value is very close to the maximum measured value, i With this flow rate, the ADS-4 valves discharge [ ]*** lbm of liquid in about [ l'6# seconds. The l [ ]*6* lbm is the mass accumulated in the system in one period of oscillation (measured value): l [ j'6# lbm is in the upper plenum, [ l'6* lbm in the pressurizer surge line, [ ]'** in the SG plenums and [ ]'6" lbm in the hot legs. In the test, the ADS-4 total flow rate is about [ ]'6# lbm/sec. for about [ ]'*# seconds, then it l decreases to [ ]'*# in about [ ]'*# seconds. An average of [ ]'6# lbm is discharged in [ ]'6" seconds. Note that the ADS-4 maximum flow rate occurs about [ ]'*# seconds later than j when the upper plenum starts to drain. This is the transportation time from the upper plenum to the ADS-4 valve as evaluated in Subsection 6.1.3.2. In the test, the upper plenum pressure can be considered constant only during the first [ ]**# seconds of ADS-4 discharge, but then the pressure decreases because the hot-leg nozzles uncover and steam is vented with the liquid throug!, the ADS-4 valves. Thus, after the first [ j'*# seconds of ADS-4 high flow rate (pure-liquid) discharge, the two-phase ccatribution to the ADS-4 pressure drops evaluation also has to be considered. In order to quantify this contribution, it is assumed that during the ADS-4 discharge, all the steam provided by the core (dl, = [ ]'6# lbm/sec.) was discharged together with the liquid. For example, the ADS 4-1 calculation is iterative: O th,=[ ]'*# lbm/sec. 6.1.3-34 The total mass flowrate, rty, is the sum of steam flowrate, rh,, and liquid flowrate th,; dy = rh, + di, 6.1.3 35 By definition the quality, x, is: X = th/ tty 6.1.3-36 Assume di, = [ ]'** lbm/sec. 6.1.3-37 i Quality = X = [ ]'*#% 6.1.3-38 O msp60cc3manon:1b-ioo395 6.1.3-18 REVISION: 1 i
1 l d The two-phase ADS-4 valves pressure drops (across the orifice), AP 3 , are evaluated utilizing the formula: AP3 = AP 3 ,,$' 6.1.3-39 1 where AP3, is the single-phase liquid pressure drop considering all fluid to be liquid and $2 is the l two phase pressure drop multiplier evaluated with the Griffith correlationa2> as: 2
$ = 1 + CX(p, / p,) 6.1.3-40 l
1 l 'Ihe constant C for orifices is 0.8 ! i p/p, = 1575 at P=15 psia 6.1.3-41 l where p,is the liquid density and p, is the vapor density. I i 2
$ = 37.8 at P=15 psia and X=[ ]'6* % 6.1.3-42 l l
D The single phase differential pressure across the valve is evaluated by rationing the previously calculated pressure drop AP,a (=[ ]'6# psia) corresponding to a liquid flow rate th,g of [ ]'6* lbWa AP3 , = (rh, / th,y2AP,g = [ ]'6# psia 6.1.3-43 where: rh , a rh,, i Therefore, from equation (6.1.3-39) AP3=[ ]'6" (37.8) = [ ]'6' psia 6.1.3-44 which is too large. Thus, the estimated [ ]'b'lbm/sec. is not compatible with the available pressure drop. Assume th, = [ ]'b' lbm/sec. 6.1.3-45 rh,=[ ]' 6" 6.1.3-46 m:\ap600\2344w-64.non:Ib 100395 6.1.3-19 REVISION: 1
AP,,, , = [ ]** psia 6.1.3-47 X=[ ]*'%, P=[ ]*' psia ==> $ = [ 2
]** 6.1.3-48 AP2 , = [ ]
- Psi 6.1.3-49 Assuming slip ratio = 1, the void fraction is evaluated as follows:
(Density ratio p/p ) = R = [ ]** at [ ]** psia 6.1.3-50 By definition, the void fraction, n, is: a = 1/[ ] = 1/ [ ]**)/[ 1**)([ ]*')]=[ ]** 6.1.3-51 Therefore the elevation pressure drop, AP,, line is: AP,, % = [ ]** psia 6.1.3-52 and the pressure in the upper plenum P,,is: Pur = [ l** Psia 6.1.3-53 O which is the initially assumed pressure in the upper plenum. Rus, with the pressure drop available to discharge all the steam generated by the core ([ ]'6* lb/sec.), the total flow from ADS-4 is limited below [ ]** lbm/sec. Note that the assumption of slip ratio = 1.0 tends to underestimate the gravity head of the ADS line. Generally, it will be slip ratio >> 1.0 resulting in a lower void fraction in die ADS line and consequently a higher elevation pressure drops AP,,w. He ADS-line friction losses would tend to slightly increase, the total pressure drops, but these were not considered. In conclusion, the presence of steam limits the ADS-4 flow rate and it can be inferred that during the high ADS-4 flow rate phase, only pure liquid can be discharged by the ADS-4 valves. The steam can be discharged only whcn the steam path to the ADS-4 valves is open as the hot-leg nozzles uncover. When this occurs, the liquid flow rate is very low. 6.1.3.3.1.5 Upper Head Depressurization (Low ADS-4 Mass Flow Rate Phase) When the level in the upper ple lum drops below 11 in., the hot-leg nozzles uncover and ADS 4-1 and 4-2 start to vent steam. The steam flow from the upper plenum to the downcomer reduces significantly. O m:w6m2344w-64.non:ib-too395 6.1.3-20 REVISION: 1 1
In this case, the hot-leg and upper plenum pressure decrease to the minimum 'value: l Pressure upstream ADS-4 = 14.7 + (AP 2 ,in the ADS line) 6.1.3-54 AP2 , in the ADS line is the sum of the pressure drops across the valve AP 2,, valve and the pressure drop:2 in the line AP24 , line. l AP2 , = AP 2,, ,,,, + AP 2,, in. 6.1.3-55 l 2 AP2 ,. vsv. = AP ,,, $ 3 6.1.3-56 where: I 2
$ = Two-phase flow pressure drop multiplier Two situations are considered:
a) Pure steam discharge h, All steam generated in the core is discharged by the ADS-4 valves without entrainment ofliquid from V the hot legs. rty = rh, = [ ]** lbm/sec. and X=[ ]**% 6.1.3-57
$' = [ ]** at X=100%, P=15 psia 6.1.3-58 The single phase differential pressure (liquid) across the valve, AP ,,,. is evaluated by rationing the 3
previously calculated (old) value corresponding to a liquid mass flow rate of [ ]** lbm/sec.: Ap3,,, = (rigfrtQ2AP, = [ ]** psia 6.1.3 59 when th, = tty 2 AP2 , vsv. = AP ,, , $ = [ 3
]** psia 6.1.3-60 D
(V m:W344 -64.non:ib-ioo395 6.1.3-21 REVISION: 1
ne pressure losses and the gravity head AP,,% of the ADS line are negligible: AP,,% 2 = AP%, % + AP,,% = 0. (the line is pure steam) 0.1.3-61 AP:, = [ ]'6# + 0.0 = [ ]'6# psia 6.1.3-62 b) Quality [ ]** percent discharge i The slip ratio can be assumed to be 1.0 to simplify the calculation: 1 rh, = th, + rb, = thy /X = [ ] = [ ]'*# lbm/sec. 6.1.3-63 l The single phase (liquid) differential pressure across the valve is: l AP ,, , = (thdrh,i,)2AP,,, = [ ]'6* psia 6.1.3-64 3 l l l 2
$=[ ]' *# at X = [ ]**% and P = [ ]*' psia 6.1.3-65 AP2 ,. ,s . = I l6 Psia 6.1.3-66 1
The ADS line void fraction is evaluated assuming slip =1 from the equation
- I a = 1/ [ ] 6.1.3-67 l where p,/p,is the density ratio which is [ ]# at the atmospheric pressure:
a=[ ] ]** 6.1.3-68 AP2 ,. w = AP,, % = [ ]' 6# 6.1.3-69 AP2 , = [ l** Psia 6.1.3-70 Finally, the upper plenum pressure oscillation minimum in the two situations is evaluate # In case (a) P,p = 14.7 + [ ]' 6 ' = [ ]'6* psia 6.1.3-71 In case (b) P,, = 14.7 + [ ]'6#=[ ]** psia 6.1.3-72 O mvan3u+u.wn:1b-too395 6.1.3-22 REVIs!ON: 1
. - .- -. .--_. - - ~.. -. _ - - . _ - - - -.. _ _ - -- ._ .
i l l i he oscillation amplitude (the maximum pressure is [ ]*** psia) is then: , (a) P,,, , - P,,, , = [ ]
- psia 6.13-73
- (b) P,,, , - P,,, , = [ ]'** psia 6.1 3-74 The measured minimum pressure is about [ ]' 6d psia and the pressure oscillation amplitude is about
[ ]'** psia. His is consistent v/ith tr.e calculated value if the quality is about [ ]'** percent. In this case, the ADS 4-1 flow rate is [ ]#' lbm/sec. and ADS-4-2 is estimated to be about [ ]'** lbm/sec. The total ADS-4 fLw is about [ ]"#, which is less than the average injection ([ ]'b# lbm/sec.L 'he imbalance a:4 flow results in the level recovery. l 6.133.1/ upper Plenum Level Recovery (End of One Cycle) l As the upper plenum pressere and level oscillate, the total injection oscillates as well. Re amplitude . can be considered small (+/- [ ]*d percent of the mean value) and the injection can be considered I fairly constant in comparison with the discharge. The average injection is abon [ J'** lbm/sec. To recover the [ ]**'lbm discharged during the upper plenum draining phase requires [ ]'** secorlds, which is the measured oscillation period. As soon as the hot leg nozzles sre recovered, the generated steam again flows into the upper head. Meanwhile, the condensate layer in the downcomer is mixed sufficiently for the interface to be subcooled again. Condeissauon c .ctcs again and the cycle is repeated. 6.133.2 Oscillations in the Other Tests Osclilations in tests: SB06, SB09, SB10, SB12, SB13, SB14, SB15, SB18, SB19, SB21 and SB23 are included in the following discussion. These tests are grouped into subsets, which show strong similadties in the oscillation characteristics. The following subdivision also emphasizes the break location. His grouping shows that the oscillation parameters are more strongly related to the break location than to the break size. t With Figure 6.13-22 as the starting point for each test, the following set of seven plots is presented:
- 1. Total DVIs injection mass flowrate (DVl-1 + DVI-2)
- 2. DVI nozzles liquid temperature (injection temperature)
Q 3. Liquid temperature in the downcomer at the elevation of the hot leg nozzles V m:Whw.noa:1b.noo395 6.13-23 REVISION: 1
- 4. 'lotal ADS-4 mass flowrate (ADS 4-1 plus ADS 4-2)
- 5. Core steam generation rate
- 6. Upper plenum collapsed liquid level
- 7. Upper head pressure 6.1.3.3.2.1 Subset 1 - SB18, SB19, SB23, SB06, SB21 - Cold Leg Break Test SB18 (Figures 6.1.3-71 to 6.1.3-77) was a 2-in. cold leg break (a duplicate of SB01); SB19 (Figures 6.1.3-78 to 6.1.3-84) was a 2-in. cold-leg break with a backpressure; SB23 (Figures 6.1.3-94 to 6.1.3-100) was a 0.5-in. cold leg break, SB06 (Figures 6.1.3 22 to 6.1.3-28) was a 4-in. cold-leg break and SB21 (Figures 6.1.-85 to 6.1.3-93) was a double 4-in. cold-leg break.
In SBl8, steam generation started at about [ ]'6# seconds and the oscillations started at about [ ]'6" seconds when the steam generation reached [ Jlbm/sec which is consistent with the calculated steam generation at the start of the oscillations in Test SB01. The average oscillation period in the time interval [ ]'6# seconds was about [ ]'6d seconds. In Test SB01, the oscillation period was higher (about [ ]'6" seconds) and this is consistent with a [ ]'6# 'F lower, top downcomer temperature measured in SB01. The lower the temperature in the downcomer, the higher the condensation rate and the longer it takes to develop a saturated condensate liquid layer, which stops condensation and starts the pressurization of the upper head. The oscilladons in Test SB19 are very similar to the ones in SB18. Steam generation started at about the same time ([ ]'6' seconds) and the oscillations stared at about [ ]'6" seconds when the steam generation reached about [ ]lbm/sec. The temperature at the top of the downcomer was very close to SB18. The average oscillation period, in the time interval [ l'6' seconds is [ ]# seconds, was very close to the value of SB18. In Test SB23, steam generation started at about [ ] seconds and the oscillations started at about [ ]'6# seconds when the steam generation reached [ l'6"lbm/sec. The oscillations are not as well defined as in SB18 and in the time interval [ ]'6' seconds, the average period is about [ ]'6d seconds. In Test SB06, the calculated core steam generation rate RPVASOU2 is presented instead of RPVASOUT. RPVASOU2 considers the net core flow (the difference between the total DVI flow and the DVI to Cold leg break bypass flow). In SB06, the DVI to break by-pass flow is higher and l consequently, the core flow rate was lower during the IRWST injection. In this case, the predicted core steam generation rate after [ ]'6# seconds was about [ J'6# lbm/sec. As a result, th; oscillations started very early in the transient at about [ ]'6" seconds. The average oscillation period is [ ]'6# seconds, which is shorter than in SB01, maap60m2h64 noa:15too395 6.1.3-24 REVISION: 1
O in Test SB21, the oscillations started at about 7200 seconds with an oscillation period of about [ ]'6' seconds. In SB21 the upper plenum collapsed hquid level (CLDP-ll3 in Figure 6.1.3-91) is not operable and the total vessel collapsed liquid level (CLDP-127 in Figure 6.13-90) is utilized instead. However the vessel collapsed liquid level does not represent the true collapsed liquid level, because of the presence of the lower core plate, the upper core plate and the upper support plate, which introduce significant pressure drops. Derefore the collapsed liquid level overestimates the vessel level. 6.13 3.2.2 Subset 2 - SB12, SB13 - Direct Vessel Injection Line Breaks Tests SB12 and SB13 (Figures 6.13-43 to 6.13-56) are the DVI breaks (DEG and 2-in., respectively). No oscillations were observed in the DEG DVI (SB12) as a result of the break location and size. He break kept a low inventory of mass in the system and the upper plenum collapsed liquid level was always below the top of the hot-leg nozzles so that the steam core always vented to ADS-4. De mechanism for oscillation cannot occur if the upper plenum level is lower and steam vents out of the hot leg. Steam generation starts at about [ l'6" seconds, and is vented through ADS-4. The upper head pressurization does not occur. In the 2-in. DVI break (SD13), steam generation occurs through the entire transient. He oscillations start at about [ ]'6" seconds when the steam generation rate is [ j'*# lbm/sec. He fact that steam generation is higher is a consequence of the partially lost injection flow though the break in DVI since the lower the injection, the higher the steam generation rate. He average period of oscillation in the time period between [ ]'*# and [ ]'6# - unds is [ ]'*# seconds. 6.1.33.2.3 Subset 3 SB15 - Hot-Leg Break Test SB15 (Figures 6.13 14 to 6.13 70) was a 2-in. hot-leg break. Steam generation started at about [ ]'*d seconds and the oscillations started at about [ ]'*# seconds when the steam generation rate was about [ l'** lbm/sec. He average period of oscillation was [ ]'*# seconds in the time interval [ ]'*# seconds, and the period is reducing during the transient: between [
] seconds is [ ]'** seconds and between [ ]'*# seconds the period is
[ ]'*# seconds. De reason that the oscillation period is smaller in SB15 than in the other tests is 4 because there is an additional flow path through the hot-leg break so the hot-leg nozzle can uncover faster than in the other tests. 6.1.33.2.4 Subset 4 - SB09, SB10 Cold Leg / Core Makeup Tank Balance Line Break Tests SB09 and SB10 (Figures 6.13-23 to 6.13-42) were the 2-in. cold-leg balance line break and DEG cold-leg balance line break. He system response was very similar during the oscillations. Steam started to generate at about [ ]'*# seconds in SB09 and [ ]'*# seconds in SB10. The oscillations started at about [ ]'6" seconds in SP N and [ ]'*d seconds in SB10 when the O' steam generation rate was [ J'*# lbm/sec. The average oscillation period is [ ]'*# seconds in maap600c344w-64.noa:ib-ioo395 6.1 3-25 REVISION: 1
l l SB09 and [ ] seconds in SB10. The steam generation rate at the onset of the oscillations was higher in Tests SB09 and SB10 ([ ]'6' lbm/sec.) than for the tests of group 1 ([ ]dd lbm/sec., SB18, SB19, and SB23). The difference is that the tests in group-l are bottom cold-leg breaks while l the tests of group-4 are top cold-leg breaks. In this case, some steam was vented by the break, and i downcomer pressmization was delayed. In both cases, the downcomer collapsed liquid level was l around the top of the cold-leg nozzles in both SB09 and SBIO: since a break is at the top of the cold leg, some steam was collected at the top of the cold leg and was discharged by the break in both SB09 and SB10. I 6.1.3.3.2.5 Subset 5 - SB14 - Inadvertent ADS opening i In this transient (Figures 6.1.3-57 to 6.1.3-63), the observed oscillations time interval was very short l and includes only [ ]'6 cycles. There was not time enough to establish a clean periodic behasior as shown in the other tests. This transient was on the borderline of the conditions required to establish the oscillations. Steam generation started at about [ ]*# seconds. At about [ ]'6# seconds, steam generation reached [ ]'6# lbm/sec. and just before [ l'6' seconds, the oscillations started. The injection temperature was higher than in the other transients because the IRWST was warmed-up for a longer time by the ADS 1-2-3 discharge. The liquid temperature at the top of the downcomer was higher, thus the condensation in the downcomer was reduced and as a result, the oscillation time period which is driven by the downcomer condensation lasted for a shorter period of time. At about [ l'6' seconds, the downcomer top reached saturation and the collapsed liquid level in the downcomer dropped below the cold-leg nozzles elevation and the oscillations ceased. 6.1.3.4 Expected Effect in AP600 Facility The oscillations observed in the OSU AP600 scaled test facility during IRWST injection may also occur in the AP600 facility. Timing of the initiation and termination of these oscillations may be different in the AP600 plant than was observed in the tests, and will depend on the following factors:
- Temperature and flow rate from the IRWST, as it affects the condensation rate of steam flowing 'o the downcomer compared to the production rate in the core
- Hydraulic flow characteristics of the steam path from the upper head to the downcomer, which determine the steam flow rate
- Liquid level in the downcomer relative to the cold leg, which will determine when the oscillations begin 9
mAap600c344w-64.noo:lt>100395 6.1.3-26 REVISION: 1
[ Since the pressures in AP600 plaat and the OSU AP600 scaled test facility will be close to atmospheric pressure during IRWST injection, there is no pressure scaling necessary for the condensation process. l The oscillation period in the AP600 plant can be estimated as follows, The oscillation period, t, is proportional to the time that is required to remove the mass, M, from the system in order to uncover the hot leg. That is:}}